View source for BVY 05-072, Vermont Yankee, Technical Specification Proposed Change No. 263 - Supplement No. 30 Extended Power Uprate - Response to Request for Additional Information

{{Adams
| number = ML052230384
| issue date = 08/01/2005
| title = Vermont Yankee, Technical Specification Proposed Change No. 263 - Supplement No. 30 Extended Power Uprate - Response to Request for Additional Information
| author name = Wanczyk R
| author affiliation = Entergy Nuclear Operations, Inc, Entergy Nuclear Vermont Yankee, LLC
| addressee name = 
| addressee affiliation = NRC/Document Control Desk, NRC/NRR
| docket = 05000271
| license number = 
| contact person = 
| case reference number = BVY 05-072, TAC MC0761
| document type = Letter type:BVY, Technical Specification, Bases Change
| page count = 473
| project = TAC:MC0761
| stage = Supplement
}}

=Text=
{{#Wiki_filter:Entergy Nuclear Vermont Yankee, LLC Entergy Nuclear Operations, Inc.
Brattleboro, Vr 05302-0500 August 1, 2005 Docket No. 50-271 BVY 05-072 TAC No. MC0761 ATTN: Document Control Desk U.S. Nuclear Regulatory Commission Washington, DC 20555-0001

==Subject:==
Vermont Yankee Nuclear Power Station Technical Specification Proposed Change No. 263 - Supplement No. 30 Extended Power Uprate - Response to Request for Additional Information

==References:==
: 1) Entergy letter to U.S. Nuclear Regulatory Commission, "Vermont Yankee Nuclear Power Station, License No. DPR-28 (Docket No. 50-271), Technical Specification Proposed Change No. 263, Extended Power Uprate," BVY 03-80, September 10, 2003
: 2) Entergy letter to U.S. Nuclear Regulatory Commission, 'Vermont Yankee Nuclear Power Station, License No. DPR-28 (Docket No. 50-271), Technical Specification Proposed Change No. 263, Supplement No. 24 - Response to Request for Additional Information," BVY 05-024, March 10, 2005
: 3) U.S. Nuclear Regulatory Commission (Richard B. Ennis) letter to Entergy Nuclear Operations, Inc. (Michael Kansler), uRequest for Additional Information - Extended Power Uprate, Vermont Yankee Nuclear Power Station (TAC No. MC0761)," July 27, 2005 This letter provides additional information regarding the application by Entergy Nuclear Vermont Yankee, LLC and Entergy Nuclear Operations, Inc. (Entergy) for a license amendment (Reference 1) to increase the maximum authorized power level of the Vermont Yankee Nuclear Power Station (VYNPS) from 1593 megawatts thermal (MWt) to 1912 MWt.
The major aspects of this submittal are:
: 1) An update to Entergy's response to request for additional information (RAI) item SRXB-A-6 regarding certain analytical methodologies of General Electric (GE) that are used for the design and evaluation of VYNPS' fuel. The prior response to SRXB-A-6 was provided with Entergy's letter of March 10, 2005 (Reference 2) and is being superseded by this submittal.
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BVY 05-072 Docket No. 50-271 Page 2 of 5
: 2) An executive overview summarizing Entergy's understanding of the key issues remaining to provide reasonable assurance of steam dryer integrity at EPU conditions and also summarizing the framework for Entergy's response to those issues.
: 3) Responses to a significant number of those RAls requested by NRC letter of July 27, 2005 (Reference 3). The remaining RAls that pertain to the steam dryer and piping/nozzle stress evaluations are not included, but will be transmitted as a separate submittal by August 4, 2005.
GE Analytical Methods In its letter of March 10, 2005, Entergy had proposed in its response to RAI SRXB-A-6 a means of addressing the NRC staffs questions regarding GE methods. The response was consistent with the Methods Interim Process proposed by GE in its letter of March 25, 2005 (MFN 05-005). Although Entergy remains confident that the concepts originally advanced in the response to RAI SRXB-A-6 are valid, an alternate, VYNPS-specific approach is provided by this letter. Entergy is revising and superseding the prior response to SRXB-A-6 with this submittal.
The alternate approach, discussed in the revised response to RAI SRXB-A-6 (Attachment 1), considers those core operating parameters and associated limits that could be impacted if all the uncertainties in methodology postulated by the staff were present during EPU operation, and then evaluating what, if any, operating restrictions should be imposed to compensate for this theoretical condition by providing additional safety margins to the affected limits. Using this approach Entergy has determined that a change of 0.02 to the safety limit minimum critical power ratio (SLMCPR) provides sufficient additional conservatism and adequate margin to address the postulated uncertainties in GE's methodology. Entergy is therefore proposing a license condition for EPU operation that imposes this additional 0.02 SLMCPR restriction until such time that the generic issues associated with GE analytical methods are adequately resolved with respect to VYNPS.
The alternate approach also describes Entergy's basis for confirming the adequacy of existing margin to accommodate the postulated uncertainties and assessing their impact on each of the remaining affected core operating parameters and associated limits. In addition, actual VYNPS operational experience with regard to core thermal limits is provided in the revised response to RAI SRXB-A-6.
Steam Dryer Analyses  provides an overview of Entergy's understanding of          the fundamental issues left to be resolved in order to provide reasonable assurance      that steam dryer integrity will be maintained at EPU conditions. These issues are         drawn from 129 individual questions posed by the NRC staff. Attachment 3 provides       a restatement of

BVY 05-072 Docket No. 50-271 Page 3 of 5 Entergy's overall approach to the steam dryer integrity issue and the framework of Entergy's strategy in addressing the remaining fundamental issues so that the answers to individual questions can be reviewed in that context. Attachment 5 provides responses to questions associated with computational fluid dynamics and steam dryer loads at EPU conditions. The remainder of the steam dryer-related RAls are in review and are expected to be submitted by August 4, 2005.
Response to Requests for Additional Information Attachments 4, 5, 7, 8, and 9 respond to individual RAls, according to NRC review branch. Of the 200 individual RAls requested by the NRC in Reference 3, 107 which pertain primarily to uncertainties in the acoustic circuit model, Scale Model Test benchmark adequacy, and applicability of the insights gained from the Quad Cities 2 instrumented dryer tests will be addressed in a future submittal, expected to be provided by August 4, 2005.
The revised response to RAI SRXB-A-6, as well as other responses to Reactor Systems Branch RAls, (Attachments 1 and 9) contain Proprietary Information as defined by 10CFR2.390 and should be handled in accordance with provisions of that regulation.
Attachments 1 and 9 are considered to be Proprietary Information in their entirety.
Attachments 2 and 10 are non-proprietary versions of Attachments 1 and 9, respectively.
Affidavits supporting the proprietary nature of the documents are provided as  (for Attachment 1), and as Attachment 12 (two affidavits for Attachment 9).
'Exhibits," which provide supporting information to certain RAI responses are included in 1.
This submittal provides a substantial portion of the information needed to support the preparation of the NRC's safety evaluation report for EPU and is therefore being submitted in advance of the responses to the remaining questions. In compiling and analyzing the information for this submittal, Entergy remains convinced that the VYNPS can be safely operated at up to 120% CLTP. It is our understanding that an audit of the underlying details supporting elements of this submittal will be conducted on or about August 22, 2005. Entergy anticipates that the nature of the audit will be confirmatory and respectfully requests that additional requests for information, if any, be communicated as soon as practical.
The following attachments are included in this submittal:
Attachment    l                                Title                             I I          Revised Response to RAI SRXB-A-6 (proprietary version) 2          Revised Response to RAI SRXB-A-6 (non-proprietary version) 3          Overview of Steam Dryer Issues 4          Responses to RAls EEIB-A-1 through EEIB-A-5 (no proprietary information) 5          Responses to RAls EMEB-B-18 through EMEB-B-149, non-inclusive (non-proprietary version) 6          Affidavit for Attachment 1

BVY 05-072 Docket No. 50-271 Page 4 of 5 7           Responses to RAls SPSB-C-47 through SPSB-C-52 (no proprietary information) 8           Responses to RAls SPLB-A-25 through SPLB-A-29 (no proprietary information) 9          Responses to RAls SRXB-A-7 through SRXB-A-58 (proprietary version) 10          Responses to RAls SRXB-A-7 through SRXB-A-58 (non-proprietary version) 11          RAI Response Exhibits (10) 12          Two affidavits for Attachment 9 13          New Regulatory Commitments (2)
There are two new regulatory commitments contained in this submittal that are incorporated into the responses to RAls EEIB-B-1 and EEIB-B-5 regarding actions associated with the postulated station blackout event. They are summarized in 3.
This supplement to the license amendment request provides additional information to clarify Entergy's application for a license amendment and does not change the scope or conclusions in the original application, nor does it change Entergy's determination of no significant hazards consideration.
Entergy stands ready to support the NRC staff's review of this submittal and suggests meetings (or audits of design files) at your earliest convenience.
If you have any questions or require additional information, please contact Mr. James DeVincentis at (802) 258-4236.
I declare under penalty of perjury that the foregoing is true and correct.
Executed on August a2        , 2005.
Sincerely, 6~./
Robert J. Vnyk                (/a Director, Ntcl'ar Safety Assurance Vermont Yankee Nuclear Power Station Attachments (13) cc:     (see next page)

BVY 05-072 Docket No. 50-271 Page 5 of 5 cc: Mr. Richard B. Ennis, Project Manager Project Directorate I Division of Licensing Project Management Office of Nuclear Reactor Regulation U.S. Nuclear Regulatory Commission Mail Stop 0 8 B1 Washington, DC 20555 Mr. Samuel J. Collins (w/o attachments)
Regional Administrator, Region 1 U.S. Nuclear Regulatory Commission 475 Allendale Road King of Prussia, PA 19406-1415 USNRC Resident Inspector (w/o attachments)
Entergy Nuclear Vermont Yankee, LLC P.O. Box 157 Vernon, Vermont 05354 Mr. David O'Brien, Commissioner (w/o proprietary information)
VT Department of Public Service 112 State Street - Drawer 20 Montpelier, Vermont 05620-2601

BVY 05-072 Docket No. 50-271 Attachment 2 Vermont Yankee Nuclear Power Station Proposed Technical Specification Change No. 263 - Supplement No. 30 Extended Power Uprate Response to Request for Additional Information Revised Response to RAI SRXB-A-6 NON-PROPRIETARY VERSION Total number of pages In Attachment 2 (excludina this cover sheet) Is23.

Attachment 2 BVY 05-072 Docket No. 50-271 Page 1 of 23 REVISED RESPONSE TO RAI SRXB-A-6 PREFACE This attachment provides a revised response to the NRC staffs request for additional information (RAI) dated December 21, 2004,1 regarding RAI SRXB-A-6. The response provided below supersedes the response provided by Entergy in its letter dated March 10, 2005.2 In its letter of March 10, 2005, Entergy had proposed in its response to RAI SRXB-A-6 a means of addressing the NRC staffs questions regarding GE methods. The response was consistent with the Methods Interim Process proposed by GE in its letter of March 25, 2005 to the NRC staff 3 . Although Entergy remains confident that the concepts originally advanced in the response to RAI SRXB-A-6 are valid, an alternate, VYNPS-specific approach appears to offer the most efficient path to resolving the NRC staff's concerns. Therefore, Entergy is revising and superseding the prior response to SRXB-A-6 with this submittal.
The alternate approach, discussed in the revised response to RAI SRXB-A-6 below, considers those core operating parameters and associated limits that could be impacted if the uncertainties in methodology postulated by the NRC staff were present during EPU operation, and then evaluating what, if any, operating restrictions should be imposed to compensate for this theoretical condition by providing additional safety margins to the affected limits. Using this approach Entergy has determined that a change of 0.02 to the safety limit minimum critical power ratio (SLMCPR) provides sufficient additional conservatism and adequate margin to address the postulated uncertainties in GE's methodology. Entergy is therefore proposing a license condition for EPU operation that imposes this additional 0.02 SLMCPR restriction until such time that the generic issues associated with GE's analytical methods are adequately resolved.
The alternate approach also describes Entergy's basis for confirming the adequacy of existing margin to accommodate the postulated uncertainties and assessing their impact on each of the remaining affected core operating parameters and associated limits. In addition, actual VYNPS operational experience with regard to core thermal limits is provided in the revised response to RAI SRXB-A-6.
  '  U.S. Nuclear Regulatory Commission (Richard B. Ennis) letter to Entergy Nuclear Operations, Inc.
(Michael Kansler), uRequest for Additional Information - Extended Power Uprate, Vermont Yankee Nuclear Power Station (TAC No. MC0761)," December 21, 2004 2 Entergy letter to U.S. Nuclear Regulatory Commission, "Vermont Yankee Nuclear Power Station, License No. DPR-28 (Docket No. 50-271), Technical Specification Proposed Change No. 263, Supplement No. 24 - Response to Request for Additional Information," BVY 05-024, March 10, 2005 3 GE Nuclear Energy (George B.Stramback) letter to U.S. Nuclear Regulatory Commission (Herbert Berkow), "Methods Interim Process (TAC No. MC5780)," MFN 05-005, March 25, 2005

Attachment 2 BVY 05-072 Docket No. 50-271 Page 2 of 23 Reactor Systems Branch (SRXB)
Boilinci Water Reactors and Nuclear Performance Section (SRXB-A)
(The RAI is stated below as provided in NRC's letter of December 21, 2004.)
RAI SRXB-A-6 Table 1-1 in Attachment 6 of the application dated September 10, 2004 [sic], lists all the nuclear steam system codes used for the EPU request. Section 1.2.2 of Attachment 6, 'Computer Codes," indicates that the VYNPS application of these codes complies with the limitations, restrictions, and conditions specified in the applicable NRC safety evaluation report (SER) that approved each code, with exceptions as noted in Table 1-1.
Similarly, review the fuel vendor's analytical methods and code systems used to perform the safety analyses supporting the VYNPS EPU application and provide the following information:
(a)     Confirm that the steady state and transient neutronic and thermal-hydraulic analytical methods and code systems used to perform the safety analyses supporting the EPU conditions are being applied within the NRC-approved applicability ranges.
(b)     Confirm that for the EPU conditions, the calculational and measurement uncertainties applied to the thermal limits analyses are valid for the predicted neutronic and thermal-hydraulic core and fuel conditions.
(c)     Confirm that the assessment database and the assessed uncertainty of models used in all licensing codes that interface with and/or are used to simulate the response of VYNPS during steady state, transient or accident conditions remain valid and applicable for the EPU conditions.
Revised Response to RAI SRXB-A-6 Margin in GE Analytical Methods Supporting VYNPS EPU Submittal Summary As part of Entergy's submittal of a license amendment request for the VYNPS extended power uprate (EPU), Entergy is proposing an alternate approach to address NRC questions related to GE's standard methodologies to facilitate NRC approval of the request. The alternate approach includes the addition of an operational restriction on bundle critical power ratio to be implemented via an adjustment of 0.02 ACPR to the safety limit minimum critical power ratio (SLMCPR). The following discussion provides the bases for the adequacy and additional conservatism, with respect to the aforementioned NRC questions, of the margins in pertinent safety parameters for the VYNPS EPU provided by GE's standard methodologies as further augmented by the proposed Entergy alternative operational restriction. It is intended that the operational restriction in the alternate approach would only be implemented as a condition of the

Attachment 2 BVY 05-072 Docket No. 50-271 Page 3 of 23 EPU License Amendment until the aforementioned NRC questions are otherwise satisfactorily resolved.
Introduction In its review of GE's generic MELLLA+ submittal4 , the NRC has asked questions related to the adequacy, given the absence of recent gamma-scan test data, of the standard uncertainties and biases utilized in GE's bundle lattice and core simulation methodologies (see 'GE Bundle Lattice and Core Simulation Methodology & Utilization of Gamma Scan and Fuel Isotopic Data" section below) for current fuel designs and operating strategies and the potential effect on safety parameters influenced by such uncertainties and biases. As noted in the "GE Bundle Lattice and Core Simulation Methodology & Utilization of Gamma Scan and Fuel Isotopic Data" section, GE has benchmarked its methods using industry standard techniques and utilized gamma scan data to retrospectively confirm the adequacy of certain elements of its methods and benchmarking. GE has provided, and continues to provide, information to the NRC supporting the adequacy of GE's methodologies for application to BWRs and BWR expanded operating domains.
The following discussion addresses the NRC questions regarding both gamma-scan and isotopic data and supports the alternate approach of addressing the NRC's questions regarding uncertainties and biases which is being proposed by Entergy as an element of the VYNPS extended power uprate (EPU) license amendment request. The alternate approach includes a proposed operational restriction in bundle critical power ratio implemented via an increase in the SLMCPR of 0.02 ACPR. This operational restriction provides additional margin and addresses the aforementioned NRC questions. The following discussion identifies the fuel parameters related to the gamma scan and isotopic data and addresses the effect of uncertainties in those fuel parameters to the extent each is applicable to the six pertinent safety parameters which are influenced by those fuel parameters. For each of the six pertinent safety parameters:
: 1. the fuel parameters which affect it are identified,
: 2. the treatment of fuel parameter uncertainties in the safety parameter limit development is considered, and
: 3. the adequacy of the existing treatment in conjunction with the alternate approach is supported.
Safety Parameters Influenced by Noted Uncertainties and Biases GE has reviewed its methodologies to determine the uncertainties and biases which were confirmed by earlier gamma scan test data or measurements of irradiated fuel isotopics and to confirm that the existing types of uncertainties already included in GE's NRC-approved treatment of uncertainties and biases address the NRC staff questions regarding the absence of 4 MELLLA+ is not part of Entergy's VYNPS EPU application; however, NRC requests for information on the related subject are contained in that docket and are provided here as reference.

Attachment 2 BVY 05-072 Docket No. 50-271 Page 4 of 23 recent confirmatory test data.
The associated fuel parameters related to such test data and measurements that are not otherwise measurable directly or indirectly by existing operating plant instrumentation, e.g., local power range monitors (LPRMs) and TlPs, are:
: 1. Local fuel pin power and exposure (depletion) vs. axial position,
: 2. Relative local fuel pin power and exposure (local in-bundle peaking),
: 3. Void reactivity coefficient, and
: 4. ((                                                      1]
The fuel parameter uncertainties of interest are thus related to relative local and pin power peaking, void reactivity coefficient, and ((                              )). Other nodal fuel and bundle parameters, e.g., lattice reactivity, bundle power, and bundle axial power shape, are satisfactorily and adequately confirmed by comparisons to operating plant data or tests, e.g.,
traversing in-core probes (TIP) data and shutdown margin demonstrations.
The safety parameters potentially influenced by the local and relative local pin power uncertainties and the ((                           )) uncertainty are:
: 1. Critical power (controlled by the SLMCPR and OLMCPR),
: 2. Shutdown margin (controlled with a technical specification limit of 0.38% AkMk),
: 3. Fuel rod thermal-mechanical performance (controlled by limits on linear heat generation rate, LHGR),
: 4. LOCA-related nodal power limits (controlled via the maximum average planar linear heat generation rate, MAPLHGR),
: 5. Stability (protected by the SLMCPR, OLMCPR, and stability solutions), and
: 6. Licensed pellet exposure (e.g., 70 GWd/MT for GE14 fuel)
Each of the uncertainties in question is currently included and addressed in the treatment of uncertainties and biases in GE's NRC-approved methodologies to determine these safety parameters. GE believes it is appropriate to continue to utilize the NRC-approved GE treatment of uncertainties and biases. If consideration of larger uncertainties is deemed appropriate, such uncertainties can be utilized in the existing treatments of propagation and combination of uncertainties. Direct application of biases into best estimate codes in an attempt to address potential uncertainty concerns is not appropriate because such introduction of unqualified biases would lead to potential non-conservatisms in resulting predictions. Therefore, the fidelity of GE's codes and methods is best maintained by not artificially adding biases. Conservative limits on safety parameters, developed with consideration for such uncertainties, provide reasonable assurance of safety.
A discussion of the adequacy of the margin existing in, and, as applicable, augmented margin for, each of these safety parameters is provided below, again based on the alternate approach being proposed by Entergy for the VYNPS EPU submittal.

Attachment 2 BVY 05-072 Docket No. 50-271 Page 5 of 23 SLMCPR Margin The safety limit minimum critical power ratio (SLMCPR) is directly affected by the fuel parameters confirmed by gamma scan data. The local pin power peaking (axial and in-bundle) and ((                               )) uncertainties are factors which affect SLMCPR. SLMCPR is not affected by void reactivity coefficient uncertainties. The other safety parameters, discussed in succeeding sections, already incorporate other conservatisms which encompass the pin power/exposure, ((                              1] and void reactivity coefficient uncertainties or are not influenced by those uncertainties.
GE's NRC-approved process for determining the SLMCPR incorporates the applicable uncertainties in the lattice and core physics parameters, and the method of determining SLMCPR assures that fuel is protected from boiling transition when such uncertainties are incorporated. Uncertainties in local pin power peaking and ((                                   )) (and bundle power) are explicitly included in the SLMCPR determination and considered separately, then cumulatively below.
The potential effect of larger pin power uncertainty on the SLMCPR has been considered. First, in lieu of an arbitrary increase in the uncertainty, a review of ((
                                                     )) Inthe determination of SLMCPR, the use of additional pin power uncertainty so derived, i.e., ((
                                                  )), providing real additional critical power margin relative to GE's standard methodology and addressing local peaking uncertainty concerns.
[I                                       )) is a component of the total bundle power uncertainty.
The total bundle power uncertainty for application within GE's NRC-approved SLMCPR determination process consists of the component uncertainties in Table SRXB-A-6-1 (From Table 4.2, page 4-2 in NEDC-32694P-A).

Attachment 2 BVY 05-072 Docket No. 50-271 Page 6 of 23 Table SRXB-A-6-1 Quantity                       Uncertainty                        Source
((
I I                                        l                              a GE has continued to provide the NRC with BWR fleet information on the consistency of integral TIP comparisons on periodic basis, e.g., in fuel technology updates. In 2005, GE formally provided a large amount of data for uprated plants loaded primarily with 10 x10 fuel in methods related RAI responses under the MELLLA+ docket (MFN 05-029, TAC No. MC5780).
Examination of these data confirms the applicability and conservatism of the original ((        ))
uncertainty documented in GE's NRC-approved topical report describing the SLMCPR methodology power distribution uncertainties (NEDC-32601 P-A/NEDC-32694P-A).
((

Attachment 2 BVY 05-072 Docket No. 50-271 Page 7of23 1))
BWRs have always operated at void fractions higher than 70% with some of the earlier gamma scan data from fuel exceeding 80% void fractions so that the effect of void fraction is included in confirmation of local and bundle power peaking uncertainty and, thus, not a significant concern.
Instead, the largest differences in bundle power are the result of depletion and are not the result of differing product lines, composition, or core power. This key aspect is already addressed in the current NRC approved value ((
                                         )) Therefore, the procedure of using the current gamma scan data to determine a conservative bound on the uncertainty is reasonable and valid.
((
                                                                      )) This additional critical power margin provides a real additional assurance of safety and is developed consistent with current NRC-approved bundle power uncertainty methodology.
The effects of ((                                                             Jlin Table SRXB-A-6-1 on the bundle power uncertainty for SLMCPR determination ((
[1                                                                                                  ]10.02 ACPR effect on SLMCPR based on the conservatively increased local peaking ((
                    )) uncertainties. ((                            ))is further conservative.
In summary, use of alternative, even more conservative values for uncertainties in the local peaking factor ((                                flresults in an increase in the SLMCPR for VYNPS of 0.02 relative to that calculated with current GE standard methodology and provides additional reasonable assurance of safety for VYNPS EPU with respect to SLMCPR.
OLMCPR Margin The analysis of anticipated operational occurrences (AOOs) examines the change in critical power ratio relative to the original starting point conditions and determines the most limiting

Attachment 2 BVY 05-072 Docket No. 50-271 Page 8 of 23 transient event. The fuel parameters identified previously, i.e., the local pin power peaking, void reactivity coefficient, and ((                                      )) are factors in the evaluation of limiting A0Os. ((
                                                    )) This assures that the analysis is both realistic but conservative.
Accommodation for uncertainties in local pin power peaking and ((                                J] (and bundle power), i.e., consideration of bundle and nodal powers higher (or lower) than expectations, is directly incorporated in the licensing methodology. Thus, there is no effect on ACPR due to the NRC staff questions regarding the local pin power peaking and ((
             )) uncertainties.
Both the ODYN and TRACG transient methodologies have established application ranges for void coefficient uncertainty. The approval of and GE confidence in the basis for these methodologies are based upon comparison of calculations for a wide variety of plant transients in which the nominal void coefficient is used. The acceptable performance of these codes relative to the data justifies that no large errors in void coefficient exist.
The ODYN model uncertainty is based on comparisons to the benchmark Peach Bottom turbine trip tests. ((
                                       )) Because inputs to the OLMCPR analysis are conservative, and the pressurization transients are conservatively analyzed by ODYN and typically establish the limiting ACPRs, conservatisms existing in the process of determining OLMCPRs address NRC questions related to gamma scans and fuel isotopics as they relate to OLMCPR.
In summary, the standard GE methodologies utilized to establish the OLMCPR conservatively address uncertainties issues and provide reasonable assurance of safety for VYNPS EPU with respect to OLMCPR.

Attachment 2 BVY 05-072 Docket No. 50-271 Page 9 of 23 Shutdown Margin (SDM)
The analysis of SDM considers whether core reactivity can be safely controlled. The fuel parameters identified previously, i.e., the local pin power peaking and ((
         )), are indirect factors in the evaluation of SDM since uncertainties in those parameters may ultimately influence prediction of fuel depletion and, thus, fuel reactivity. Void reactivity coefficient is not a contributor since essentially zero voiding is present at hot or cold shutdown conditions. As described in the "GE Bundle Lattice and Core Simulation Methodology &
Utilization of Gamma Scan and Fuel Isotopic Data" section, the GE bundle lattice and core simulation methodologies are best estimate predictions so that validation of operating benchmark data, core follow, and core licensing can proceed using consistent methodology.
Comparisons to actual plant cold critical states are an important part of this verification because any error in bundle or nodal power (or exposure) would tend to degrade the ability of the core simulator to establish a stable bias (in eigenvalue), a measure of the ability of the model to reliably predict core hot and cold critical conditions. While the Technical Specification for SDM is 0.38% Ak/k reactivity (for an in-sequence check only), normal GE design procedure is to provide design cold shutdown margins of 1%or more depending on customer request and GE procedure. For VYNPS, the standard design SDM is 1.1% Ak/k to provide additional flexibility in cycle length and operations. The uncertainty in cold critical predictive capability is considered and included in this choice of SDM requirement.
However, it is very important to note that actual SDM is a demonstrated quantity (plant verification) during plant startups or by use of local criticality confirmations. Inaddition, trending of hot eigenvalue (i.e., reactivity anomalies), also required by Technical Specifications and another direct confirmation of the adequacy of GE's methods with respect to fuel depletion and reactivity predictions, is performed. Because such plant verification data from power uprated plants and plants with modern fuel designs, including GE14, have continued to confirm that adequate SDM exists and that eigenvalue biases in GE's methods are stable and well understood, there is sufficient justification for the adequacy of GE's bundle lattice and core simulation methodologies and the uncertainties in the nodal and bundle power and exposure even without recent confirmatory gamma scan or fuel isotopic data.
In summary, the current design process and design goal, in combination with the existing processes of plant verification of SDM and trending of hot eigenvalues, provide reasonable assurance of adequate SDM.
LHGR Margin For each GNF fuel design, including GE14, thermal-mechanical based linear heat generation rate (LHGR) limits are specified for each fuel rod type (for both U02 and gadolinia-bearing rods) such that, if each rod type is operated within its LHGR limits, all thermal-mechanical design and licensing criteria, including those which address response to anticipated operational occurrences (AOOs), are explicitly satisfied and fuel rod integrity is maintained. The fuel parameters identified previously, i.e., the local pin power peaking, void reactivity coefficient,

Attachment 2 BVY 05-072 Docket No. 50-271 Page 10 of 23
[l                                   11' are factors, to differing extents, in the development of LHGR limits. The fuel parameters ultimately determine the local power, which is explicitly addressed by the LHGR limit.
Fuel rod thermal-mechanical licensing criteria explicitly considered in the specification of LHGR limits include fuel centerline temperature, cladding plastic strain, and fuel rod internal pressure.
Each of these criteria is limiting over a portion of the fuel rod lifetime. For development of the final limit curve, the peak power node is conservatively assumed ((
                                 )). In addition, model and operating uncertainties are explicitly addressed in the development of limits, including an additional power uncertainty of [i                 ))
power that is not specifically assigned to any cause, as well as a ((              )) conservative power bias in the fuel rod internal pressure calculation. The uncertainty and bias also apply to exposure because, in the determination of LHGR limits, the exposure is the integrated power.
Moreover, the model uncertainties in GE's NRC-approved thermal-mechanical analysis methodology (GSTRM) are based upon temperature benchmark data and are also validated via fission gas benchmark data for which the nominal power history is produced in the steady-state core simulator. Because the large uncertainties included by this process encompass the uncertainties in local and rod power reflected in the NRC staff questions and because separate experimental benchmarking information confirms that the model uncertainties remain valid, an adjustment to provide additional LHGR margin is unnecessary.
In summary, the standard GE methodology for determining LHGR limits includes conservative consideration for, and provides reasonable assurance of adequate margin to address, the power and void reactivity uncertainties in question.
MAPLHGR Margin The purpose of the maximum average planar linear heat generation rate (MAPLHGR) limits is to assure adequate protection of the fuel during a postulated loss-of-coolant accident (LOCA) with the defined operation of emergency core cooling system (ECCS). The fuel parameters identified previously, i.e., the local pin power peaking and ((                                        )),
are factors, to differing extents, in the development of LHGR limits. The fuel parameters ultimately determine the local power, which is the subject of the MAPLHGR, a local limit. The void reactivity coefficient is not a factor in the ECCS-LOCA analysis.
The ECCS-LOCA analysis applicable to the VYNPS EPU follows the NRC-approved SAFER/GESTR application methodology documented in NEDE-23785-1-PA Rev. 1, "The GESTR-LOCA and SAFER Models for the Evaluation of the Loss-Of-Coolant Accident, Volume l1l, SAFER/GESTR Application Methodology," October 1984. The analytical models used to perform ECCS-LOCA analyses are also documented in NEDE-23785-1-PA together with NEDE-30996P-A, -SAFER Model for Evaluation of Loss-of-Coolant Accidents for Jet Pump and Non-jet Pump Plants, Volume I, SAFER - Long Term Inventory Model for BWR Loss-of-Coolant Analysis," October 1987, and NEDC-32950P, "Compilation of Improvements to GENE's SAFER

Attachment 2 BVY 05-072 Docket No. 50-271 Page 11 of 23 ECCS-LOCA Evaluation Model," January 2000.
When SAFER/GESTR methodology is applied, the hot bundle is initialized with a ((
                                                                 *3. In addition, a ((
               ]I. In order to ensure that the SAFER analysis is bounding for all exposures, the hot rod of the hot bundle is placed at the exposure corresponding to the ((
                                    )). In addition to these analytical conservatisms, margin to the MAPLGHR limits is maintained during plant operations.
Total bundle power is also important to the severity of the ECCS-LOCA analysis. ((
                                           )). Furthermore, the ECCS-LOCA basis target MCPR is set lower than the OLMCPR so that the OLMCPR is not set by the ECCS-LOCA analysis (thus set by the AOO analysis).
Pin power peaking for the hot rod is set to a [
                                   )) to further insure that the ECCS-LOCA results are bounding.
Lastly, the axial power profile ((
                                                                       ))
The above considerations indicate that significant conservatisms related to initial local pin and bundle powers exist in the GE SAFER/GESTR ECCS-LOCA methodology.
In addition to the above conservatisms, the VYNPS Licensing Basis peak clad temperature (PCT) determined by the methodology described above is 19600F. This result includes application of Appendix K modeling assumptions. The maximum nominal PCT is about
[I        )) lower than the Appendix K value. When the nominal PCT is adjusted to account for model uncertainties (at 95% probability), the PCT (also known as the Upper Bound PCT in the SAFER/GESTR methodology) is about ((                  )) lower than the Licensing Basis PCT. The 95% probability PCT includes an uncertainty of ((           )) on the LHGR (((             ))).
The SAFER/GESTR methodology assumes a bounding post-LOCA core power trajectory and, thus, core kinetics are not modeled. The average and hot bundle void profile is determined by SAFER at the limiting initial conditions described above as well as at the post-LOCA conditions.
Uncertainties in predictions of void reactivity have no impact in the SAFER/GESTR methodology. The overall SAFER/GESTR methodology is designed to maximize the PCT.
In summary, the conservatism of the present ECCS-LOCA methodology used to determine

Attachment 2 BVY 05-072 Docket No. 50-271 Page 12 of 23 MAPLGHR limits adequately considers the effects of the uncertainties in local and bundle power and provides reasonable assurance that those limits provide adequate margin to protect the fuel.
Stability BWR thermal-hydraulic stability analyses are performed to assure that SLMCPR is protected in the event of a thermal-hydraulic instability event. The fuel parameters in question affect stability performance.

==Background:==
VYNPS has implemented the Option l-D solution documented in "Application of the
 'Regional Exclusion with Flow-Biased APRM Neutron Flux Scram' Stability Solution (Option l-D) to Vermont Yankee" Licensing Topical Report, GENE-637-018-0793, July 1993. Option l-D has (1) "prevention" elements, these being the Exclusion and Buffer Regions, and (2) a "detect &
suppress' element, this being Safety Limit MCPR (SLMCPR) protection provided by the flow-biased APRM flux scram trip (for the dominant core wide mode of coupled thermal-hydraulic/neutronic reactor instability).
The prevention portion of the solution includes separate administratively controlled exclusion and buffer regions which are evaluated for every reload. The detect-and-suppress portion of the solution is a flow-biased APRM flux scram trip that prevents oscillations of significant magnitude. This scram ensures the Fuel Cladding Integrity SLMCPR is met for thermal hydraulic oscillations.
Stability analyses for both the EPU and fuel cycle specific conditions are performed to define the exclusion and buffer regions as well as to confirm that the scram setpoints meet the design basis. With respect to power distribution uncertainties of the nuclear simulator data, the results pertaining to the exclusion region may be slightly affected, but this is not considered to have any safety significance for reasons described below. The power distribution uncertainties of the nuclear simulator data are considered in the determination of the limiting bundle conditions and therefore have insignificant impact on the flow-biased APRM flux scram trip setpoint and the SLMCPR protection. An increase to the void reactivity used in the GE stability analysis models (the frequency domain code ODYSY and the time-domain code TRACG) may also affect the predicted results. However, the current stability models have been used to model actual instability events, and the decay ratio acceptance criteria have been established consistent with the uncertainty as documented in the approved licensing reports. Furthermore, recent instability events at two domestic BWRs have also been evaluated with the stability models and shown to meet the previously established criteria. This provides high confidence that the GE methodology is adequately simulating recent fuel designs and fuel power densities. Therefore, no adjustment to stability models or analysis is necessary due to potential void reactivity uncertainties.
Exclusion Region calculation: The NRC-approved ODYSY methodology (NEDC-32992P-A, July 2001) is used in the exclusion region calculation for every reload. The calculation of the

Attachment 2 BVY 05-072 Docket No. 50-271 Page 13 of 23 exclusion region boundary is based on a very conservative core wide decay ratio (((
       ))) that may be influenced by the core wide axial power distribution calculation.
((
                                              )) An additional protection feature includes a cycle-specific buffer region, which is 5% in rated core power or 5% in rated core flow, beyond the exclusion region. Manual monitoring of the decay ratio is required while operating in the buffer region.
The decay ratio calculation includes a cycle-specific confirmation that core wide oscillation is the predominant reactor instability mode .for VYNPS and that regional mode instability is not probable. The dominance of the core-wide mode oscillation is confirmed for every reload at the most limiting state point on the EPU power/flow map. The calculation to confirm that regional mode of instability is not likely may be affected by uncertainties in power distribution because it considers the limiting bundle power; ((
Therefore, reasonable potential local or bundle power distribution uncertainties do not affect the confirmation that regional oscillations are not likely for VYNPS.
Detect and Suppress calculation: The detect and suppress evaluation for the VYNPS EPU is performed under the approved LTR basis (NEDO-32465-A, General Electric Company, "Reactor Stability Detect and Suppress Solutions Licensing Basis Methodology for Reload Applications,"
August 1996,). The flow biased APRM scram setpoints are initially established with conservative margin such that they are found applicable to future fuel cycles during reload confirmation calculations. The calculation of the scram setpoints is based on the limiting fuel bundle being at the Operating Limit MCPR (OLMCPR) and the SLMCPR not being exceeded during the instability oscillation.
The detect and suppress calculation requires the use of the DIVOM (which is defined as the Delta CPR over Initial MCPR Versus the Oscillation Magnitude) curve. Per the new BWROG Guideline (GE-NE-0000-0031-6498-RO, "Plant-Specific Core-Wide Mode DIVOM Procedure Guideline," June 2005), a plant- and cycle-specific DIVOM evaluation is used to establish the plant specific relationship between the Hot Channel Oscillation Magnitude (HCOM) and the relative change in MCPR such that the initial MCPR value corresponds to the OLMCPR and the limiting MCPR value remains above the SLMCPR. ((
I'

Attachment 2 BVY 05-072 Docket No. 50-271 Page 14 of 23
((
3] The scram setpoint analytical limit is established such that the hot channel power is maintained below acceptable values.
In summary, the uncertainties in power distribution calculation and void reactivity do not significantly impact the safety margin in the stability analysis for VYNPS.
Margin to Licensed Fuel Exposure GE fuel designs are licensed to a peak pellet exposure (i.e., 70 GWd/MTU for GE14). This is equivalent to a GE14 rod average exposure of -61.4 GWd/MTU, but there is not an explicit rod average exposure limit for GE14 or other GE fuel designs. This limit is used to assure that fuel is not operated beyond its analysis basis. The fuel parameters identified previously (i.e., the local pin power peaking, void reactivity coefficient, f:                                         ]3) are factors, to differing extents, in the development of LHGR limits, and, thus, the fuel exposure limit. The fuel parameters ultimately determine the local power, which is explicitly addressed by the LGHR limit.
Fuel rod internal pressure is the limiting licensing criterion at end-of-life for GE fuel designs. The fuel cladding creep rate is a function of cladding temperature and in turn of LHGR. As discussed previously, the LHGR limits for GE14 are deliberately conservative with respect to local rod power, assume a conservative pellet swelling rate uncertainty, and are also specified such that the margin to the criterion for limiting pellet-cladding gap increase due to rod internal pressure is actually smaller several GWd/MTU before end-of-life than at the peak pellet exposure (end-of-life) limit. Thus, existing uncertainties and margins in GE's NRC-approved fuel thermal-mechanical methodology adequately address the NRC question regarding local peaking uncertainty with respect to the licensed fuel exposure limit.
In summary, the GE standard fuel thermal-mechanical analysis basis considers and provides adequate margin for uncertainties in local and bundle power.
Additional Margin Summary If it is desirable to address NRC questions regarding the adequacy of GE's standard uncertainties in local power/exposure, ((                             )), or void reactivity coefficient for EPU conditions in the absence of recent confirmatory gamma scan and fuel isotopic data via an alternate approach of incorporating additional margin in appropriate safety parameters, the evaluation above provides the basis for a determination that an operational restriction implemented via an adjustment to the SLMCPR of 0.02 ACPR provides additional and reasonable assurance of safety for VYNPS at EPU conditions. Significant conservatisms already exist in the processes for determination of the other safety parameters, i.e., OLMCPR margin, SDM, LHGR, MAPLHGR, thermal-hydraulic stability protections, and fuel (peak pellet)

Attachment 2 BVY 05-072 Docket No. 50-271 Page 15 of 23 exposure, to address the NRC staff questions, and adjustments or operational restrictions for these are, thus, not required.
GE Bundle Lattice and Core Simulation Methodology & Utilization of Gamma Scan and Fuel Isotopic Data Summary GE's bundle lattice and core simulation codes, TGBLA and PANACEA, are best-estimate methods with uncertainties and biases in inputs and outputs of those codes addressed by the conservative treatment, previously approved by the NRC, of uncertainties and biases propagation in GE's calculations of conservative limits for various fuel safety parameters. The bundle lattice methods have been benchmarked, using industry standard practice, against Monte Carlo calculations for all GE fuel types. These benchmarks have been further confirmed for certain GE fuel types, retrospectively, with gamma-scan data available to GE. The core simulator methodology has been benchmarked, again using industry standard practice, against the operating plant instrumentation, e.g., traversing in-core probes (TIPs). ((
                                             )) Operating plant data are continuously utilized to evaluate the accuracy of predictions of the bundle lattice and core simulator methodologies on both a plant-specific and BWR fleet-wide basis, and such trending is periodically (approximately annually) reviewed with the NRC staff in fuel technology update meetings.
In accordance with its understanding of previous NRC-approved licensing topical reports and NRC-issued safety evaluations for GE's methods, GE has evaluated and reflected the accuracy of its methodologies as it has introduced new fuel designs and operating strategies. GE believes that its bundle lattice and core simulator methodologies, including the associated uncertainties and biases utilized by GE, in combination with its NRC-approved treatment of uncertainties and biases, are adequately predicting the performance and assuring the safety of BWRs at up to and including 120% EPU conditions.
Qualification Process GE utilized rod gamma scans, i.e., measurements of gamma emissions from certain fission product isotopes in individual irradiated BWR fuel rods, to further confirm the ability of its benchmarked methods to adequately predict local (fuel pin) power and exposure (i.e., burnup or depletion). GE utilized bundle gamma scans, i.e., scans of entire BWR fuel bundles, to confirm an appropriate value for uncertainty related to the ((                                           )).
GE utilized irradiated fuel rod isotopic measurements, i.e., radiochemistry determination of inventory of certain fission and activation products, which are necessarily limited in number due to the difficulty in obtaining such measurements, in lattice physics code development but not as part of code benchmarking.
GE evaluates methods on multiple geometrical bases. The process of monitoring operational

Attachment 2 BVY 05-072 Docket No. 50-271 Page 16 of 23 core parameters provides an up-to-date (hourly) evaluation of steady-state core reactivity control and provides a way to evaluate the core simulator eigenvalue bias. Comparison of calculated to measured TIP signals provides confirmation of the three-dimensional field of flux/power on a very timely basis (monthly) but with a resolution scale that only reflects the coarse mesh resolution of the three-dimensional simulator. Natural noise in the TIP instrumentation conservatively results in a fundamental contribution of 1% to the evaluated comparison (NEDC-32694P-A, page A-7). Bundle-wise or pin-wise gamma scans allow for a better resolution in space but result in a poor temporal comparison because the present concentration of the typically measured fission product (140Ba) requires an integration of the power history for the prior sixty days. Moreover, because of limitations, gamma scans may only be achievable once per cycle for operating power reactors. Bundle gamma scans usually entrain an experimental uncertainty of 1%(Ia) in the measured values while rod gamma scans entrain an uncertainty of 2%(Ia).
Because the injection of experimental error of non-routine benchmarking may confound physical phenomena of interest and for purposes of more timely and comprehensive evaluation, it is meaningful to compare production lattice physics methods (TGBLA) to Monte Carlo methods whose efficacy has been established through comparison to critical benchmarks. Assuming adequate trials have been considered, the local accuracy provides significant insight for examination of relative local pin peaking accuracy. If the local power is being produced correctly, the subsequent depletion of the fuel is occurring at the correct rate and location.
Furthermore, assuming the nominal production lattice physics code produces stable core eigenvalue behavior (evaluated in the operational core follow examination), use of the depleted isotopic compositions from the deterministic code for comparisons to Monte Carlo later in the life of the fuel is both meaningful and produces further insight into modeling accuracy. The conclusion is that it is meaningful and proper to consider comparisons between TGBLA and Monte Carlo methods in evaluation of methods accuracy.
In summary, the GE standard fuel lattice and core simulator methodology qualification process utilizes a large volume of contemporaneous operating plant data supported by available confirmatory, retrospective gamma scan to assure high-quality best-estimate predictions of local, nodal, bundle, and core power. As discussed above, GE's NRC-approved treatment of the uncertainties in the power predictions assure conservative limits for the safety parameters influenced by the local, nodal, bundle, and core power.
VYNPS Core Follow / Required Technical Specification Surveillance Information In addition to the above arguments related to GE analytical methods, VYNPS and Global Nuclear Fuels personnel perform core follow and required Technical Specification surveillance activities in the effort to ensure the VYNPS core is operating as expected.
CASMO-41SIMULATE-3 Overview The CASMO-4 and SIMULATE-3 codes are part of the Studsvik Scandpower Core Master

Attachment 2 BVY 05-072 Docket No. 50-271 Page 17 of 23 System. CASMO-4 is a multigroup transport theory code which feeds cross sections and discontinuity factors into SIMULATE-3. SIMULATE-3 is an advanced two-group nodal diffusion code with the ability to perform pin-power reconstruction. The code package is primarily used to independently verify vendor calculations and confirm the core is behaving as predicted during the cycle. Entergy receives the most up-to-date versions of the codes when available.
Entergy, through the core follow procedure, uses the CASMO-4/SIMULATE-3 package to trend the online performance of key core parameters. The key parameters, indicated in the core follow procedure, include MFLCPR, MFLPD, MAPRAT and gamma TIPS.
Thermal Limits 3D-MONICORE T &'(3DM) is the plant adaptive online software. At the heart of 3DM is the PANACEA-11 (P11) software engine that runs in shape adaptive mode to calculate core state points. Offline non-adaptive P11 is used by Global Nuclear Fuels (GNF) engineers to show the agreement between the offline and online comparison during core follow. Offline P11 and SIMULATE-3 are used to model online data provided by 3DM to ensure future predictions with the codes are correct.

Attachment 2 BVY 05-072 Docket No. 50-271 Page 18 of 23 MFLCPR (Maximum Fraction Limiting Critical Power Ratio)
UpdaIed Jul.005 W C24 MFLCPR J
1.N::::
0.95  -                   ..                                           ...............
O.R. -      -------              ..........   ..   ......                 .........
gr0.75 -    --------------------------------
__.*                                    -----  _---r   .. _---o
                                                                                - ........       -_-----ow.
                                                                                                            .. , ...... I...............
I I
                                                                      -Sir3             Core Followw C 0.70    - - - - -- - -
                                                                       -a,-- PitI Core Followw. Cl
                                                                      .      3DM 0.65                  .
0              1             2        3    4         5         6              7          8   9         10     11      12 13  14   is Exposure (GWdIST)
Figure SRXB-A-6-1 Cycle 24 MFLCPR SIMULATE-3 MFLCPR is calculated through Entergy's own in-house code JAFCPR 2.1 which is fed power distributions and core parameters from SIMULATE-3. The code uses a similar approach to P1 1's CPR routine giving high confidence in its accuracy to calculate MFLCPR.
SIMULATE-3 shows good agreement with GNF methods as shown in Figure SRXB-A-6-1 and there is substantial margin where the codes do not agree. The data points beyond the last 3DM point represent the as designed expected values for each thermal limit, in this case the Cycle 24 (current cycle) Cycle Management Report (CMR), Supplement 2.

Attachment 2 BVY 05-072 Docket No. 50-271 Page 19 of 23 MFLPD (Maximum Fraction Limiting Power Density)
Updated Jul2005 VY C24 Pin Power ReconstructIon Based MFLPD 1.M -_
0.95 - -.
0.90 -
I 085-0.
e:
080-..
0    1   2   3    4    5    6      7     8      9 10  11 12 13      14   15 Exposure (GWd/ST)
Figure SRXB-A-6-2 Cycle 24 Pin Power Reconstruction Based MFLPD SIMULATE-3 uses its own pin power reconstruction module to determine MFLPD. As in Figure SRXB-A-6-2, in most cases, SIMULATE-3 and P11 calculate a larger and more conservative MFLPD than is representative in actual 3DM online operations, but the two offline codes agree relatively well.

Attachment 2 BVY 05-072 Docket No. 50-271 Page 20 of 23 MAPRAT (Maximum Average Planer Linear Heat Generation Ratio)
Updated Jul2065 VY C24 MAPRAT
                                                                                 . Sirn3
                                                                                      .              Core Followw. C 0.95 ,--P--                        ...                   ..               v Pl1lCore Followw. C
                                                                                 -.     ,3DM 0.90 - --       - - -    - -   ---    -    - - -  -  - . _J. _........ . __._...._........ .,        ___i_......
0.75     - ------       ..
r     ;          r r1 0.75      0 6       - . . ..       -- - -                -- - -        ..   . .         -- -       -- -- --       --
a  ..8 .. . . .. ..  --    - -     --   --     -- --  --    -          ---    --  ......     .......      , , ...
                                                                                                                          .. .. l...
0.....                                               .........
0.50           <                 <,,
0            1          2             3        4             5            6             7           a             9   10 11  12 13  14   15 Exposure (GWdfST)
Figure SRXB-A-6 Cycle 24 MAPRAT MAPLHGR is a nodal parameter and requires no additional SIMULATE-3 module for its calculation. SIMULATE-3 and P11 have a fairly consistent bias over 3DM as shown in Figure SRXB-A-6-3. MAPRAT is typically the least limiting thermal limit at VYNPS.
Gamma TIPS SIMULATE-3 TIPS are generated to confirm the accuracy of the model relative to the plant TIP data set. SIMULATE-3 TIP comparisons to 3DM are created when TIP data becomes available.
The TIPS are produced in a Studsvik Scandpower post processing code known as S3post.
The following TIP output, Figure SRXB-A-6-4, depicts the comparison for VYNPS Cycle 24. The dashed line represents SIMULATE-3 and the solid line, 3DM. The radial, axial, and nodal RMS error values are calculated for each combined core average TIP. Any larger than expected

Attachment 2 BVY 05-072 Docket No. 50-271 Page 21 of 23 deviations are reviewed with plant reactor engineers, nuclear engineering analysis (NEA) engineers, and the NEA manager, per the core follow procedure, to investigate the discrepancy and, if necessary, take action. The low RMS errors provide a high confidence that the VYNPS SIMULATE-3 model is correctly calculating the power distribution.
Technical Specification Reactivity Anomaly Surveillance Technical Specification 3/4.3.E requires that at least every equivalent full power month, the critical rod configuration is compared to the expected configuration. These configurations are required to be within 1%Ak/k. The comparisons are performed using the eigenvalue calculated at that statepoint. As can be seen in Figure SRXB-A-6-5, for VYNPS Cycle 24 (current cycle),
the hot eigenvalue has compared well with the predicted eigenvalue, as well as the GNF core follow eigenvalue. Agreement between these eigenvalues provides confidence that the actual plant operation follows core design.
It should be noted that VYNPS Cycle 24 has not yet ended and data is provided only through a partial cycle.

VERMONT YANKEE CYCLE 24 SIMULATE TO 3DM TIP COMPARISON - JULY 05 I
U      Avor2go Azifl TIP TrnqcNodal Camparfson
                                                                        .I.                                           0 tRIN     mrPwo_           . rk     llW_. VW*        ,i,.yI. T              . Mm MrrLeTh   YvcmL1!24QvOc PL     0- W   WJ43I  6.1.11 -mJv        IWRI t
* wm      m     -0
- *A7LW a, t M             r*E1E     vWCYCLE2ACoOFCE     . OKflrm ci-4-A 4.717  bWr e       TF13  . Vmqur      -      0) CD10>
lFigur6 SRXB-A-6-4                                                                         -K   ) 0  ,
                        .Cycle 24 TIP Comparison Data l                                                           CA) " N    N

Attachment 2 BVY 05-072 Docket No. 50-271 Page 23 of 23 Cycle 24 ReactivityAnomaly Elgenvalue Curve 1.UZ 1.0186 1.0172
                '--      __l *__.5-.
1.3158 1.0102 1.013                                              ~
1.0102                                                                                   __                              _
1.0088         ~
1.0074                                       A 1.0046 1.0032 tx 1.0004 0.999 0.9976                                                        -
0.9962 0.9948                                                               _                              -.    -- _
0.9934                                                                                                              -
                                                                                                                                  *-  - -u  - -      -- _.6   -
0.992 0.9908 0.989 U.86b4I 0.98   I5 0      0.5      1     1.5      2    2.5   3   3.5  4    4.5    5    5.5    6     8.5   7. 7.5        8 8.5     9    9.5    10     10.5    11     11.5 EXPosure (GWdIST)
                                  - e MA      EAh...$b7cd)     -     Tch     Kpf              .TedhSpcI    .N                 - ..- -Aj*.4.73%
I  -_-z- -Ad    is  5%                    O.Lba3DKxff              _-GNFOT.          K..ff Figure SRXB-A-6-5                                                                        I Figure S.RXB-A-6-5 Cycle 24 Reactivity Anomaly Eigenvalue Curve Conclusion SIMULATE-3 gives high confidence through independent means that GNF methods are adequate to model the VYNPS reactor core. The model also gives high confidence that future design cycles are valid and sufficiently accurate to model EPU conditions.
The Technical Specification Reactivity Anomaly surveillance provides confidence that the actual plant operation follows core design.

BVY 05-072 Docket No. 50-271 Attachment 3 Vermont Yankee Nuclear Power Station Proposed Technical Specification Change No. 263 - Supplement No. 30 Extended Power Uprate Response to Request for Additional Information Overview of Steam Dryer Issues Total number of pages In Attachment 3 (excluding this cover sheet) is 4.

Attachment 3 BVY 05-072 Docket No. 50-271 Page 1 of 4 Overview of Steam Drver Issues By letter dated July 27, 2005 and meetings on May 9, 2005 and June 15-16, 2005, the NRC provided additional questions with respect to the VYNPS steam dryer analysis provided by Entergy in submittals dated March 31, 2005, April 5, 2005 and June 2, 2005. The purpose of this discussion is to provide an overview of Entergy's understanding of the fundamental issues to be addressed, the approach to address and the results. The major issues that have been raised by the staff relative to the VYNPS dryer analysis can be summarized as follows:
* Uncertainties that could exist in the Acoustic Circuit Analysis (ACA)
* Benchmarking and uncertainty of the computational fluid dynamics (CFD) model
* Development/evaluation of the 120% VYNPS load definition
* Adequacy of the GE scale model test (SMT) benchmark of the ACA
     . Applicability of knowledge gained from QC2 instrumented dryer tests to VYNPS assumptions, methods and conclusions The major issues above are the distillation of 129 individual questions. The purpose of this discussion is to provide an overview of Entergy's strategy in addressing the issues to assist the reviewers in understanding the context of the responses to the individual questions.
There is no change to the fundamental approach Entergy has taken, which consists of developing an analytical model that can convert empirical main steam system strain gage and pressure measurements to loads on the steam dryer. These dryer loads will then be compared to specific acceptance criteria in the power ascension dryer monitoring program.
An essential concept in this approach to providing reasonable assurance that steam dryer integrity will be maintained at EPU is the fact that there is significant margin between the dryer loads at 100% CLTP and the structural fatigue limits. Based on the absence of any significant detectable main steam system acoustic excitation at 100% CLTP and industry experience to date, VYNPS dryer loads at EPU conditions are still expected to provide significant margin to the structural fatigue limits. The Entergy power ascension dryer monitoring program contains acceptance criteria which will be used to verify that structural limits are not challenged during power ascension. By taking this approach, Entergy relies on detecting, and appropriately responding to, any unexpected phenomena that might occur during power ascension. This can be accomplished with reasonable certainty, rather than attempting to create a bounding 120% VYNPS EPU load prediction.
In response to the NRC questions, Entergy has updated the VYNPS steam dryer structural integrity analysis to incorporate the following:
1.) To address the issue of uncertainties that could exist in the ACA, Entergy calculated these uncertainties for both the methodology and the measurement techniques and applied them to the analysis to determine the overall impact on the steam dryer load definition. A significant amount of evaluation has been performed to more precisely determine the ACA model uncertainty. This information is described below and will be presented in response to the applicable RAI's. Three of the four elements of this uncertainty have been quantified and completed including methodology, strain gage, and venturi uncertainty and the remaining element, instrument location uncertainty, is in final review. The four uncertainty contributors will be combined to quantify a conservative total analysis uncertainty. It is expected that this will be submitted, along with all of the RAls applicable to ACA model uncertainty, for staff review by

Attachment 3 BVY 05-072 Docket No. 50-271 Page 2 of 4 8/4/05. Entergy anticipates the staff's review of the bases for the model uncertainty determination during the audit scheduled for 8/22/05.
2.) To address the issue of benchmarking and uncertainty of the computational fluid dynamics model, Entergy has provided benchmark references and performed sensitivity analyses for mesh size and turbulence intensity in order to quantify the CFD load uncertainty.
3.) To address the issue of development/evaluation of the 120% VYNPS load definition, Entergy developed 120% CFD transient loads and evaluated the impact of these loads on the structural analysis. Entergy also evaluated the characteristics of the ACA signatures based on VYNPS operating data from 80% to 100% CLTP and other industry experience.
4.) To address the issue of adequacy of the GE scale model test (SMT) benchmark of the ACA, Entergy performed additional benchmark data assessment including review of the model setup and operation and the impact of phasing sensitivity. All of the elements of this effort, with the exception of the phasing sensitivity evaluation that is in final verification, are complete. It is expected that this will be submitted, along with all of the RAls applicable to SMT benchmark, for staff review by 8/4/05. Entergy anticipates the staffs review of the bases for the SMT benchmark evaluation during the audit scheduled for 8/22/05.
5.) To address the issue of applicability of knowledge gained from QC2 instrumented dryer tests to VYNPS assumptions, methods and conclusions, Entergy reviewed the QC2 reports for any findings or conclusions that could adversely impact assumptions, methods, or techniques used in the VYNPS analysis. This review has verified that the approach implemented by Entergy for VYNPS remains valid, is reinforced by the QC2 79OMWe benchmark data, and the uncertainties applied to the VYNPS analysis contain sufficient conservatism. It is expected that results will be submitted, along with all of the RAls applicable to the quantification of the impacts of the QC2 test experience on the VYNPS dryer evaluation, for staff review by 8/4/05. Entergy anticipates the staffs review of the bases for the QC2 impact assessment during the audit scheduled for 8/22/05.
Acoustic Circuit Analysis Uncertainty The uncertainty associated with the ACA is a combination of measurement uncertainty and methodology uncertainty. Entergy evaluated the measurement uncertainty of the ACA by calculating the error associated with strain gage and pressure sensor measurement methods. The methodology uncertainty was assessed empirically by comparing predicted results against measured loads in two benchmark tests (SMT and QC2 full scale).
Additionally, the frequency response uncertainties observed in the benchmark tests were converted to load uncertainties and included in the methodology uncertainty. The impact of uncertainty associated with the instrument location relative to the dryer on the analysis was also evaluated.
CFD Model Benchmarking and Uncertainty The Fluent CFD methodology has been benchmarked against measured data and determined to provide an accurate representation of the modeled system. Included with the justification are several papers that describe benchmarking of the methodology.
To assess the uncertainty associated with CFD loads, a sensitivity analysis of the model

Attachment 3 BVY 05-072 Docket No. 50-271 Page 3 of 4 mesh size and inlet turbulence parameters was conducted. Results demonstrate that the model is relatively insensitive to further refinements of mesh size in the steam dome. The Large Eddy Simulation (LES)-appropriate fine mesh in the plenum region is sufficient to accurately compute the hydrodynamic forces acting on the dryer. The CFD model was also determined to be insensitive to large variations in inlet turbulence.
Development/Evaluation of 120% VYNPS Drver Load Definition The VY steam dryer loads are generated from two fluid models. An acoustic circuit model (ACA) developed by Continuum Dynamics Inc. was used to convert empirical data at between 80% and 100% CLTP to loads on the dryer face. This methodology will be used to measure acoustic loads above 100% CLTP. Because the ACA model may not completely account for low frequency hydrodynamic loads associated with vortex shedding, a computational fluids dynamics model (CFD) was developed by Fluent Inc. to evaluate these loads. Stresses from both models are combined in the VYNPS dryer structural analysis. The ACA loads are based on empirical measurements. The ACA methodology that converts these measurements to dryer loads was benchmarked to demonstrate the ability of the model to predict loads on the dryer not only at 100% CLTP, but at increasing steam velocities as power is increased. The Entergy approach recognizes that the ability to project acoustic loads at higher power levels could be challenged by the excitation of acoustic resonances that occur as flow changes. This is why the fundamental approach does not rely on -a predicted load curve above 100% CLTP. The performance criteria that Entergy has developed are based on reducing the code allowable stress for the component by the dryer load that is based on the CFD analysis, then determining what additional load is available for acoustic loading and establishing acceptance criteria that ensure acoustic loads are within the available margin. Entergy has evaluated CFD loads for both 100% CLTP and 120% EPU conditions and the maximum CFD loads are applied to dryer components.
Although the Entergy approach does not rely on a prediction of acoustic loads at EPU conditions, it should be noted that a reasonable extrapolation of ACA load data calculated between 80% and 100% CLTP predicts that there would continue to be a wide margin between dryer loads and structural limits at 120% EPU conditions. The absence of any significant detectable main steam system acoustic excitation at 100% CLTP and industry experience to date also support this expectation.

Attachment 3 BVY 05-072 Docket No. 50-271 Page 4 of 4 Scale Model Benchmark of the ACA Since the Entergy approach relies heavily on the ability of the ACA methodology to accurately convert empirical data collected from the Main Steam system to acoustic pressure loads on the face of the dryer, the principle objective in the SMT benchmark test was to demonstrate not only the ability of the ACA to model the transmission of anticipated main steam system signals, but also to demonstrate the ability of the model to detect new acoustic sources at EPU flows that are different from those detected at 100% CLTP or those observed at Quad Cities (QC). To accomplish the goal of demonstrating the ability to detect new acoustic sources, the Entergy benchmark program used broadband signals as inputs to excite a wide spectrum of acoustic frequencies, not just discrete frequencies that might be expected based on Quad Cities data. Additionally, a wide range of test conditions such as multiple flow rates (including a zero flow case) and two diverse broadband sound sources were used. Multiple runs were performed to verify repeatability of test data at each test condition. Finally, microphones were calibrated before, and calibration checked, after testing to ensure reliability of the data.
Many of the RAls focus on the impact of differences between the configuration of the SMT facility and VYNPS plant-specific characteristics. Because the purpose of the benchmark test was not to predict actual VYNPS loads, these minor differences in configuration are less important to establishing the validity of the ACA than they might be in an effort to predict actual dryer loads. The most important requirement of the scale model is that it accurately represents the actual acoustic features at VYNPS.
Applicability of QC2 Instrumented Dryer Tests/Benchmark to VYNPS The QC2 benchmark of the ACA performed at 79OMWe used the same version of the ACA model used in the SMT benchmark and for VYNPS (e.g. damping, acoustic speed, reflective boundary assumptions). Results are similar to SMT benchmark which is generally conservative.
The QC2 results at 79OMWe, along with the SMT benchmark results, were used to develop the ACA load uncertainty applied to the VYNPS dryer load definition.
Conclusions The CDI acoustic circuit analysis methodology used for the VYNPS dryer load definition is appropriate and reasonably accurate. The ACA methodology was benchmarked against the GE SMT and the Quad Cities 2 instrumented dryer. The QC2 benchmark at OLTP involved steam flow velocities similar to those for VYNPS at EPU power levels. The QC2 benchmark provides evidence that further supports Entergy's application of the ACA methodology for VYNPS' load definition. A review of the QC2 results supports use of the ACA loads combined with the CFD loads for the VYNPS loads and application of appropriate uncertainties is appropriate.
Entergy is confident that the FIV load definition and uncertainty values provide conservative loads and adequate margin to the acceptance limit curves. With the use of these curves, Entergy can ensure dryer structural integrity is maintained through the power ascension to EPU conditions and continued operation at full EPU power levels. This will be accomplished through careful monitoring of main steam system fluctuating pressures during controlled power ascension, and assessment against acceptance curves.

BVY 05-072 Docket No. 50-271 Attachment 4 Vermont Yankee Nuclear Power Station Proposed Technical Specification Change No. 263 - Supplement No. 30 Extended Power Uprate Response to Request for Additional Information Electrical and Instrumentation and Controls Branch RAls Total number of pages In Attachment 4 l (excludino this cover sheet) Is 6.       l

Attachment 4 BVY 05-072 Docket No. 50-271 Page 1 of 6 RESPONSE TO REQUEST FOR ADDITIONAL INFORMATION REGARDING APPLICATION FOR EXTENDED POWER UPRATE LICENSE AMENDMENT VERMONT YANKEE NUCLEAR POWER STATION PREFACE This attachment provides responses to the NRC Electrical and Instrumentation and Controls Branch's (EEIB) individual requests for additional information (RAIs) in NRC's letter dated July 27, 2005.1 Upon receipt of the RAI, discussions were held with the NRC staff to further clarify the RAI. In certain instances the intent of certain individual RAls may have been modified based on clarifications reached during these discussions. The information provided herein is consistent with those clarifications.
The individual RAls are re-stated as provided in NRC's letter of July 27, 2005.
Electrical and Instrumentation and Controls Branch (EEIB)
Electrical Engineerinq Section (EEIB-A)
RAI EEIB-A-1 The licensee's submittal dated March 24, 2005 (Supplement No. 25), provided revised station blackout (SBO) analyses for VYNPS in response to a finding documented in the NRC's inspection report dated December 2, 2004. The finding relates to the requirements in Title 10 of the Code of Federal Regulations (10 CFR) Section 50.63, "Loss of all alternating current power." Specifically, 10 CFR 50.63(c)(2) requires, among other things, that the time required for startup and alignment of the alternate alternating current (AC) power source and required shutdown equipment be demonstrated by test. The licensee has not indicated in their submittal that they are planning to do any kind of integrated test, with all parties involved, to show they can meet the 2-hour basis for starting and aligning the alternate AC power source (i.e., the Vernon Hydroelectric Station (Vernon Station)), should it have to be re-started during a regional blackout. The staff considers such a test to be critical to showing that appropriate procedures and protocols are in place to coordinate between the multiple entities that would be involved. Provide a discussion on how the licensee intends to meet the 10 CFR 50.63(c)(2) test requirements.
Response to RAI EEIB-A-1 Entergy's letter of March 24, 2005,2 described the regular, periodic testing currently conducted to demonstrate the ability of the alternate AC (AAC) source to power required electrical loads under a postulated station blackout event. The testing consists of two components:
' U.S. Nuclear Regulatory Commission (Richard B.Ennis) letter to Entergy Nuclear Operations, Inc.
(Michael Kansler), ORequest for Additional Information - Extended Power Uprate, Vermont Yankee Nuclear Power Station (TAC No. MC0761)," July 27, 2005 2 Entergy letter to U.S. Nuclear Regulatory Commission, "Vermont Yankee Nuclear Power Station, Technical Specification Proposed Change No. 263 - Supplement No. 25, Extended Power Uprate -
Station Blackout and Appendix RAnalyses," BVY 05-030, March 24, 2005

Attachment 4 BVY 05-072 Docket No. 50-271 Page 2 of 6
    . Actual black start of the AAC source by TransCanada in accordance with ISO New England operating procedure OP-11, "Black Start Capability Testing Requirements."
This testing is conducted and documented annually.
Surveillance testing of the Vernon Tie in accordance with Entergy.VYNPS procedure OP-4142, 'Vernon Tie and Delayed Access Power Source Backfeed Surveillance."
Performance of this test demonstrates actual ability to power required SBO loads.
This testing is conducted and documented during each refueling outage.
These tests perform all actions required to restart the Vernon Hydroelectric Station (VHS) and, upon delivery of power to the Vernon Tie, the reenergization of a VYNPS 4KV bus. The only step not performed is the opening and closing of certain breakers in the interfacing switchyard as this would cause an unnecessary blackout to the general public. These breaker manipulations are performed in a continuously manned regional operations center and controlled by system procedures. Completion of these external activities will take less than five minutes of the two hour coping duration and are considered simplistic actions not requiring periodic validation.
Entergy recently held discussions with TransCanada, the owner/operator of the VHS, and the regional grid control center regarding procedural requirements and communication protocols for a postulated SBO event. These communications have resulted in system restoration procedure improvements and have served to promote a better understanding of the expectations relative to Entergy's reliance on the VHS during a SBO.
Entergy has established administrative controls to assure performance of a once per operating cycle tabletop review of the procedures that complete the actions to repower a VYNPS 4KV bus from the VHS. Pursuant to discussions with co-host REMVEC, a system-wide annual tabletop review will take place in October 2005. During this meeting Entergy will lead a tabletop review of all actions required to support the restoration of 4KV AC to VYNPS.
This review will discuss the interfaces with the operator of VHS and the regional grid operator to verify that roles and responsibilities and timelines are understood and that there have been no changes that would impact the assumption in the VYNPS SBO coping strategy.
Entergy will also provide the participants with additional insights regarding offsite power issues for nuclear power stations including plant response to and consequences of a SBO.
RAI EEIB-A-2 Since the operators of the Vernon Station are not Entergy personnel or Entergy contractors or vendors, and the station is not manned 24 hours per day, 7 days per week, the NRC staff requires additional information to have reasonable assurance that the operators will respond to the station as required and perform their duties in a reliable manner as needed to provide the alternate AC power source to VYNPS. Please address the following issues:
a) Are specific operators designated "on-call" to respond to the Vernon Station, as needed, during periods when the station is unmanned?
b) Are the operators subject to any fitness-for-duty requirements?

Attachment 4 BVY 05-072 Docket No. 50-271 Page 3 of 6 c)    Are the operators, responsible for responding to the station when it is unmanned, required to remain within a certain distance from the Vernon Station?
d) The licensee's submittal states that since the Vernon Station is designated a "black start" facility under arrangements with the regional grid operator, this designation requires that the facility be capable of being black-started within 90 minutes after the operator is notified. From the onset of a regional grid collapse, during a period for which the Vernon Station is unmanned, discuss all assumptions regarding the time required for the operator to reach the station (e.g., adverse weather conditions, distance traveled), and the time required for the operator to perform the necessary actions to black start the station.
Response to RAI EEIB-A-2 Entergy has discussed these questions with TransCanada, the owner/operator of the Vernon Hydroelectric Station (VHS). Insupport of its commitment with ISO-NE to provide black start capability, TransCanada currently has four personnel that work out of the VHS during the normal day shift and report to that location. These personnel may be used to support activities at the other TransCanada units. These workers are trained to return the unit to service and make it available to be connected to the system. They have an on-call supervisor assigned who covers the Vernon, Bellows Falls and Wilder hydroelectric facilities and carries a pager. For off hour events, TransCanada's control station, which is continuously manned and located in Wilder VT, would contact the on-call supervisor who would call in the necessary support personnel to restart the VHS. Any alarm indication that the VHS has tripped off line is treated as a critical alarm and would prompt the call-in immediately upon receiving the alarm. Based on their experience, which includes off hours events in which the VHS needed to be re-started, TransCanada indicated that they had restarted the unit within the required ISO-NE response timeframe. They also indicated that they had not experienced situations where personnel were unavailable to support restart of the unit.
TransCanada indicated that they did have company policies that include a fitness-for-duty program. Although random drug testing is not performed on all personnel, supervisory observations that identified a potential for alcohol or drug abuse would lead to drug testing. It is TransCanada's expectation that the on-duty supervisor be fit to perform this duty when on-call.
A key assumption in the coping evaluation time line is that the personnel in the Wilder control station would be aware of a regional black-out almost immediately. During off hours, which maximizes response time, the control center would contact the on-call supervisor who would contact and dispatch personnel to restart the VHS. Given current agreements, testing practices and past experience, the timeline assumes that this will be completed within 90 minutes.
For a regional blackout, ISO-NE would direct the system restoration and order the transmission owner to close the switchyard breakers supplying VYNPS. These breakers can be operated remotely by the transmission owner. VYNPS' operators would then close the breakers supplying power to our emergency bus. The VYNPS breakers are operated from the VYNPS main control room, and these actions can occur very quickly. The coping study

Attachment 4 BVY 05-072 Docket No. 50-271 Page 4 of 6 uses a two hours duration which bounds the actions discussed above.
To account for additional unforeseen circumstances (e.g., adverse weather beyond assumed travel time) the coping study is done with conservative inputs that provide additional margin.
For example, should the SBO event occur during a winter snow storm that could delay VHS startup, the conservatisms in heat sink temperature (which assumes peak summer allowable temperature) would allow for additional coping time.
Entergy believes that use of a two hour coping time together with conservatisms inherent in the coping analysis, in addition to conservatively estimated response times, provides reasonable assurance that the VHS will be available to support mitigation of a SBO event.
RAI EEIB-A-3 Provide a discussion on the agreements between Entergy and other entities to bring the Vernon Station online from black-start conditions in order to provide electric power to VYNPS during an SBO event (i.e., whether there are formal written agreements supported by written procedures). If there are no formal written agreements, discuss why there is reasonable assurance that the alternate AC power source will be available, as needed, consistent with the time frames assumed in the revised VYNPS SBO analyses.
Response to RAI EEIB-A-3 Vermont Yankee has an Alternative AC Source Agreement dated July 31, 2002, with Green Mountain Power (GMP), which is the retail electricity provider in the area, to make available at the point of interconnection between the Vernon switchyard and the Vernon Tie up to 3 MW of energy from the Vernon Station during an emergency affecting Vermont Yankee.
This agreement requires GMP to take all reasonable steps to keep the Vernon Tie energized at all times. GMP has in turn entered into an agreement, dated as of July 31, 2002, with USGenNE (the operator of the Vernon Station at the time) and now assigned to supply this power to Vermont Yankee in an emergency, and a Service Agreement for Network Integration Transmission Service Between New England Power Company and Green Mountain Power Corporation, effective July 31, 2002, which commits New England Power (the owner of the transmission facilities at the Vernon Station) to keep the Vernon Tie energized during normal utility operations and to make reasonable efforts to keep the line energized during emergency situations, subject to ISO, NEPOOL and REMVEC requirements. TransCanada has affirmed that they are committed under tariff to provide black start capability of the VHS to ISO-NE.
Both the NEPOOL and REMVEC procedures state that "the most critical power requirement after a blackout is the assurance of reliable shutdowns of nuclear generators, and that expeditious restoration of alternative off-site power sources to nuclear units is imperative to promote the continued reliability of shutdown operations.' NEPOOL Operating Procedure No. 6, System Restoration, App. A (System Restoration Guidelines); REMVEC II Operating Procedure No. 6, System Restoration, App, B (Guidelines and Philosophies).]
As a backup to local indication available to grid operators of a regional blackout, VYNPS procedure OT 3122 "Loss of Normal Power" Appendix A 'Station Blackout Procedure" directs

Attachment 4 BVY 05-072 Docket No. 50-271 Page 5 of 6 operators to immediately contact the regional grid control center to initiate a blackstart of the VHS if the Vernon Tie is unavailable due to a regional grid blackout. The regional grid control center actions are directed by Operating Procedure OP-6 "System Restoration." This directs certain hydro station operators (including the VHS) to initiate blackstart procedures, and upon notification that the units are started, provide instructions to align power to WNPS and to communicate when these actions are complete to the VYNPS control room. The owner of the VHS has a procedure for the actual blackstart.
RAI EEIB-A-4 With respect to the battery capacity requirements during an SBO event, verify that sufficient DC power is available, under worst case conditions during the two hour coping period, to close the 4160 volt breakers associated with the alternate AC power source.
Response to RAI EEIB-A-4 A review of the worst case scenario has been performed by adding the 6 amp load of the Vernon Tie breaker closing for a full minute, for conservatism, to the end of the 2 hour duty cycle. The run was compared to calculation WC-2154, Rev. 0, MCC 4 and shows no change to the required battery capacity. This additional load also has no effect on end voltage.
RAI EEIB-A-5 Provide a discussion regarding the changes required to plant procedures for SBO coping (e.g., which procedures will be changed and brief discussion of the changes to be made).
Also, address when the operator training on the revised procedures is expected to be completed.
Response to RAI EEIB-A-5 Procedure changes for SBO coping include the following:
* Operating Procedure OP 2124, " Residual Heat Removal System" has been revised for training" to include direction on how and when to place the second RHRSW pump per RHR heat exchanger in service when placing the torus cooling mode of RHR in service.
* Operational Transient procedure OT 3122, "Loss of Normal Power" has been revised to direct operators to immediately contact the regional grid control center to initiate a blackstart of the VHS if the Vernon Tie is unavailable due to a regional grid blackout. In addition, the procedure has been "revised for training" to commence a cooldown within one hour of the SBO event, and when power is restored, to place two RHRSW pumps in service per OP 2124 (see above). Also, a note was added about exceeding the drywell air temperature of 280'F for a short period of time without exceeding the 280'F drywell shell temperature.

Attachment 4 BVY 05-072 Docket No. 50-271 Page 6 of 6 Training on the changes to OP 2124 and OT 3122 is currently underway. Once the training cycle is complete (scheduled for September 1, 2005), the procedures will be revised and issued for use.
Various operating, surveillance and administrative procedures will be revised to incorporate a higher condensate storage tank inventory limit as either a precaution or an administrative limit by October 15, 2005.

BVY 05-072 Docket No. 50-271 Attachment 5 Vermont Yankee Nuclear Power Station Proposed Technical Specification Change No. 263 - Supplement No. 30 Extended Power Uprate Response to Request for Additional Information Mechanical and Civil Engineering Branch l Total number of pages in Attachment 5 l (excludina this coversheet) is 28.        l

Attachment 5 BVY 05-072 Docket No. 50-271 Page 1 of 28 RESPONSE TO REQUEST FOR ADDITIONAL INFORMATION REGARDING APPLICATION FOR EXTENDED POWER UPRATE LICENSE AMENDMENT VERMONT YANKEE NUCLEAR POWER STATION PREFACE This attachment provides responses to the NRC to Mechanical and Civil Engineering Branch's (EMEB) individual requests for additional information (RAis) in NRC's letter dated July 27, 2005.1 Upon receipt of the RAI, discussions were held with the NRC staff to further clarify the RAI. In certain instances the intent of certain individual RAls may have been modified based on clarifications reached during these discussions. The information provided herein is consistent with those clarifications.
The individual RAis are re-stated as provided in NRC's letter of July 27, 2005.
Mechanical and Civil Engineering Branch (EMEB)
Component Integrity and Testing Section (EMEB-A)
Civil and engineering Mechanics Section (EMEB-B)
RAI EMEB-B-41 On Page 22 of Attachment 6 to Supplement 26, GE discusses the modified steam dryer analysis results. The acoustic and CFD pressure loadings were calculated up to 100% of the current licensed thermal power (CLTP) level. For the square type dryer similar to the ones at VYNPS and Quad Cities, no significant cracking was identified during operation at 100% OLTP.
However, this type of steam dryer failed when Quad Cities Units 1 and 2 operated at EPU conditions. Analysis up to 100% CLTP does not demonstrate dryer integrity at EPU conditions.
Provide the evaluation of VYNPS modified steam dryer for EPU conditions.
Response to EMEB-B-41 Entergy's modified steam dryer analysis relies on measured main steam system data for input to the acoustic circuit analysis in order to generate dryer loads. The CLTP dryer stresses calculated based on those loads, incorporating uncertainty, show substantial margin to the ASME fatigue limit. In addition, the CFD analysis performed assuming 120% EPU steam flow conditions does not show a significant increase in hydrodynamic loads. Entergy agrees that analysis up to 100% CLTP does not conclusively demonstrate dryer integrity at EPU conditions.
The square type dryer installed at VYNPS is similar not only to Quad Cities, but to other BWR's.
One such BWR that has been operating at EPU conditions since 1998 recently conducted a comprehensive dryer inspection that found no indications of significant cracking. Although no significant cracking was identified at Quad Cities or VYNPS during operation at 100% OLTP, VYNPS is the only square hood plant to perform a comprehensive internal and external inspection for indications prior to operation at EPU conditions. Additionally, Quad Cities
'U.S. Nuclear Regulatory Commission (Richard B. Ennis) letter to Entergy Nuclear Operations, Inc.
(Michael Kansler), "Request for Additional Information - Extended Power Uprate, Vermont Yankee Nuclear Power Station (TAC No. MC0761)," July 27, 2005

Attachment 5 BVY 05-072 Docket No. 50-271 Page 2 of 28 instrumented dryer and main steam system measurements show evidence of acoustic source excitation at 100% OLTP, whereas VYNPS main steam measurements do not. Industry evidence suggests that the increased steam flow rate associated with EPU does not create new acoustic phenomena, rather it exacerbates phenomena that existed at OLTP. Based on the absence of any significant acoustic excitation at 100% CLTP and industry experience to date, VYNPS dryer loads at EPU are still expected to provide wide margin to the structural fatigue limits.
Entergy plans to collect data during the controlled and methodical power ascension for use in evaluating EPU dryer loads with respect to defined performance criteria and to verify that structural limits are not challenged during power ascension. The performance criteria that Entergy has developed are based on reducing the code allowable stress for the component by the dryer load that is based on the CFD analysis, then determining what additional load is available for acoustic loading and establishing acceptance criteria that ensure acoustic loads are within the available margin. Entergy has evaluated CFD loads for both 100% CLTP and 120% EPU conditions and the maximum CFD loads are applied to dryer components. By taking this approach, Entergy relies on detecting, and appropriately responding to, any unexpected phenomena that might occur during power ascension. This can be accomplished with reasonable certainty, rather than attempting to create a bounding 120% VYNPS EPU load prediction.
RAI EMEB-B-49 On Page 29 of Attachment 7 to Supplement 26, CDI compares the calculated steam dryer loads for Quad Cities Unit 2 and VYNPS. CDI states that the added energy content present in Quad Cities Unit 2 is the result of distinct deterministic mechanisms that exist at the feed flow rates at which Quad Cities Unit 2 operates, and that these mechanisms are not excited at the much lower feed flow rates at which VYNPS operates. Discuss the basis for the assumption that the higher steam flow rates to be achieved during EPU operation at VYNPS, combined with the plant-specific steam system configuration, will not result in the excitation of distinct deterministic mechanisms similar to Quad Cities Unit 2.
Response to EMEB-B-49 There is no explicit or implicit assumption made in the discussion cited by this RAI regarding excitation of deterministic mechanism at VYNPS as a result of EPU operation. The comparison of VYNPS acoustic circuit analysis loads with Dresden and Quad Cities in Attachment 7 to Supplement 26 is based on use of measured main steam system data for input. Figure 15 of  provides a peak load comparison between VYNPS and QC2 at OLTP. The comparison discussion asserts that the QC2 peak loads are a result of deterministic mechanisms that are excited at the unit's OLTP feed flow rates. Rated feed flow of QC2 at OLTP is 50% greater than VYNPS. This difference in rated feed flow is cited as the justification for why no such sources are detected at VYNPS OLTP operating conditions.
As stated in response to RAI EMEB-B-41, Entergy agrees that analysis up to 100% CLTP does not conclusively demonstrate dryer integrity at EPU conditions. Entergy will collect data during

Attachment 5 BVY 05-072 Docket No. 50-271 Page 3 of 28 power ascension for use in evaluating EPU dryer loads to verify that structural limits are not challenged during power ascension. By taking this approach, Entergy relies on detecting, and appropriately responding to, any unexpected phenomena that might occur during power ascension.
The discussion cited in this RAI reflects that QC2 has deterministic excitation mechanisms at OLTP. Quad Cities instrumented dryer and main steam system measurements confirm acoustic source excitation at 100% OLTP. Based on the absence of any significant acoustic excitation at 100% CLTP, VYNPS dryer loads at EPU are still expected to provide wide margin to the structural fatigue limits.
RAI EMEB-B-73 With regard to Section 3.1.2 of Attachment 6 to Supplement No. 26, the vortex shedding pressure loading was input into ANSYS from a Fluent LES compressible flow simulation of the dryer at 100% of CLTP conditions. Entergy should provide documentation that benchmarks and validates the CFD code's ability to predict vortex shedding in the complex flows of the dryer.
Response to EMEB-B-73 Exhibits EMEB-B-73-1 through 3, provided in Attachment 11, are papers that benchmark the CFD codes prediction capability through the comparison of LES simulation results measured data.
EMEB-B-120 With regard to Attachment 1 (Fluent Final Report TM-675, "CFD Modeling of the Vermont Yankee Steam Dryer Phase - II) to Supplement 29, please provide a more detailed discussion of the lessons-learned during the Phase I analysis of the steam dryer along with more information on how these lessons-learned were addressed by the Phase II analysis.
Response to EMEB-B-120 The Phase I analyses consisted of an incompressible URANS simulation on the full upper vessel geometry and an incompressible LES simulation on a clipped geometry. Comparisons of these two results yielded an improved understanding of the flow field and flow structures within the dryer. Low frequency shifts in the flow field were noted as a result of the steam jet interaction occurring at the inlet to the dome. The lesson learned was not to neglect the interaction between the dome and the plenum. Strong vortices were observed near the steam line inlets as the steam turned into the steam lines. These vortices were noted to rapidly make small shifts in position with time, thereby causing large pressure gradient fluctuations on the opposing vertical faceplate. Large pressure gradients within the vortices were considered unrealistic and assumed to be caused by flow incompressibility. The resulting lesson learned was to avoid modeling the fluid (steam) as incompressible. Variations in mass flow at the outlets indicated potential steam line flow interaction. High frequency features were difficult to characterize with the limited data collected for this phase. The resulting lesson learned was to

Attachment 5 BVY 05-072 Docket No. 50-271 Page 4 of 28 collect sufficient data in order to characterize high frequency features.
Lessons learned from the Phase I analysis that were addressed in the Phase II analyses included: coupled plenum-dome flow domain to account for transients in dome, flow compressibility, extension of outlet steam lines to common outlet plenum. Surface pressure (difference) data was collected for all dryer walls.
RAI EMEB-B-121 On page 1 of Attachment 1 to Supplement 29, the Fluent report states that the model geometry was modified to raise the overall accuracy of the Phase I model. What specific measure or basis is used to support the improvements in accuracy?
Response to EMEB-B-121 Accuracy improvements were expected because some assumptions from the Phase I model were removed and the geometry was better presented in Phase II.
Assumptions in Phase I that were removed:
* Flow is incompressible. This assumption has an affect on the pressure gradient across the vortices in front of the nozzles. Using a compressible gas would allow the fluid to compress and expand and thus to absorb a portion of centrifugal force in a strong vortex in a density change.
* Interaction between dome and plenum is negligible. A decoupled plenum simulation with a stagnant boundary condition does not reflect the dynamic interaction between the flow in the dome and the plenum.
Better represented geometric components:
* Inclusion of dams
* Curvature at steam line nozzle
* Extended pipe network
* 2" gap around steam dryer The modifications made in Phase II have led to some changes of the simulation results that represent the problem better than in Phase I. Those modifications and the resulting improvements included:
* pressure difference across the vortex in front of steam lines is lower because of compressibility effects;
* large scale, low frequency oscillations in the dome are observed and feed into the plenum at the face plates;
* curved steam line nozzles allow flow to enter pipes more smoothly;
* some flow in the 2" gap has been observed; and
* some flow rate oscillations in steam lines were observed.

Attachment 5 BVY 05-072 Docket No. 50-271 Page 5 of 28 EMEB-B-1 22 On page 1 of Attachment 1 to Supplement 29, the Fluent report states that the model geometry was modified to raise the overall accuracy of the Phase I model. Are there any further refinements of the geometric details that could be made? Please discuss any remaining geometric simplifications that are present in the model and the basis for assuming these simplifications do not impact the predictions.
Response to EMEB-B-122 The following simplifications were made:
: 1. Material thickness was neglected. The thickness of the gussets and other interior walls is small compared to the size of the dryer. By neglecting the thickness the flow area has increased by a small amount. Neglecting the wall thickness is common practice in CFD simulations when the thickness is small compared to the size of the flow path.
: 2. Hooks and bolts were neglected. Small features such as hooks and bolts create some local disturbances but have very little impact on the main flow pattern. These features can be neglected if they are small compared to flow path.
: 3. Dryer vane geometry was not included. Instead a porous media model was used to represent the pressure loss in the vane bank. The dryer vane banks are located relatively far from the face plate. The influence of the vane banks on the face plate loading was deemed to be small because of its distance to face plate. The question of the influence of the porous media and the dryer vane bank on the hydrodynamic loads at the face plate is discussed in RAI EMEB-141.
: 4. Steam line simplifications:
* The steam lines were shortened downstream of the strain gauges, any turns were neglected.
* The valve assembly was simplified.
* Details in the steam lines such as branch lines, dead ends, and contractions were neglected.
The main purpose of including the steam lines was to obtain pressure data at the location of the strain gauges and to couple the steam lines to enable communication between the steam lines.
It was assumed that the details of the flow in the steam lines do not impact the face plate loading because the steam line flow is downstream of the face plate. Turbulent disturbances cannot be convected upstream except through pressure waves and acoustic modes. However, capturing acoustic modes was not the focus of the analysis.

Attachment 5 BVY 05-072 Docket No. 50-271 Page 6 of 28 EMEB-B-1 23 On page 1 of Attachment 1 to Supplement 29, the Fluent report states that "the Phase I study showed that an LES model could be successfully applied to the dryer model." Please define the success criteria used to support this conclusion.
Response to EMEB-B-123 The LES model was applied successfully in the sense that the simulation was completed within a practical time frame. The time and length scales are such that the volume around the area of interest could be resolved with an appropriate mesh while being able to obtain a solution for one scenario within one or two months. Furthermore, the simulation provided some qualitative insight into the flow pattern in the steam dryer.
EMEB-B-1 24 On page 1 of Attachment 1 to Supplement 29, the Fluent report states that improvements were made to the physical models. Assuming that certain assumptions are made in a typical CFD simulation, please discuss the remaining major assumptions and limitations of the physical models used in relationship to the flow physics and conditions expected in this specific problem.
Please describe how computational limitations affected your modeling approach.
Response to EMEB-B-124 Remaining assumptions of the physical model include:
* Steam represented as an ideal gas
* Liquid fraction can be neglected
* Rigid walls (as opposed to elastic walls), no two-way fluid-structure-interaction Viscous heating was neglected in the simulation because the amount of energy dissipated compared to the thermal energy in the fluid is relatively small. Viscous heating is considered for applications with very high shear rates, such as torque converters and other rotating machinery, which is not the case in this simulation. Pressure work and turbulent kinetic energy were not included in the energy equation because the pressure variations and the turbulent kinetic energy were assumed to be small compared to the thermal energy in the fluid based on general experience.
Thermal boundaries were treated as adiabatic.
The pressure variations in the domain are small. Therefore, the ideal gas law should be sufficient to describe the density sufficiently in the vicinity of the operating pressure. The molecular weight was adjusted to match the real gas density of steam at the operating point.
The liquid fraction was small after the flow went through the dryer vane bank and therefore was neglected.

Attachment 5 BVY 05-072 Docket No. 50-271 Page 7 of 28 The CFD analysis did not consider any effect of the structural deformations on the flow.
Potentially, there could be some fluid structure interaction. The structure can transmit some of the energy absorbed from the fluid back to the fluid.
Computational limitations primarily affected the model mesh size, time step, and extent of the geometry.
EMEB-B-1 25 On page 2 of Attachment 1 to Supplement 29, the Fluent report states that a "highly refined mesh, consistent with LES requirements" was used in a specific domain. Discuss the specific rationale used to determine the mesh size in the LES-appropriate mesh region and provide the specific computations used in this determination. Where appropriate, compare mesh sizes to relevant length scales, such as the Taylor or Komolgorov micro-scales or the boundary layer thickness (wall y+). Discuss the size of the mesh in the rest of the model regions (noted to be too large for ideal LES simulations) in terms of relative size of the mesh compared to an ideal LES mesh to give an idea of the relative coarseness of this mesh.
Response to EMEB-B-125 The mesh used in the LES section of the Phase II model was taken from the Phase I LES study.
Results from the Phase I URANS calculation were used to define mesh requirements for the Phase I LES model. Meshing guidelines for LES calculations require that computational cells be sized such that they are larger than the Taylor length scale and smaller than the integral length scale. The Taylor length can be estimated as (10vk/-)" 2 and the integral length scale can be estimated as C.k0'2/I. Based on the estimated k and E fields from the URANS calculation, typical values in the plenum for Taylor and integral length scales are 0.03 inches and 1.5 inches, respectively. Typical cell size used in the plenum for the LES section of the model was 0.65 inches. Only one plenum was modeled with this cell size, which was sufficiently representative for the dryer region of greatest interest. In the opposite plenum, the cell size was approximately 1.5 inches. The cell size in the dome ranged from the approximately 1.5 inches to 6 inches. These cell sizes are appropriate for these regions of the model for the purpose of capturing the large scale unsteadiness and flow pattern in the dome.
EMEB-B-1 26 On page 3 of Attachment 1 to Supplement 29, the Fluent report discusses a comparison of LES and unsteady Reynolds-Averaged Navier-Stokes (URANS) in an attempt to demonstrate that the LES model would not produce significantly different results from the URANS model. On page 12, it is noted that Figures 8 through 15 show that the LES model sufficiently represents the URANS solution. In light of the complex jet interactions in the upper dome of the reactor, explain the basis for using the URANS solution as a reference point given the lack of specific validations for this complex flow pattern. In addition, discuss the sufficiency argument for the LES to URANS comparison in the context of Figures 14 and 15 which show a significant variation in velocity magnitude as well as the spatial variation in the velocity.

Attachment 5 BVY 05-072 Docket No. 50-271 Page 8 of 28 Response to EMEB-B-126 Phase II model was seen as an improvement over Phase I. The URANS model showed some unsteady behavior that was important to capture in Phase II.
The acceptance criteria for using the LES approach included:
* Mean flow contours are similar.
* LES model is less diffusive than URANS and therefore shows more pronounced gradients and higher peaks.
* LES model does not show any results that are apparently unphysical.
Some variations between the LES model and the URANS approach were expected. The URANS solution is sufficient for use as a reference point because it was the best available description of the flow features in the dome at the time of the comparison.
The LES model showed similar unsteady behavior in the dome based on visual inspection of velocity magnitude animations.
The LES to URANS comparison is sufficient based on the similarity in flow features. Flow features that are similar in both the URANs and LES cases can be found in Figures 8 through 15.
Pressure contours in Figure 10 and Figure 11 are similar. The difference between minimum and maximum pressure is similar.
Figure 12 and Figure 13: High pressure spot in same location. There is a layer of low pressure on the surface of the dryer. It is more pronounced in the LES results.
Figure 14 and Figure 15: There are differences in peak velocities, however, the spatial distribution of the flow is quite similar. Flow features that can be found in both URANS and LES include:
* streak of high velocity in 1 o'clock position along dome,
* horizontal streak of high velocity at about 1/4 height of the dome (from outermost dryer jet),
* streak of high velocity pointing in 11 o'clock direction.
The LES results show an additional feature on the right hand side above the horizontal streak of high velocity.
EMEB-B-127 On page 5 of Attachment 1 to Supplement 29, the hybrid mesh is described and it is noted that the flow domain is coupled by resolving the entire domain for the LES simulation. Conceptually, discuss the significance of the coarse mesh on this "boundary condition" at the edge of the LES appropriate mesh region. Given the strong dependence of LES solutions on the applied boundary conditions, discuss the potential for sensitivity of the results to variations in the coarse mesh size.

Attachment 5 BVY 05-072 Docket No. 50-271 Page 9 of 28 Response to EMEB-B13127 The influence of boundary conditions on the result in the flow domain depends in general on the forces that act upon the flow in the domain. If the forces in the domain are very strong then the boundary conditions are less important. If the forces are weak then the boundary conditions have a big impact on the flow field. For example, in a pipe flow the inlet boundary condition will prolong for a long distance. If the distance is long enough a fully developed pipe flow evolves. If there is a turn in the pipe then the flow is likely to separate on the inner wall. The flow profile downstream of the turn is more influenced by the turn than by the inlet profile.
For the sensitivity of the solution on the 'boundary condition' at the edge of the LES appropriate mesh regions, two aspects can be considered: first, the distribution of flow entering the LES region, second the content of turbulent kinetic energy in the flow.
Figure EMEB-B-127-1 shows the velocity distribution on a plane located a small distance in front of the face plate. Velocity magnitude contours from a steady state RANS simulation and time averaged velocity magnitude contours from the LES simulation (100 % load) are shown.
From these figures it can be seen that the LES case represents the flow distribution well compared to the RANS study. The spatial distribution and the velocity magnitudes are in good agreement including the flow distribution in the dome. From this perspective, the boundary conditions used in the LES model are sufficient.
Potentially, there could be an influence of the solution in the LES region on the 'flow condition' in the area where the mesh transitions from the coarse mesh to the fine mesh. Vortices that are generated upstream could be convected downstream and could influence the solution in the LES region. If the dome region had been adequately resolved, more turbulence would be generated in the dome and carried over to the face plate.
The distribution of turbulent kinetic energy on the mesh transition face is shown in Figure EMEB-B-127-2 for both the LES and steady state RANS cases. The mesh employed in the steady-state RANS case is adequate to resolve local turbulent kinetic energy (TKE) production in the dome. The steady state RANS simulation shows the distribution of TKE on this face. In both cases, the TKE carried over from the dome to the face plate is smaller than the TKE generated at the face plate. The peak TKE levels are higher in the LES simulation.

l I
(Iime=7.4521e+00)
Contours of Mean Velocity Magnitde (nWs)                                Jul 28, 2005 contours of Velocity Magntude (rnts)
FLUENT   6.2 (3d, segregated, LES, unsteady)
C 0
To 03c0r I Figure EMEB-B-127-1: Velocity Magnitude Contours from RANS (left) and LES (right)
J XD Z Opson(
0o0
                                                                                                                                    -h EDX-02 tO 0 =:
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Attachment 5 BVY 05-072 Docket No. 50-271 Page 11 of 28 Potentially, there could be an influence of the solution in the LES region on the 'flow condition' in the area where the mesh transitions from the coarse mesh to the fine mesh. Vortices that are generated upstream could be convected downstream and could influence the solution in the LES region. If the dome region had been adequately resolved, more turbulence would be generated in the dome and carried over to the face plate.
The distribution of turbulent kinetic energy on the mesh transition face is shown in Figure EMEB-B-127-2 for both the LES and steady state RANS cases. The mesh employed in the steady-state RANS case is adequate to resolve local turbulent kinetic energy (TKE) production in the dome. The steady state RANS simulation shows the distribution of TKE on this face. In both cases, the TKE carried over from the dome to the face plate is smaller than the TKE generated at the face plate. The peak TKE levels are higher in the LES simulation.

0 0
0 Contours from    a CD co>
Figure EMEB-B-127-2: Steam Dome and Dryer Face Turbulent Kinetic Energy RANS (left) and LES (right)                                                           O~Co naD 0.' . CD
                                                                                      -4   N) 0 ,
x-4M -
00 -r%N)CA

Attachment 5 BVY 05-072 Docket No. 50-271 Page 13 of 28 EMEB-B-1 28 On page 5 of Attachment 1 to Supplement 29, the hybrid mesh is described and it is noted that the flow domain is coupled by resolving the entire domain for the LES simulation. Discuss the spatial and temporal filtering that could be expected to occur as the result of excess diffusion in the coarse tetrahedral region in the upper dome. What is the potential impact of not having resolved potentially higher frequency spatial and temporal variations on the conceptual "input boundary" to the LES appropriate mesh region? Could the lack of higher frequency information in the final results be partially attributed to a lack of resolution of these frequencies in large portions of the domain?
Response to EMEB-B-128 The effect of the coarse mesh in the dome is a reduced ability of the simulation to resolve local turbulence (in the dome) with high frequency information. At a mean flow velocity of about 10 m/s and a mesh size of about 6" or 0.15 m (as can be found in the dome) the maximum frequency based on the Nyquist criterium is: mean flow velocity/ (4* mesh size) =
10/(4*.15)=16.6. The Nyquist criterium requires a minimum resolution of a sinusoidal signal with 4 points. With the assumption that a disturbance is convected with the mean flow velocity the above relationship can be found. Therefore, the high frequency portion of the spectrum is underresolved in the dome.
Figure EMEB-B-128-1 shows the comparison of the LES results and the steady state RANS.
The LES and RANS simulation generate about the same level of turbulence in the dome.
The time averaged velocity magnitudes from a steady state RANS and the LES simulation in the dome are shown in Figure EMEB-B-128-2. The LES results represent the mean flow patter well compared to the RANS solution in terms of spatial distribution and peak velocities. The coarse mesh also leads to excess numerical diffusion which causes the jets to diffuse more quickly.
The flow velocities in the plenum are larger than in the dome. The lengths scales are smaller.
Therefore, significant contributions at higher frequencies are more likely to be generated in the plenum than in the dome. Figure EMEB-B-128-3 shows the PSD of 62.2 Hz at the face plate.
The signal this PSD is based on includes pressure fluctuations from acoustic and hydrodynamic sources. The hydrodynamic sources are represented by the local peaks.

   *3.00ef01                     -                               3.ODe401 i   2.70e401                                                     2.'l eM0l
    -Ml5e01                                                      2a.:ie l 2.25e401i                                                    2.25e~Ol 2.1M     I2.1                                De4.0 lT.95erl                                                     I l5eMil 1.20e.0i                                                     i52eO0l HOMO                                                         9C00e            K        e 577.0e O                                                       7.50e;00 6.0l0                    Fs         atFe                                                                                               t 3.00e+00                                                [H50~OM 4 za x0 Y j                                        _  _   _   _ _ _   _  _  _   _ _  _  _  _   _ _ _ _ _ _ _ _ _  _  _   _ _  _   _   _
CnorofIejm 7451eO                      Jul 29, 2005 Contours of Turbulerre Kinett Energy (k)(m212)                              Jul 29, 2005 ES, nstedy)FLUENT FWEN        6.2(3d segegatd,                                                    6.2 3d, segregated, rke) 0 0
CD Figure EMEB-B-128-1: Steam Dome Turbulent Kinetic Energy Contours from RANS (left) and LES (right)                                      lz C0      0
                                                                                                                                                    -, NM 0

0 0
0 T   77 Figure EMEB-1-128-2: Velocity Magnitude Contours from RANS (left) and LES (right) CD0 % 0 I                                                         cen en o n O0 KN1
                                                                                      ? 0N flCD1 K) -4  -4

Attachment 5 BVY 05-072 Docket No. 50-271 Page 16 of 28
* S.                                                                 Sl Figure EMEB-B-128-3: PSD of 62.2 Hz at the Dryer Face Plate EMEB-B-129 On page 5 of Attachment 1 to Supplement 29, it is stated that "the success of obtaining an LES solution on the hybrid mesh hinged on the capacity for the LES model to sufficiently resolve the flow field on the coarse mesh." The concept was tested by comparing the LES and URANS solutions. How can one determine that the LES model sufficiently resolves the flow on the coarse mesh without comparison to data? Is comparing a specific turbulence model to another turbulence model considered to be a validation?
Response to EMEB-B-129 Relevant engineering data required of the simulation is based on the resolution of <-70 Hz signals. The LES model simulation is therefore required to be able to resolve the flow characteristics in this frequency range. A sufficiently fine mesh and time step size is required for this resolution. The model was set up to meet and exceed this resolution requirement in the regions where the turbulence production was considered to be significant, i.e., in the plenum region.
In the dome, a coarser mesh was used. The comparison study, as well as a follow-on steady-state RANS study, showed that the LES model on the coarser mesh in the dome region was

Attachment 5 BVY 05-072 Docket No. 50-271 Page 17 of 28 able to resolve the mean flow features and low frequency oscillations. The steady-state RANS study verifies this quantitatively as discussed in responses to RAI 127 and 128.
In the absence of experimental data, a comparison of results with that of a second or third turbulence model along with mesh refinement studies is an appropriate validation. The LES model is the most universal turbulence model available and applicable to most flow regimes.
Therefore, comparing results obtained from different turbulence models can support the validation of results.
EMEB-B-130 On page 9 of Attachment 1 to Supplement 29, it is noted that a central differencing scheme is used. What is the order of accuracy for the solution and what impact does the boundary treatments have on this accuracy? Does the accuracy at the boundaries impact the overall solution accuracy? Were higher order schemes attempted? What is the expected sensitivity of the solutions to the central differencing scheme and the wall treatment accuracy?
Response to EMEB-B-130 Second-order upwind schemes have been widely used for RANS computations. Numerical diffusion introduced by upwind schemes is accepted in RANS simulations because the eddy viscosity is orders of magnitude larger than the molecular viscosity. In LES simulation the subgrid scale viscosity is much smaller than the eddy viscosity in a RANS simulation. The numerical diffusion can therefore overwhelm the physical diffusion when upwind schemes are used. To improve the accuracy of LES solutions, Fluent has developed the bounded central differencing scheme that is less diffusive and dissipative than upwind schemes. The central differencing scheme is the recommended scheme for LES problems.
Sensitivity to wall treatment has been discussed in response to RAI 133.
References to differencing schemes can be found in:
Kim, S.E., Mohan, L.S.: "Prediction of Unsteady Loading on a Circular Cylinder in High Reynolds Number Flows Using Large Eddy Simulation", 24th Intemational Conference on Offshore Mechanics and Arctic Engineering, Greece, 2005 Kim, S.E.: Large Eddy Simulation Using Unstructured Mesh and Dynamic Subgrid-Scale Turbulence Models", AIAA Paper 2004-2458 EMEB-B-131 On page 9 of Attachment 1 to Supplement 29, it is noted that a dynamic Smagorinsky-Lilly model for subgrid scale stresses is used. What is the expected sensitivity of the solutions to variations in the subgrid scale stress model?

Attachment 5 BVY 05-072 Docket No. 50-271 Page 18 of 28 Response to EMEB-B-131 Upon review of this RAI modeling question, it was discovered that the dynamic Smagorinsky model was not enabled during the simulation as was reported in Attachment 1 to Supplement
: 29. Regardless, the question of sensitivity is addressed in response to this RAI.
The subgrid scale stresses are determined, in part, by the Smagorinsky constant, Cs. The default constant is set to 0.1. The dynamic Smagorinsky model allows for a variation in this constant based on local scales and can vary from 0 to 0.23. If the constant is larger, the subgrid scale stresses are larger.
On coarse meshes, the computed subgrid scale stresses are likely to be of the same order or smaller than numerically derived stresses (numerical diffusion) and hence the sensitivity to the subgrid scale model is small.
On adequately resolved meshes, the subgrid scale stress term becomes important. For free shear flows, as is found in the VYNPS steam dryer simulation, the Smagorinsky constant is generally found to be at the higher end of the spectrum (0.2). Since a lower value was actually used (0.1), the local stresses were computed to be lower, and the result would likely be a conservative over-estimation of local pressure forces.
EMEB-B-1 32 On page 9 of Attachment 1 to Supplement 29, it is noted that the turbulence level of the boundary condition is not important because the flow will develop as it traverses through the dryer. Please comment on the accuracy in the turbulence level as it reaches the LES appropriate mesh. Does the turbulence develop appropriately in the porous models within the.
dryer? Does the turbulence develop appropriately in the relatively coarse tetrahedral mesh in the upper dome? Are any sensitivities completed to justify the statement that the turbulence level at the inlet "does not have any impact" on the flow in the dryer?
Response to EMEB-B-132 A steady-state RANS simulation was performed to assess the sensitivity of the resolution of the flow field to turbulence inlet boundary conditions. In one case, a low level of turbulence was applied (1% intensity with a viscosity ratio of 2). In a second case, a high level of turbulence was imposed (50% intensity with a viscosity ratio of 20000). Distributions of turbulent kinetic energy (TKE) from these two cases can be seen in the figures below. Figure EMEB-B-132-1 compares the TKE near the inlets. The left figure clearly shows a much higher input of turbulent energy. The overall impact of the change in the dome can be seen in Figure EMEB-B-132-2 where it can be seen that the distribution of turbulent kinetic energy within the dome is comparable between the high and low inlet turbulence cases.

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CD
                                               -z Dryer In T ence Ine  0 0
TKE for 50% and 1%Assumed                       OT n

0 0
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0O IC Figure EMEB-B-132-2: Resulting Dome TKE for 50% and 1% Inlet          o
                                                                        ,CD D
0 to C-N)MC

Attachment 5 BVY 05-072 Docket No. 50-271 Page 21 of 28 EMEB-B-133 On page 10 of Attachment 1 to Supplement 29, it is noted that y+ values are in the range of 3,000 to 15,000. What are "best practice" values for y+ in an LES simulation? Assuming the y+
values used are larger than "best practice" guidelines would require, discuss the potential impact of the wall treatment (wall boundary condition) on LES predictions.
Response to EMEB-B-133 In general, best practice y+ values for LES simulations should be on the order of 1. However, these constraints can be relaxed for conditions when wall-bounded effects are unimportant. For the present study, free shear flows dominate the region: e.g., free jets, vortical structures impinging on the vertical faceplates and turning into the steam lines.
Boundary layers are important to resolve in external aerodynamic application such as the fuselage of an aircraft, or in flows that are dominated by friction near a wall, such as pipe flow.
In case of the VYNPS steam dryer simulation, the large eddies are dictated by the arrangement of the geometry and the flow distribution. These eddies are much larger than the thickness of the boundary layer. The size of eddies generated in the boundary layer are expected to be at the same order of magnitude as the boundary layer thickness and associated time scales are proportionally smaller.
EMEB-B-134 On page 10 of Attachment 1 to Supplement 29, the time step size is listed. Please provide further details on the determination of the time step size for this type of LES analysis.
Response to EMEB-B-134 The time step size was selected to resolve the hydrodynamic effects of the smallest, resolved eddies. It was not the objective of this simulation to model the acoustic effects. An appropriate time scale can be estimated from local velocity and length scales with an assumed Courant number (UAtIAx) of 1. By imposing the Courant number constraint, disturbances traveling at a velocity U can be resolved a distance Ax away in a time span of At. Local mean velocities within the plenum did not exceed 40 in/sec. With a cell size of -0.65 in (0.016m), the corresponding At is -0.0004 seconds.
EMEB-B-135 On page 11 of Attachment 1 to Supplement 29, the determination of pressure differences is discussed. Are pressure fluctuations expected on the back side of the plate? Would this model determine pressure fluctuations in this region given the location of the inflow boundary condition

Attachment 5 BVY 05-072 Docket No. 50-271 Page 22 of 28 and the non-prototypical influence of the porous media regions? What impact might these assumptions have on the reported pressure fluctuations?
Response to EMEB-B-135 As a mass flow inlet boundary condition type was used, pressure fluctuations could arise at the inlet to enforce the constant mass flow rate condition. These fluctuations were not expected to be significant, but were nevertheless accounted for in the calculation for pressure differences.
The presence of the porous media would tend to isolate any disturbances in pressure within the regions separated by the media. Turbulence generation is expected to be low upstream of the dryer vanes because of the low flow velocity in this region. Neither location of the inflow boundary condition nor porous media assumptions have significant impact on pressure fluctuations.
EMEB-B-1 36 Given the complex nature of the jets and interactions highlighted in Figure 32 of Attachment 1 to Supplement 29, discuss the impact of the coarse tetrahedral mesh on these flows and its potential subsequent impact on the "boundary" to the LES appropriate mesh region.
Response to EMEB-B-136 See the response to RAI 127.
EMEB-B-1 37 On page 17 in the first paragraph of Attachment 1 to Supplement 29, it is implied that improvements to accuracy are obtained locally in the specific region that is resolved using the LES appropriate mesh. The basic question is whether or not additional fluctuations would be present if the entire domain were modeled with an LES appropriate mesh. Please summarize the basis for assuming that the prediction in a local region is accurate when the predictions in the surrounding regions can be assumed not to be accurate. This discussion should consider the importance of boundary conditions in LES predictions.
Response to EMEB-B-137 The basis for using a coarse mesh in the dome was the assumption that the bulk of the production of the turbulence occurs in the plenum. Figure EMEB-B-137-1 shows the turbulent production distribution in a steady state RANS simulation. The highest production of turbulence occurs in the plenum.

     .77e5                          __-7e05 L 1.8i5WS                                                                              .68e405 e05              ___U15e405
   - l51esO               _      l'                                                     151e405
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     '-24* 51.24e;05 U.5ei.                                                                             i.l5eit5 e5                                .,                                          1.i. 05 e                                                                             9.75e044 886e*                                                                              86e04 7.974]                                                   7.97e:44 6.20e*6.20e044 5.32e                                                                              5_32ef04
__.7s4j=    ..                          S               t-         --  4,3es04 4                                                                               8B6es03 537e-06                                                                      ._     537e6                     \  \X       \       ,
Gortours of Produdbn of k pbm-s3)                                    Jul 29, 2005 Contours of Prtucdbn of kgVm-s3)                   Jul 29, 2005 6.2 (, segregated, rke) l FWENT                                                                     segregate, rke)
FWENT 6.2 3d, 0
0 U CD      >
I                                                                                      (no     <'
Figure EMEB-B-137-1: RANS Turbulent Production Distribution                                                  w Cn o 3
                                                                                                                                                  -F     30   CD I-4-4 Co   - K) M

Attachment 5 BVY 05-072 Docket No. 50-271 Page 24 of 28 EMEB-B-1 38 With regard to the work described in Attachment 1 to Supplement 29, discuss in general the applicability of a single CFD prediction (no sensitivity studies for key assumptions or parameters) in a flow domain where no data are used for validation or benchmarking.
Response to EMEB-B-138 Generally, CFD is applied as a complementary tool to other analytical techniques, including experimentation. CFD is useful in extracting absolutes (validation quality) and for sensitivity studies. Ideally, results from a model are validated with specified experimental data before perturbations to the model are studied. In the absence of experimental data, sensitivities to modeling parameters are warranted. Given the immense CPU requirements associated with LES modeling, such sensitivity studies can become impractical. Under these conditions, the modeler must rely on using sound engineering judgments and best practices to achieve a solution developed from the governing equations of motion. Supported by complementary studies, single CFD predictions can still be considered a useful tool. Sensitivity studies were performed to investigate the impacts of varying mesh size and inlet turbulence intensity. These studies were performed using a steady state RANS version of the VYNPS steam dryer CFD model. Results of these sensitivity studies are contained in responses to RAI's 127, 132, 136 and 140. Benchmarking of Fluent LES methodology against experimental data is described in three papers enclosed as Exhibits EMEB-B-73-1 through 3.
EMEB-B-140 On page 9 of Attachment 1 to Supplement 29, a uniform inlet turbulence intensity of 2% is specified. The inlet dissipation rate (which along with turbulence intensity defines characteristic flow length scales) is not specified. Entergy is requested to discuss the sensitivity of the simulated steam dryer loading to the inlet turbulence intensity and dissipation rate.
Response to EMEB-B-140 The LES model was set-up such that it used a constant mass flow rate at the inlet. The mass flow inlet boundary condition does not allow specifying any perturbations of the inlet velocity.
Velocity variations could have occurred because of pressure and density fluctuations. An inlet dissipation rate does not apply to the LES model.
The response to RAI 132 highlights the lack of sensitivity to inlet turbulent boundary conditions for this particular case using a RANS model.
EMEB-B-1 41 On page 9 of Attachment 1 to Supplement 29, a porous media is described which simulates the head loss across the dryer. Entergy is requested to describe the effects of the porous media model on the turbulence intensity and characteristic length scales of the jet flow emanating from

Attachment 5 BVY 05-072 Docket No. 50-271 Page 25 of 28 the top of the dryer. Also, the sensitivity of the simulated steam dryer loading to the porous media should be provided.
Response to EMEB-B-141 The porous media model does not have any direct influence on the turbulence intensity and length scales. Through the pressure loss term in the momentum equations the porous media models the flow resistance of the dryer vane bank. The flow distribution adjusts accordingly.
Turbulence is generated only from stream line curvature and shear in the flow.
The loss coefficient in flow direction is derived from known pressure drops.
EMEB-B-142 The PSDs averaged over the dryer face plate are compared for 100% of CLTP and 120% of CLTP conditions (or EPU conditions) in Figures 45 and 46 of Attachment 1 to Supplement 29.
The comparison shows significant increases in loading at EPU condition at frequencies of about 32, 45, and 62 Hz. The amplifications strengthen near the MSL inlets. Entergy indicated on page 48 that low-frequency increases are due to hydrodynamic effects, and high-frequency increases are due to acoustic effects. Entergy also indicated that the time step size used in the compressible LES solutions is too large to resolve acoustic effects accurately. Entergy is requested to explain whether the peaks at 32, 45, and 62 Hz are due to acoustic amplification, and if so, explain how the time step size affects the amplitudes of the peaks. The evaluation should be based on an assessment of significant acoustic modes of the MSLs and dome contributing to pressure fluctuations on the steam dryer.
Response to EMEB-B-142 In order to verify the acoustic nature of frequencies larger than 30 Hz a Fast Fourier Transform (FFT) was performed on time series of pressures at the dryer surface, including the top, the face plate and the skirt. The FFT generates coefficients of cosine and sine functions for each frequency under consideration. The signal at each frequency is represented as a_n*cos(wt)+b.n*sin(wt). Figures EMEB-B-142-1, EMEB-B-142-2 and EMEB-B-142-3 show the a n -coefficients for 32Hz, 45.5 Hz and 62.2 Hz, respectively. The data is taken from the 120%
load case.
The contours plots indicate that the dome and the gap around the skirt interact to form acoustic modes. The 32Hz mode oscillates in vertical direction. The 45.5 Hz and 62.2Hz modes oscillate from one plenum to the opposite plenum. There could also be a component in circumferential direction in the gap around the dryer. The face plates are impacted by these acoustic modes as can be seen from the contour plots.
Accurate predictions of the propagation of pressure waves can be obtained if the time step is chosen such the Courant number based on the speed of sound is close to unity. The time step chosen was based on the fluid flow velocity rather than the speed of sound and is about 10 times larger than required for tracking sound waves. With a large time step pressure waves are propagating at the speed of sound. However, a sharp pressure front could be diffused and

Attachment 5 BVY 05-072 Docket No. 50-271 Page 26 of 28 attenuated with distance from its origin. The large time step relative to acoustic phenomena could have a damping effect on any acoustic signal.

Attachment 5 BVY 05-072 Docket No. 50-271 Page 27 of 28
-B-1i4Z-Z: an coefficient Tor 40.OHz

Attachment 5 BVY 05-072 Docket No. 50-271 Page 28 of 28 figure LMtL-ti-14Z-J: an co

BVY 05-072 Docket No. 50-271 Attachment 6 Vermont Yankee Nuclear Power Station Proposed Technical Specification Change No. 263 - Supplement No. 30 Extended Power Uprate Response to Request for Additional Information GE Affidavit for Attachment 1 Total number of pages in Attachment 6 l (excludina this cover sheet) Is 3.        l

General Electric Company AFFIDAVIT I, Robert E. Gamble, state as follows:
(I) I am Manager, ESBWR, General Electric Company ("GE"), have been delegated the function of reviewing the information described in paragraph (2) which is sought to be withheld, and have been authorized to apply for its withholding.
(2) The information sought to be withheld is contained in Enclosure I of GE letter, GE-VYNPS-AEP-393, Margin in GE Analytical Methods Supporting Vermont Yankee EPU Submittal, dated July 29, 2005. The proprietary information in Enclosure 1, Basis for Margin in GE Analytical Methods Supporting Vermont Yankee EPU Submittal, is delineated by a double underline inside double square brackets. Figures and large equation objects are identified with double square brackets before and after the object.
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The development of the evaluation process along with the interpretation and application of the analytical results is derived from the extensive experience database that constitutes a major GE asset.
(9) Public disclosure of the information sought to be withheld is likely to cause substantial harm to GE's competitive position and foreclose or reduce the availability of profit-making opportunities. The information is part of GE's comprehensive BWR safety and technology base, and its commercial value extends beyond the original development cost. The value of the technology base goes beyond the extensive physical database and analytical methodology and includes development of the expertise to determine and apply the appropriate evaluation process. In addition, the technology base includes the value derived from providing analyses done with NRC-approved methods.
The research, development, engineering, analytical and NRC review costs comprise a substantial investment of time and money by GE.
The precise value of the expertise to devise an evaluation process and apply the correct analytical methodology is difficult to quantify, but it clearly is substantial.
GE's competitive advantage will be lost if its competitors are able to use the results of the GE experience to normalize or verify their own process or if they are able to claim an equivalent understanding by demonstrating that they can arrive at the same or similar conclusions.
The value of this information to GE would be lost if the information were disclosed to the public. Making such information available to competitors without their having been required to undertake a similar expenditure of resources would unfairly provide competitors with a windfall, and deprive GE of the opportunity to exercise its competitive advantage to seek an adequate return on its large investment in developing these very valuable analytical tools.
I declare under penalty of perjury that the foregoing affidavit and the matters stated therein are true and correct to the best of my knowledge, information, and belief.
Bxecutcd on this 30th day of July 2005.
Robert E Gamblc General 1lecetic Company

BVY 05-072 Docket No. 50-271 Attachment 7 Vermont Yankee Nuclear Power Station Proposed Technical Specification Change No. 263 - Supplement No. 30 Extended Power Uprate Response to Request for Additional Information Probabilistic Safety Assessment Branch Total number of pages in Attachment 7 (excludina this cover sheet) is 5.

Attachment 7 BVY 05-072 Docket No. 50-271 Page 1of 5 RESPONSE TO REQUEST FOR ADDITIONAL INFORMATION REGARDING APPLICATION FOR EXTENDED POWER UPRATE LICENSE AMENDMENT VERMONT YANKEE NUCLEAR POWER STATION PREFACE This attachment provides responses to the NRC Probabilistic Safety Assessment Branch's (SPSB) individual requests for additional information (RAls) in NRC's letter dated July 27, 2005.1 Upon receipt of the RAI, discussions were held with the NRC staff to further clarify the RAI. In certain instances the intent of certain individual RAls may have been modified based on clarifications reached during these discussions. The information provided herein is consistent with those clarifications.
The individual RAIs are re-stated as provided in NRC's letter of July 27, 2005.
Probabilistic Safety Assessment Branch (SPSB)
Containment and Accident Dose Assessment Section (SPSB-C)
RAI SPSB-C-47 The response to RAI SPSB-C-41 is not clear as to why required NPSH values, based on lower pre-EPU suppression pool temperatures, satisfy pump requirements at the higher EPU suppression pool temperatures. Does the increased suppression pool temperature affect the magnitude of the required NPSH or the time period at a given required NPSH, or both?
Response to RAI SPSB-C-47 The pump NPSH requirements are based on the lower pre-EPU suppression pool temperatures only to the extent that the pump vendor used the minimum available NPSH based on those temperatures and no overpressure credit as their starting point in evaluating the core spray and residual heat removal pumps capability. Attachment 5 of calculation VYC-0808 2 , the vendor's summary report, recommended time-dependent minimum NPSH requirements which bounded the minimum available NPSH values. These requirements are not dependent on any particular scenario or suppression pool temperature. Therefore, neither the magnitude nor the time period at a particular NPSH is affected.
The minimum NPSH requirements continue to be met at the higher pool temperatures associated with EPU conditions by crediting containment overpressure. Therefore, the license amendment requested credit for containment overpressure for DBA-LOCA and ATWS events.
1 U.S. Nuclear Regulatory Commission (Richard B. Ennis) letter to Entergy Nuclear Operations, Inc.
(Michael Kansler), "Request for Additional Information - Extended Power Uprate, Vermont Yankee Nuclear Power Station (TAC No. MC0761)," July 27, 2005 2 Calculation VYC-0808, Rev. 8 was provided to the NRC in VYNPS Proposed Change No. 263, Supplement No. 18, BVY 04-106, October 5,2004

Attachment 7 BVY 05-072 Docket No. 50-271 Page 2 of 5 RAI SPSB-C-48 In response to RAI SPSB-C-39, Entergy stated that the emergency operating procedure NPSH curves are independent of specific event scenarios. However, credit was taken for the minimum available NPSH curves. Aren't these curves event-specific to the large break loss-of-coolant accident (LBLOCA) at pre-EPU conditions? Why is this acceptable?
Response to RAI SPSB-C-48 The emergency operating procedure (EOP) NPSH curves are based on the minimum required NPSH from Attachment 5 of calculation VYC-0808. The minimum required NPSH is a characteristic of the pump in question and is not dependent on any particular event. As noted above in the response to RAI SPSB-C-47, the required NPSH values are only based on the lower pre-EPU suppression pool temperatures in the sense that the pump vendor used the available NPSH at those lower temperatures as the starting point in evaluating the core spray and residual heat removal pumps' capability to function under those conditions.
For any specific event, operators refer to the core spray and residual heat removal NPSH curves included in the EOPs to confirm operation within NPSH limits. The EOP curves allow the operators to account for flow, pool temperatures, and containment overpressure.
RAI SPSB-C-49 Supplement 25, dated March 24, 2005, states that credit for containment accident pressure is no longer required in determining adequate available NPSH for the residual heat removal and core spray pumps for the postulated Appendix R fire and the SBO scenario. Was it necessary to take credit for the minimum available NPSH curves in reaching this conclusion?
Response to RAI SPSB-C-49 Yes. The NPSH curves used in the evaluations for the postulated Appendix R fire and the SBO scenario are based on the minimum required NPSH from Attachment 5 of calculation VYC-0808. These are the same curves used to evaluate NPSH margin for current licensed thermal power operating conditions.

Attachment 7 BVY 05-072 Docket No. 50-271 Page 3 of 5 RAI SPSB-C-50 Provide a comparison of the emergency core cooling system (ECCS) pump flow rates for the 10 CFR 50.46 LBLOCA analysis with the ECCS flow rates used for the short-term and long-term NPSH analyses.
Response to RAI SPSB-C-50 For conservatism, minimum flows are assumed for the 10 CFR 50.46 analysis and maximum flows are assumed for short-term and long-term NPSH analysis. The 10 CFR 50.46 values (as presented below in Tables SPSB-C-50-1, SPSB-C-50-2, and SPSB-C-50-3) are defined as a function of the differential pressure between the reactor vessel and containment.
Table SPSB-C-50-1 One RHR Pump injecting into one recirculation loop (LPCI mode) 10 CFR 50.46 Analysis Pressure Difference (psid)                            Pump Flow (gpm) 260                                              0 20                                            6,300 0                                             6,550 RHR pump NPSH evaluations for a single pump, regardless of mode (LPCI, containment spray, torus cooling) assume a maximum flow rate of 7,400 gpm constant.
Table SPSB-C-50-2 Two RHR Pumps injecting into one recirculation loop (LPCI mode) 10 CFR 50.46 Analysis Pressure Difference (psid)                            Pump Flow (gpm) 260                                              0 20                                           11,251 0                                           11,765 RHR pump NPSH evaluations for two pumps injecting into one recirculation loop assume a maximum flow rate of 14,200 gpm constant.
Table SPSB-C-50-3 One CS Pump injecting into the reactor vessel 10 CFR 50.46 Analysis Pressure Difference (psid)                           Pump Flow (gpm) 260                                             0 20                                           2,800 0                                           4,000 CS pump NPSH evaluations assume 4,600 gpm constant for short-term (i.e. first ten minutes of the DBA-LOCA event) and 3,500 gpm constant for long-term (beyond the first ten minutes of the

Attachment 7 BVY 05-072 Docket No. 50-271 Page 4 of 5 DBA-LOCA event, the pump is assumed to be throttled).
RAI SPSB-C-51 Describe how instrument uncertainty is taken into account, either in the EPU containment analyses (both peak pressure and minimum pressure for NPSH), surveillance procedures, or in some other way.
Response to RAI SPSB-C-51 Instrument uncertainty is factored into the input parameters in certain key areas. See Table SPSB-C-51-1 below. The input parameters defining initial conditions are based on design basis or technical specification limits.
Table SPSB-C-51-1 Parameter         Input                           Instrument Uncertainty Value Initial Reactor       1950        A 2% uncertainty is assumed. This accommodates all the Power                MWt          instrument uncertainties that provide input to the plant heat balance calculations, which are used to monitor compliance with the licensed operating power.
Initial              900 F        This is the Technical Specification maximum. The required suppression pool                  condition is less than 87.30 F to accommodate worst case temperature                       instrument uncertainty.
Initial              68,000       This is the Technical Specification minimum. The required suppression pool     cu ft        condition is greater than 68,255 cu ft to accommodate volume                            instrument uncertainty.
Instrument uncertainty is also applied to the RHR and core spray flow rates assumed in the NPSH analysis. Surveillance procedures assure the pumps meet their Technical Specification requirements.      The acceptance criteria in the procedures accommodate instrument uncertainties. The instrument uncertainties are applied to assure the pumps meet minimum head and flow requirements. Since NPSH margins are minimum when the pump flow rates are maximized, the NPSH analysis is based on assuming the instrument uncertainties are applied in the opposite direction.

Attachment 7 BVY 05-072 Docket No. 50-271 Page 5 of 5 RAI SPSB-C-52 The effects of a loss-of-ventilation on a SBO event are discussed in Supplement No. 25, , Section 2.3. Please provide the references listed in that section for NRC staff review (i.e., References 3, 6, 7, 12, 13, and 16).
Response to RAI SPSB-C-52 Those sections of the requested calculations that deal with the loss of ventilation are provided in 1 as Exhibits SPSB-C-52-1 through SPSB-C-52-6. The information provided in the Exhibits is summarized in Table SPSB-C-52-1 below.
Table SPSB-C-52-1 Exhibit No. Reference                                    Discussion SPSB-         No.
C                 13     VYC-0886, Rev. 2, Calculation Change Notice (CCN) Number 4, is provided. CCN 4 specifically addresses the impact of power uprate on VYC-0886, Rev. 2. The relevant information is on pages 5 and 6 of Attachment A.
2           6      VYC-1347, Rev. 0, calculates the heat-up in the main steam tunnel upon loss of ventilation for two scenarios (1) main steam and feedwater isolated and (2) main steam and feedwater not isolated. Scenario (1) applies to SBO. This is called Case MST1 in the calculation, and the results are discussed in Section 2.2.1 on pages 20 and 21. The results show that the HPCI and RCIC isolation setpoints would not be exceeded until approximately 18 hours after the loss of ventilation.
3           7      VYC-1 502, Rev. 0, calculates the heat-up in the control room due to loss of ventilation for an Appendix R event. The heat loads for the Appendix R event bound those for the SBO event.
4          12      VYC-2405, Rev. 0, calculates the drywell heat-up for the SBO event and was done specifically for power uprate operating conditions and a 2 hour coping time.
5          13      VYC-2279, Rev. 0, evaluated the effect of power uprate on ambient space temperature during normal operation. Only those portions that deal with the steam tunnel are provided. The calculation shows that ambient steam tunnel temperatures during normal operation will increase by less than 1 degree F due to power uprate.
6          16      Only those parts of Operating Procedure OP-2192, Rev. 31, "Heating, Ventilating, and Air Conditioning System," related to the loss of control room ventilation are included.

BVY 05-072 Docket No. 50-271 Attachment 8 Vermont Yankee Nuclear Power Station Proposed Technical Specification Change No. 263 - Supplement No. 30 Extended Power Uprate Response to Request for Additional Information Plant Systems Branch l Total number of pages in Attachment 8 1 (excludina this cover sheet) is 12.       I

Attachment 8 BVY 05-072 Docket No. 50-271 Page 1 of 12 RESPONSE TO REQUEST FOR ADDITIONAL INFORMATION REGARDING APPLICATION FOR EXTENDED POWER UPRATE LICENSE AMENDMENT VERMONT YANKEE NUCLEAR POWER STATION PREFACE This attachment provides responses to the NRC Plant Systems Branch's JSPLB) individual requests for additional information (RAls) in NRC's letter dated July 27, 2005. Upon receipt of the RAI, discussions were held with the NRC staff to further clarify the RAI. In certain instances the intent of certain individual RAls may have been modified based on clarifications reached during these discussions. The information provided herein is consistent with those clarifications.
The individual RAls are re-stated as provided in NRC's letter of July 27, 2005.
Plant Systems Branch (SPLB)
Balance of Plant Section (SPLB-A)
RAI SPLB-A-25 Section 6.3.1 of Attachment 6 of the application dated September 10, 2003, indicates that in the unlikely event of a complete loss of spent fuel pool (SFP) cooling capability, the SFP will reach the boiling temperature in six hours. This conclusion does not appear to be consistent with the information that is provided for the alternate cooling system (ACS) in Updated Final Safety Analysis Report (UFSAR) Section 10.8 which indicates that upon a loss of all SFP cooling, boiling will occur in two-to-three days. Please explain this apparent inconsistency.
Response to RAI SPLB-A-25 The apparent inconsistency is due to two different scenarios-one assumes a batch off-load, the other a full core off-load.
Section 10.8.4 of the current UFSAR indicates that upon a loss of all SFP cooling, boiling will occur in two-to-three days. Section 10.8.4 also indicates that this time is based on a fuel pool heat load of 7.8 x 106 BTU/hr. This heat load is based on a batch off-load.
The six hour time to boil value for extended power uprate (EPU) is based on a full core off-load.
Following extended power uprate, the time to boil value for a batch off load is approximately two days.
' U.S. Nuclear Regulatory Commission (Richard B. Ennis) letter to Entergy Nuclear Operations, Inc.
(Michael Kansler), "Request for Additional Information - Extended Power Uprate, Vermont Yankee Nuclear Power Station (TAC No. MC0761)," July 27, 2005

Attachment 8 BVY 05-072 Docket No. 50-271 Page 2 of 12 RAI SPLB-A-26 UFSAR Section 10.8 indicates that the deep basin has a water capacity of 1.48 million gallons and that a water inventory of 1.45 million gallons is sufficient to assure seven days worth of ACS cooling capability. Please explain in detail how this conclusion was reached for post-EPU operation, quantifying all water additions and losses that are assumed to occur over this seven-day period along with how these values were determined, and how much inventory is required at the end of seven days to satisfy pump net positive suction head (NPSH) requirements.
Response to RAI SPLB-A-26 Water losses are due to evaporation, drift and external factors (e.g., pipe drainage during ACS setup, silt buildup and collapse of non-seismic portions of cooling structure). No makeup to the cooling tower is assumed for the entire duration of the ACS event. For ACS mode design basis heat loads and meteorological conditions, total seven day losses due to evaporation and drift were calculated to be 1,040,000 gallons and 29,000 gallons, respectively. Total losses due to external factors were calculated to be 266,800 gallons. The cooling tower has a minimum capacity of 1,451,700 gallons. Therefore, remaining inventory at the end of seven days of ACS operation is 116,000 gallons, equating to an inventory margin of 8%.
To maintain positive margin on NPSH over the entire seven day ACS event, two of the four RHRSW pumps are removed from service after 48 hours. The resultant reduction in suction header friction losses compensates to some degree for the reduction in suction static head caused by evaporative losses. At the end of four pump operation at 48 hours, minimum basin level is calculated to be 10 ft. above the first stage impeller of the RHRSW pumps. For the maximum pump flowrate at this point (approximately 2,105 gpm), required NPSH is 19.4 ft., and available NPSH is calculated to be 26.1 ft., yielding a margin of 6.7 ft., or 34%. At the end of seven days of ACS operation with only two pumps in operation, minimum basin level is calculated to be 4.2 ft. above the first stage impeller of the RHRSW pumps. For the maximum pump flowrate at this point (approximately 2,144 gpm), required NPSH is 19.9 ft. and available NPSH is calculated to be 29.9 ft., providing a margin of 10 ft., or 50%.
RAI SPLB-A-27 The response to RAI SPLB-A-17 in Supplement No. 28 indicates that there is sufficient margin between the minimum transient reactor feedwater pump (RFP) suction pressure and the current RFP suction pressure trip setpoint to ensure RFP operation during normal operation and the loss of one condensate pump transient. This does not appear to be consistent with the information provided in the RAI response that indicates that the condensate pumps only have a 7% flow margin to pump runout conditions, which would suggest that two condensate pumps operating are not sufficient to ensure RFP operation following the loss of one condensate pump.
Please explain the basis for concluding that continued RFP operation is assured following the loss of one condensate pump.

Attachment 8 BVY 05-072 Docket No. 50-271 Page 3 of 12 Response to RAI SPLB-A-27 For clarification, the response to RAI SPLB-A-17 in Supplement No. 28 relates to the impact on current equipment and system margins due to EPU implementation.
The 7% condensate pump flow (CP) margin discussion applies to normal EPU operation with all three CPs operating. The 7% flow margin is determined by comparing total condensate system flow during normal EPU conditions and the maximum possible condensate system flow demand assuming minimum total condensate and feedwater system resistance. That is, at EPU, the CPs will be required to produce 7% more flow to sustain the maximum possible system flow.
This is well within the capabilities of the CPs since the runout flow for the CPs is 7,450 gpm whereas the maximum required system runout flow per CP is 5,781 gpm.
The RFP suction pressure trip setpoint discussion relates to the basis for selection of the RFP suction pressure trip setpoint. This setpoint is based on assuring that, for normal and off-normal operation (including the trip of one condensate pump), sufficient pressure exists at the suction of the RFPs to provide more NPSH available than NPSH required by the pump to avoid potentially harmful pump suction cavitation.
The response to RAI SPLB-A-17 in Supplement No. 28 also described an automatic runback of the recirculation pumps that will be implemented for EPU. This feature is designed to ensure that upon a trip of a condensate pump or a feedwater pump, a recirculation pump runback will occur to quickly reduce reactor power and steam flow to values that allow the remaining operating pumps to support continued plant operation.
The loss of an operating CP at EPU conditions results in:
: 1. Reduced feedwater flow, which affects reactor water level. That is, with two operating CPs, following a loss of one of the three CPs, the condensate and feedwater system is incapable of supplying feedwater flow rates necessary for full power operation-even if the demineralizers are at the cleanest condition and the feedwater regulating valves are 100% open.
Upon a loss of a CP, to maintain full power, each remaining CP would have to provide about 8,100 gpm. However, given the system hydraulic characteristics with the feedwater regulating valves 100% open, the maximum flow possible per CP is calculated to be 7,537 gpm following a loss of a CP. Thus, it is not possible to maintain full power following a loss of a CP.
As indicated above, the reactor recirculation (RR) system runback feature is intended to quickly reduce reactor power and steam flow to values that allow the remaining operating pumps to support continued plant operation at a reduced power level.
: 2. The calculated RFP suction pressure following the trip of one CP at EPU is approximately 124 psig. This suction pressure is the minimum pressure predicted prior to the RR runback assuming the feedwater regulating valves are full open and maximum DP across the condensate demineralizers. The RFP low suction pressure trip setpoint is currently set to 98 psig, to avoid a RFP suction pressure trip following a CP trip. This setpoint will be retained for EPU conditions.

Attachment 8 BVY 05-072 Docket No. 50-271 Page 4 of 12
: 3. The remaining two operating CPs will continue to maintain sufficient NPSHA for the RFPs at the maximum resulting individual CP flow of 7,537 gpm. This maximum flow of 7,537 gpm would be available prior to the runback, assuming the feedwater regulating valves are full open and minimum DP across the condensate demineralizers.
The resulting individual CP flow of 7,537 gpm is 87 gpm or approximately 1% more than the published characteristic curve maximum flow of 7,450 gpm (i.e., 7,537 - 7,450 = 87 gpm). Since the increased flowrate is practically insignificant, reasonable estimates of CP capabilities, TDH and NPSHR, can be obtained by extrapolating the pump characteristic curve to the slightly higher flowrate. From this extrapolation, it is estimated that at the resulting flow rate of 7,537 gpm, the CP NPSHA is about 10 feet above the estimated pump NPSHR indicating that more than sufficient NPSHA exists to prevent CP cavitation.
Since the remaining two CPs continue to operate with more than sufficient NPSHA, the lowest calculated RFP suction pressure will be approximately 124 psig or well above the RFP suction trip setpoint of 98 psig.
RAI SPLB-A-28 EPU operation will result in a substantial reduction in the available condensate and feedwater system operating margin and plant modifications must now be credited for preventing challenges to reactor safety systems that would otherwise occur upon the loss of a RFP or a condensate pump. Because the plant response to loss of RFP and condensate pump events following EPU implementation is substantially different from the response at the current licensed power level, and the expected EPU response has not been confirmed by previous full power tests or plant transients, the NRC staff requires that the power ascension test program include sufficient testing at the 100% EPU power level to confirm that the plant will respond as expected following a) the loss of a RFP, and b) the loss of a condensate pump. Please provide a complete description of the full-power testing that will be completed in this regard for the staff's review and approval, and propose a license condition that will assure that the proposed testing will be completed as described and that the results are fully satisfactory as a prerequisite for continued operation at the EPU power level.
Response to RAI SPLB-A-28 This response is supported by the information previously provide in response to RAI SPLB-A-17, page 9 of 23, and RAI-SPLB-A-18, page 12 of 23, as Attachment 1 to Entergy letter BVY-05-046, dated April 22, 2005.
A modification to the condensate and feedwater system was installed at VYNPS to add a reactor recirculation (RR) pump runback as a trip avoidance feature to reduce the potential for a reactor low water level scram on the loss of either a feedwater or condensate pump at EPU conditions. Although not required to achieve full EPU operation, Entergy determined it is prudent to avoid this potential plant transient/trip in the very unlikely event that a condensate pump (CP) or reactor feedwater pump (RFP) trips at EPU operating conditions. A dynamic analysis of a single feedwater pump trip at EPU conditions indicates that an automatic reactor

Attachment 8 BVY 05-072 Docket No. 50-271 Page 5 of 12 recirculation runback can reduce core flow and thermal power to within the capability of the running feedwater pumps and avoid a reduction in reactor water level to the scram setpoint.
VYNPS has been analyzed to respond to a reactor trip on low reactor water level at EPU conditions and maintains adequate margin to safety for this event. The RR pump runback feature is not a safety function because the reactor is designed to scram if operating limits are exceeded; it is rather an operational reliability issue because it is preferred that the loss of a condensate or feedwater pump not result in a scram.
The runback logic is enabled when reactor power exceeds the capability of two feedwater pumps, as measured by total steam flow (approximately 112% of CLTP). A runback is initiated when fewer than three feedwater pumps and three condensate pumps are running and total feedwater flow exceeds the capacity of two feedwater pumps. The automatic runback will rapidly reduce core flow to approximately 60% of rated EPU core flow.
A transient analysis performed by GE in support of VYNPS using an NRC approved code (ODYN) shows that even in a degraded condition:
        "The results of the single feedwater pump trip evaluations show that even the cases with a very degraded response and 1-element feedwater control show the acceptability of the reactor recirculation run back. Reactor water level remains above the low reactor water scram setpoint and below the high reactor water level trip setpoint (Level 8) for all conditions."
Thorough logic testing was performed as part of the modification, including using breaker trips to initiate the runback circuitry and monitoring RR pump controls. Based on the analysis of the plant response to the pump trip and the offline testing performed during post-modification testing, it has been determined that no functional test is warranted.
The EPU configuration differs from the current configuration which uses two RFPs with one RFP available as a standby pump. During steady state EPU operation, the three RFPs will operate at lower pump capacity (5,831 gpm) with less stress on pumps and motors than two RFPs operating at CLTP (6,965 gpm). This reduction in individual RFP flow increases the available margin from normal operating flows to runout for the individual pump. Each CP will be required to provide the increased flow associated with EPU operation with the same number of pumps (i.e., three) currently being used. The CP flow margin between the EPU conditions and the runout with three RFPs and three CPs will be approximately 7% greater than the required EPU flow. Industry criteria typically recommend a 5%margin. As such, the available margin exceeds that typically required by industry.
The operation of the feedwater and condensate systems in terms of required response to initiating events does not fundamentally change at EPU. At CLTP the trip of a CP requires operator action to reduce RR flow/power level to a point supported by the remaining pumps.
When one RFP is out of service operations must also reduce flow/power rapidly to avoid the low level trip. Infact, without the modification the likelihood today is that the trip of a RFP without a standby pump in autostart will likely result in a reactor trip. The RFPs and CPs at VYNPS have

Attachment 8 BVY 05-072 Docket No. 50-271 Page 6 of 12 a long history of reliable operation. No pump trips at power have occurred in at least the last ten years of power operation.
Based on discussions with others in the nuclear industry, no other uprated BWR needed to perform an integrated test to confirm the RR runback performed as designed. Additionally, for BWRs that installed the runback feature as part of original plant design, only one BWR plant could be identified as having performed the test.
This modification was installed for both operational support and economic reasons. The economic reasons include keeping the plant on line and minimize operator response to a transient involving the loss of a single RFP or CP at EPU conditions. The operational support reasons are to avoid an unnecessary low level scram during EPU conditions due to a loss of a RFP or CP at EPU conditions.
NRC's Standard Review Plan (SRP) 14.2.1, Section III.B.1 states: "The reviewer should assess if the licensee adequately identified functions important to safety that are affected by EPU-related modifications, setpoint adjustments, and changes in plant operating parameters. In particular, the licensee should have considered the safety impact (emphasis added) of first-of-a-kind plant modifications, the introduction of new system dependencies or interactions, and changes in system response to initiating events. The review scope can be limited to those functions important to safety (emphasis added) associated with the anticipated operational occurrences described in Attachment 2 to this SRP, "Transient Testing Applicable to Extended Power Uprates." To assist in this review, Attachment 2 also includes typical transient testing acceptance criteria and functions important to safety associated with these anticipated events."
The RR runback based on a RFP or CP trip or low feedwater pump suction pressure does not meet any of the criteria per Attachment 2, "Transient Testing Applicable to Extended Power Uprates." However, additional insight is gained by review of the exception criteria provided in SRP 14.2.1 Section III.C.
Addressing SRP Section III.C:
: a. Previous operating experience:
Entergy is unaware of any VYNPS or industry EPU operating experience that supports performance of this test. The operational history of VYNPS and the very limited industry experience with RFP and CP trips at power supports that there is little benefit in injecting this transient.
: b. Introduction of new thermal hydraulic phenomena or identified system interactions:
There are no new thermal hydraulic phenomena (reducing RR flow based on a pump trip is not a new thermal hydraulic phenomena) or system interactions that may be introduced as a result of the RR runback. Feedwater flow changes due to plant events (e.g., pump trip, valve malfunction) result in pre-analyzed

Attachment 8 BVY 05-072 Docket No. 50-271 Page 7 of 12 transients (low water level reactor trip, high level turbine trip).
C. Facility conformance to limitations associated with analytical analysis methods:
The ODYN analysis cannot provide 100% confidence that a scram is avoided. It is dependent on the actual runback rate and various systems' performance at the time of the postulated event. There may be certain operating conditions where the plant is more vulnerable to scram. The'real margin to a scram will depend on the actual instrument setting and whatever instrument drift upwards may occur. This is acceptable as the impacted systems are not safety related and uncertainty in the analysis would not alter potential outcomes (no outcome different than analyzed events).
: d. Plant staff familiarization with facility operation and trial use of operation and emergency operating procedures:
Plant operators have been trained on both the automatic plant response and any required operator actions in response to these events. This includes simulator exercises that include these events. No different types of operator actions are required.
: e. Margin reduction in safety analysis results for anticipated operational occurrences:
There is no reduction in margin to safety in either installation of the modification performed or the lack of an online integrated test. The modification was installed to minimize operational transients and avoid a plant trip in response to the pump trip. Failure of the modification to initiate the runback, failure in the execution of the runback, or a plant response to the runback different than that modeled in the analysis will lead to a reactor low level trip or a high level turbine trip (analyzed events that have been experienced at VYNPS and elsewhere).
: f. Guidance contained in vendor technical reports:
No guidance is contained in vendor technical reports related to performing integrated tests related to the RR runback.
: g. Risk implications:
VYNPS is not proposing a risk informed basis for not performing certain transient tests.

Attachment 8 BVY 05-072 Docket No. 50-271 Page 8 of 12 Conclusion Entergy has instituted all of the necessary modifications and required post-modification testing at VYNPS to provide reasonable assurance that the feedwater and condensate system will remain highly reliable, and in the event of a pump trip, no new challenge to safety occurs. Therefore, Entergy believes that there is no need for a license condition, and no need to perform an integrated plant test to trip both a CP and separately a RFP at full EPU conditions to demonstrate the ability of the RR runback and feedwater low suction pressure modification to function as designed.
RAI SPLB-A-29 The licensee's response to RAI SPLB-A-20(a) in Supplement No. 28, is incomplete in that only the balance-of-plant (BOP) startup transient response criteria for the main steam isolation valve closure and generator load rejection transients were addressed. In accordance with the review criteria provided in NRC Review Standard RS-001 and draft Standard Review Plan (SRP)
Section 14.2.1, the staff's request applies to the BOP transient response for all of the startup tests that are potentially impacted by the proposed EPU. Please provide the additional information that is needed in this regard.
Response to RAI SPLB-A-29 For all startup tests that are potentially impacted by the proposed EPU, GE Document No 22A2217, Revision 1, Vermont Yankee Startup Test Specification, dated January 5, 1973 was reviewed to determine if any of the original VYNPS balance of plant (BOP) startup testing transient response criteria are affected. The criteria section of this document specifies Level 1 and Level 2 criteria which are defined as follows:
Level 1:
These values of process variables assigned in the design of the plant and equipment are included in this category. If a Level 1 criterion is not satisfied, the plant will be placed in a hold condition which is satisfactory, until a resolution is made. Tests compatible with the hold condition may be continued. Following resolution, applicable tests must be repeated to verify that the requirements of the Level 1 criterion are satisfied.
Level 2:
The limits considered in this category are associated with expectations in regard to the performance of the system. If a Level 2 criterion is not satisfied, operating and testing plans would not necessarily be altered. Investigations of the measurements and of the analytical techniques used for the predictions would be started.

Attachment 8 BVY 05-072 Docket No. 50-271 Page 9 of 12 Table SPLB-A-29-1 below lists the startup tests that are potentially impacted by the proposed EPU, and the Level 1 and Level 2 Acceptance Criteria per GE Document No 22A2217, Revision 1, Vermont Yankee Startup Test Specification, dated January 5, 1973.

==
Conclusion:==

After reviewing the list below, for all of the startup tests that are potentially impacted by the proposed EPU, none of the original VYNPS balance of plant (BOP) startup testing transient response criteria are affected.
Table SPLB-A-29-1 Test            Description              BOP Transient Response Acceptance Criteria Evaluation I    Chemical and                     Level 1:
Radiochemical o    Water quality must be known and must conform to water quality and fuel warranty specifications.
o    The activity of gaseous and liquid effluents must be know and must conform to license limitations.
o    Chemical factor defined in the Technical Specifications must be maintained within limits specified.
Level 2:
o    None 2    Radiation Measurements           Level 1:
o   The radiation doses of plant origin and occupancy times shall be controlled consistent with the guidelines of the standards for protection against radiation 10CFR20.
Level 2:
o    None

Attachment 8 BVY 05-072 Docket No. 50-271 Page 10 of 12 10 IRM Performance  Level 1:
o    None Level 2:
o    The IRM channels will be calibrated to read equal to or grater than the actual percent of reactor rated thermal power, and will overlap the SRM and APRM reading.
12 APRM Calibration Level 1:
o    The APRM channels must be calibrated to read equal to or greater than the actual core thermal power.
Level 2:
o    None 19 Core Performance Level 1:
o    Reactor power, maximum fuel surface heat flux, and minimum critical heat flux ratio (MCHFR) must satisfy the following limits:
o Maximum fuel rod surface heat flux shall not exceed 134 W/cm 2 (425,500 BTU/hr-ft 2).
o   Minimum CHF ratio shall not be less than 1.9 when evaluated at the operating power level. The basis for evaluation of MCHFR shall be 'Design Basis for Critical Heat Flux Condition in BWRs" APED-5286, Sept. 1966.
o   Normal reactor power shall be limited to 1593 MWt for the steady state conditions.
Level 2:
o    None

Attachment 8 BVY 05-072 Docket No. 50-271 Page 11 of 12 22 Pressure Regulator       Level 1:
o    The decay ratio must be less than 1.0 for each process variable that exhibits oscillatory response to pressure regulator changes.
Level 2:
o    The decay ratio is expected to be less than or equal to 0.25 for each process variable that exhibits oscillator response to pressure regulator changes when the plant is operating above the lower limit setting of the Master Flow Controller.
o    During the simulated failure of the operating pressure regulator, the backup pressure regulator shall control the transient such that the reactor does not scram.
23 Feedwater Control System Level 1:
Testing o    The decay ratio must be less than 1.0 for each process variable that exhibits oscillatory response to feedwater system changes.
Level 2:
o    The decay ratio is expected to be less than or equal to 0.25 for each process variable that exhibits oscillator response to feedwater system changes when the plant is operating above the lower limit setting of the Master Flow Controller.

Attachment 8 BVY 05-072 Docket No. 50-271 Page 12 of 12 24  Turbine Valve Surveillance    Level 1:
o   The decay ratio must be less than 1.0 for each process variable that exhibits oscillatory response to bypass valve changes.
Level 2:
o The decay ratio is expected to be less than or equal to 0.25 for each process variable that exhibits oscillator response to bypass valve changes when the plant is operating above the lower limit setting of the Master Flow Controller.
This transient is not expected to cause a scram.
25  Main Steam Isolation Valves   Level 1:
o    MSIV stroke time will be between 3 and 5 seconds. Reactor pressure shall be maintained below 1230 psig, the setpoint of the first safety valve, during the MSIV closure event.
Level 2:
o The maximum reactor pressure should be 35 psi below the first safety valve setpoint. This is a margin of safety for safety valve weeping.
100 Main Steam and Feedwater Not part of the original Startup Test Specification Piping Vibration            I

BVY 05-072 Docket No. 50-271 Attachment 10 Vermont Yankee Nuclear Power Station Proposed Technical Specification Change No. 263 - Supplement No. 30 Extended Power Uprate Response to Request for Additional Information Responses to Reactor Systems Branch RAls NON-PROPRIETARY VERSION Total number of pages in Attachment 10 (excluding this cover sheet) is 126.

Attachment 10 BVY 05-072 Docket No. 50-271 Page 1 of 126 RESPONSE TO REQUEST FOR ADDITIONAL INFORMATION REGARDING APPLICATION FOR EXTENDED POWER UPRATE LICENSE AMENDMENT VERMONT YANKEE NUCLEAR POWER STATION PREFACE This attachment provides responses to the NRC Reactor Systems Branch's sSRXB) individual requests for additional information (RAIs) in NRC's letter dated July 27, 2005. Upon receipt of the RAI, discussions were held with the NRC staff to further clarify the RAI. In certain instances the intent of certain individual RAls may have been modified based on clarifications reached during these discussions. The information provided herein is consistent with those clarifications.
The individual RAls are re-stated as provided in NRC's letter of July 27, 2005.
Reactor Systems Branch (SRXB)
Boiling Water Reactors and Nuclear Performance Section (SRXB-A)
RAI SRXB-A-7 Table 1-1 of the VYNPS Power Uprate Safety Analysis Report (PUSAR) (i.e., Attachment 4 of the application dated September 10, 2003), lists computer codes used for CPPU for transient analysis. Please clarify which code was used for the over-pressure protection analysis.
Response to RAI SRXB-A-7 The over-pressure protection analysis was performed using version 10 of the ODYN code2 ,
which is applicable to plants that use variable speed pumps for recirculation flow control.
RAI SRXB-A-8 Section 3.10.1 of the VYNPS PUSAR discusses the shutdown cooling (SDC) analysis for CPPU. However, SDC with single loop operation was not discussed in the PUSAR. Please clarify which criteria apply to SDC with single loop operation, and whether the criteria are satisfied at CPPU conditions.
Response to RAI SRXB-A-8
                                                )) The reactor shutdown cooling under abnormal conditions with either loss of the normal suction flow path or loss of one SDC loop, is considered a licensing basis, not a design basis.
1 U.S. Nuclear Regulatory Commission (Richard B. Ennis) letter to Entergy Nuclear Operations, Inc.
(Michael Kansler), "Request for Additional Information - Extended Power Uprate, Vermont Yankee Nuclear Power Station (TAC No. MC0761)," July 27, 2005 2"Qualification of the One-Dimensional Core Transient Model (ODYN) for Boiling Water Reactors,"
NEDO-241 54-A, Vols. 1 - 3,August 1, 1986, NEDC-241 54P-A Supplement 1,Volume 4, February 2000

Attachment 10 BVY 05-072 Docket No. 50-271 Page 2 of 126 There is no licensing basis or design basis requirement for single loop SDC analysis because VYNPS has no commitment to Reg. Guide 1.139, or Technical Specification requirement that specifies SDC operation using single RHR loop.
Evaluations have been performed for certain events (i.e., an Appendix R fire) that assume only the capability of one loop of RHR shutdown cooling. The SDC capability is evaluated at power uprate conditions as part of the specific event scenario in which SDC is credited, using assumptions appropriate for the specific event. Section 3.10.1 of the VYNPS PUSAR provides evaluation of the RHR SDC design basis consistent with the design and licensing basis for the VY SDC mode of operation.
RAI SRXB-A-9 Section 2.2 of the PUSAR states that a representative cycle core was used for the CPPU evaluation. Please define the VYNPS 'representative" cycle core and discuss which GE fuel type is limiting from the standpoint of fuel thermal limits.
Response to RAI SRXB-A-9 A Power Uprate Representative Equilibrium Cycle (PUREC) core design was generated at LPU/100% rated core flow conditions representing a full-core loading of GE14 fuel and an 18-month equilibrium operating cycle. In addition, refer to the response to NRC RAI SRXB-A-19.
RAI SRXB-A-10 Please provide the following additional information regarding the VYNPS LBLOCA analysis for the CPPU:
a)   Describe the VYNPS limiting single failure LBLOCA event for the current licensing basis conditions and for EPU conditions, respectively. Typically, the events are the same; but if the events are different for VYNPS, then explain the reasons. Also, describe the type of reactor core that was assumed for the EPU analysis (i.e., whether the core was assumed to be loaded with the same kind of GE fuel, or a mixed-core was assumed). If it was a mixed-core, then describe which GE fuel types used, their proportions and burnup level, etc.
b) The peak cladding temperature (PCT) changes due to CPPU are typically within 20 'F; but for the VYNPS EPU, it was determined to increase by 50 "F. Discuss the reasons behind such a comparatively large increase of the PCT, and why VYNPS is an exception in this regard.
Response to RAI SRXB-A-10 a)      At both current licensing basis and EPU conditions, the limiting large break LOCA case for VYNPS is the maximum recirculation line break with a DC power source (battery) failure. This case results in the large break with the least amount of ECCS available.

Attachment 10 BVY 05-072 Docket No. 50-271 Page 3 of 126 Consistent with the approved SAFER/GESTR-LOCA methodology, the VYNPS EPU ECCS-LOCA analysis assumed an equilibrium core loading.                   This approach is acceptable because ((
                                                                     ))
b)     The <20IF impact due to EPU is the expected average result across the BWR fleet.
However, the impact on a given plant may be larger due to plant specific characteristics.
A comparison of the VYNPS Appendix K LOCA response for the two operating conditions show almost identical responses with the vessel water level for the EPU case reflooding the bottom of the core only slightly later (i.e., about 4 seconds). This is expected for the EPU case which initializes with higher voids (resulting in less initial vessel inventory) and less subcooling (resulting in more inventory loss due to flashing from depressurization). At the time of reflooding the cladding heatup rate for VYNPS is about ((            JJ which results in the ((      1] difference in PCT. This rate is high because ((
                                                       ))
RAI SRXB-A-11 As shown in Supplement No. 4, Attachment 4 (NRC Review Standard RS-001, BWR Template Safety Evaluation (SE) as revised for VYNPS), Section 2.8.6, .Fuel Storage," draft General Design Criterion (GDC) 66 is applicable to the NRC's review of the affect of the proposed EPU on new and spent fuel storage. This GDC requires prevention of criticality in fuel storage systems by physical systems or processes, preferably utilizing geometrically safe configurations. The NRC staff did not find any discussion on criticality of new and spent fuel storage in the licensee's submittals. Please provide this information.
Response to RAI SRXB-A-11 For EPU, Vermont Yankee is not changing its licensing basis requirements for new or spent fuel storage, which are listed in Technical Specification 5.5. These include the following:
a)     The new fuel storage facility shall be such that the effective multiplication factor (Keff) of the fuel when dry is less than 0.90 and when flooded is less than 0.95.
b)     The Keff of the fuel in the spent fuel storage pool shall be less than or equal to 0.95.
c)      Spent fuel storage racks may be moved (only) in accordance with written procedures which ensure that no rack modules are moved over fuel assemblies.
d)     The number of spent fuel assemblies stored in the spent fuel pool shall not exceed 3353.
e)     The maximum core geometry infinite lattice multiplication factor of any segment of the fuel assembly stored in the spent fuel storage pool or the new fuel storage facility shall be less than or equal to 1.31 at 200C.

Attachment 10 BVY 05-072 Docket No. 50-271 Page 4 of 126 RAI SRXB-A-12 In Section 7.2.1, Step # 3, NUMARC 87-00, Revision 1, "Guidelines and Technical Bases for NUMARC Initiatives Addressing Station Blackout at Light Water Reactors," it is indicated that the minimum permissible usable gallons of water in the condensate storage tank (CST) should be recorded in the plant Technical Specifications (TSs). In Supplement No. 25, Attachment 2, Section 2.1, the licensee stated that in order to ensure that at least 100,000 gallons of usable CST inventory is available during an SBO, the minimum administrative limit for CST level identified in procedure OP 0150, "Conduct of Operations and Operator rounds," will be increased. The current CST minimum volume in TS 3.5.E.1.b is 75,000 gallons. Please justify why the TSs do not need to be revised, consistent with the recommendations in NUMARC 87-00.
Response to RAI SRXB-A-12 The purpose of Section 7.2.1 of NUMARC 87-00, Rev. 1 is to ensure that adequate condensate inventory is available for decay heat removal during an SBO event for the required coping duration. Procedural Step 3 of Section 7.2.1 provides one measure for assessing the adequacy of condensate inventory requirements and is not a recommendation to revise the TSs.
VYNPS relies on administrative controls (i.e., plant procedures) to meet the SBO condensate inventory requirements. Step 5 of Section 7.2.1 provides the criteria for establishing additional water sources in addition to the TS limit value of the condensate storage tank (CST). VYNPS's credit of the administratively controlled additional CST volume meets the four criteria specified in Step 5 of NUMARC 87-00, Rev. 1 and therefore is an acceptable means to demonstrate that acceptable water is available to support the coping period.
10CFR50.36(c)(2)(ii) establishes criteria for Technical Specification limiting conditions for operation. The CST inventory required for copying with an SBO event does not satisfy these criteria because it does not involve:
: 1. Installed instrumentation that is used to detect, and indicate in the control room, a significant abnormal degradation of the reactor coolant pressure boundary.
: 2. A process variable, design feature, or operating restriction that is an initial condition of a design basis accident or transient analysis that either assumes the failure of or presents a challenge to the integrity of a fission product barrier.
: 3. A structure, system, or component that is part of the primary success path and which functions or actuates to mitigate a design basis accident or transient that either assumes the failure of or presents a challenge to the integrity of a fission product barrier. An SBO event is not a design basis accident or transient event that either assumes the failure of or presents a challenge to the integrity of a fission product barrier.
: 4. A structure, system, or component which operating experience or probabilistic risk assessment has shown to be significant to public health and safety. SBO risk is assessed to be low.

Attachment 10 BVY 05-072 Docket No. 50-271 Page 5 of 126 Furthermore, in accordance with 10CFR50.36(c)(2)(iii), a licensee is not required to-propose to modify technical specifications that are included in any license issued before August 18, 1995, to satisfy the criteria in paragraph (c)(2)(ii) of this section.
RAI SRXB-A-13 As discussed in PUSAR Section 2.4, "Stability," VYNPS currently operates under the Option l-D solution. Please provide a clarification for the following areas:
a)   The current flow-biased average power range monitor (APRM) scram provides automatic detection and suppression of core wide instability. Provide the technical basis that supports the conclusion that regional mode reactor instability is not probable under EPU conditions.
b)   Describe any alternative method to provide automatic detection and suppression of any mode of instability other than through the current flow-biased APRM scram.
c)   Describe how the dominance of the core-wide mode oscillations is maintained under the EPU conditions. Specifically, describe how the effects on axial and power distributions (which change for EPU core loadings) have been taken into account in the new calculations to ensure the dominance of the core-wide mode. Are there any negative effects on stability of the EPU core loadings?
Response to RAI SRXB-A-13 a)      As stated in NEDO-31960-A 3 , the Option l-D solution is applicable to plants that have relatively tight fuel inlet orificing and relatively small diameter cores. For plants with tight inlet orificing, the probability of regional oscillations is very low and the expected mode of oscillation is core-wide. In addition, plants with small diameter cores are also less likely to experience regional oscillations because of the strong preference of the fundamental mode of the neutronics. Even though the implementation of EPU at VYNPS does not change the fuel inlet orificing or the core diameter, the EPU core design might drive a higher core decay ratio due to higher power at extended operation. The APRM flow-biased flux scram line has been redesigned with three slopes for ARTS/MELLLA operation and rescaled for EPU operation such that the SLMCPR is protected against thermal-hydraulic instability events. ((
                                             )) Calculations will continue to be performed for each VYNPS reload cycle to demonstrate that the probability of regional oscillations is very low.
b)      Response: There is no other method to provide automatic detection and suppression of any mode of instability for Option l-D plants other than through the current flow-biased APRM scram.
3 NEDO-31960-A,      -BWR Owners' Group Long-Term Stability Solutions Licensing Methodology,"
Licensing Topical Report, November 1995

Attachment 10 BVY 05-072 Docket No. 50-271 Page 6 of 126 c)       The dominance of the core-wide mode oscillation can be demonstrated by calculating the core and channel decay ratio at the most limiting power/flow state point (i.e., the intersection of Natural Circulation Line and MELLLA boundary). If the calculated channel decay ratio is = 0.56, the dominance of the core-wide mode of oscillation is demonstrated.
A limiting axial power shape with a 2.00 magnitude axial power shape peaked at node 3 is applied to the hottest channel and this gives a relatively conservative axial power shape in the hot channel decay ratio evaluation.
Calculations will continue to be performed for each VYNPS reload cycle to demonstrate that the hot channel decay ratio criterion is met and thus the probability of regional mode oscillations is very low.
RAI SRXB-A-14 Provide the technical basis that supports a conclusion that the hot bundle oscillation magnitude portion of the detect-and-suppress calculation is not dependent upon the core and fuel design.
Response to RAI SRXB-A-14 As stated in NEDO-32465-A 4 , the hot channel oscillation magnitude (HCOM) is dependent on plant-specific factors. Some of the parameters that affect the hot bundle oscillation magnitude are: core size, LPRM assignments, trip setpoints, growth rate, harmonic power distributions (contours), LPRM failures, trip overshoot, and oscillation frequency. The only parameter that could potentially be affected by the core and fuel design would be the contours. However, the EPU power distributions are not significantly different from the pre-EPU power distributions to have a significant effect on the harmonic power distributions and hence, the hot bundle oscillation magnitude. Please note that the HCOM values have been revised for Cycle 24 EPU operation to demonstrate that the APRM flow-biased flux scram line provides adequate SLMCPR protection against thermal-hydraulic instability events.
RAI SRXB-A-15 The hot channel decay ratio provided for EPU is very close to acceptable criteria limits. In addition, the core-wide decay ratio is not provided. Have the proposed EPU core loadings degraded the stability performance significantly? Provide a table of hot channel and core-wide decay ratios at the most limiting state point for the last three cycles and the proposed EPU condition. The purpose is to evaluate the impact of the EPU on relative stability of the plant, and the applicability of Option l-D to VYNPS under these new conditions.
4 NEDO-32465-A, 'Reactor Stability Detect and Suppress Solutions Licensing Basis Methodology for Reload Applications,' Licensing Topical Report, August 1996

Attachment 10 BVY 05-072 Docket No. 50-271 Page 7 of 126 Response to RAI SRXB-A-15 The limiting hot channel and procedure core decay ratios at the most limiting state point for the last three cycles including the proposed EPU condition are provided in Table SRXB-A-15-1 below:
Table SRXB-A-15-1 Cycle           Limiting Hot Channel     Procedure Core Decay Ratio              Decay Ratio 22              ((
24 (EPU)                                        ]3 Note that the Cycle 22 decay ratios were computed with FABLE/BYPSS methodology and therefore are not directly comparable to the Cycle 23 and Cycle 24 results, which were computed with ODYSY methodology. (The limiting hot channel and procedure core decay ratios at the most limiting state point were not computed for Cycle 21.) By comparing the Cycle 23 and Cycle 24 results, it can be seen that the proposed EPU core loadings do not significantly impact the stability performance of the plant. ((
                                                               )) Therefore, the dominant mode of oscillations is core-wide and Option l-D is applicable to VYNPS under these new conditions.
RAI SRXB-A-16 It appears that the APRM flow-biased scram setpoint will be maintained at the same absolute levels (in terms of megawatts) for EPU as for CLTP. Please address the following:
a)   Because the distance (in terms of megawatts (MWs)) between the most limiting power/flow operating point and the scram setpoint represents the oscillation amplitude required for scram, has this distance (i.e., the maximum oscillation amplitude) changed for EPU?
Provide a graphical power/flow map representation of the new and old operating domains and the VYNPS scram setpoints, including the exclusion region. Note that the most limiting condition in terms of the oscillation amplitude is not necessarily the most unstable point, but the one that results in the largest amplitude.
b)   If the above distance (i.e., the oscillation amplitude required for scram) has changed, is the CLTP scram setpoint still conservative for the EPU?
c)   Has the resolution of the recent DIVOM (delta critical power ratio (CPR) over initial minimum CPR versus oscillation magnitude) 10 CFR Part 21 notification had any effect on VYNPS implementation of Option l-D? What DIVOM correlation is used to justify the EPU scram line? Is it a plant-specific or generic correlation? Please provide details.

Attachment 10 BVY 05-072 Docket No. 50-271 Page 8 of 126 Response to RAI SRXB-A-16 a)   The rated rodline has been redefined for EPU and the APRM scram line has been rescaled for EPU. Therefore, as can be seen in Figures SRXB-A-16-1 and SRXB-A      2, the difference in power at natural circulation between the rated rodline and the APRM scram setpoint decreases from 15.3% of rated power for pre-EPU to 4.6% of rated power for EPU. This decrease in the power difference occurs for core flows higher than natural circulation as well. The corresponding hot channel oscillation magnitude (HCOM) is reduced for EPU. This is offset somewhat by a larger value of DIVOM for EPU. The net effect is an improvement in margin to the SLMCPR.
b)   The APRM scram setpoint has been rescaled for EPU and the scram setpoint is conservative for EPU. Calculations will continue to be performed for each VYNPS reload cycle to demonstrate that the SLMCPR is protected against thermal-hydraulic instability events.
c)
                                                                  )) This conservative DIVOM slope addresses the recent DIVOM 10 CFR Part 21 issue. Consistent with the resolution of the Part 21 issue, for future cycles, VYNPS plans to use a plant- and cycle-specific core-wide mode DIVOM curve in accordance with the BWROG Plant-Specific Core-Wide Mode DIVOM Procedure Guideline (GE-NE-0000-0031-6498-RO, June 2, 2005).

Attachment 10 BVY 05-072 Docket No. 50-271 Page 9 of 126 Vermont Yankee EPU, Power/Flow Map Core Flow (MlIb/lir) 0  5  10  15    20       25       30     35    40    45     50      55 2200 2000 1800 1600   c_
1400  r "I                                                                                     I-I Pi                                                                                1200  W 0
Ca 1000   ;
I..                                                                                    E a..
800    -
P 600 400 200 0 10 20  30   40      50       60     70    80    90    100     110     120 Core Flow (%)
Figure SRXB-A-16-1

Attachment 10 BVY 05-072 Docket No. 50-271 Page 10 of 126 Vermont Yankee Pre-EPU, Power/Flow Map Core Flow (MlIb/hr) 0                      10            20                  30                        40                       50 120    _
MELLLA Upper Boundary Line                                                              1800 110 100-                                         108%CLTPRodLine                                                            1593 IWtLt 1600 90  -
1400 APRM Scram 8O
                                                                                                                                       . 1200 70 -                                                                                                                                      r.
L.                                                                          / /, . *I'01/OO CLTP Rod Line                                        L.*
(U 1000 60  -                                                                                                                                      0 15.3% Power Difference   -oo                                                                                                       5-800   co E 50 c                                                                                                                                             .0 Minimum Pump Speed                                                                           E-
-    40 -
t  600 Natural Circulation                                        Minimum Power Line 30  -
                                                                                                                                        . 400 20  -                                 --                    --------        1 1,-       - - -- - - - - - -     - -_
I0O%CLTP        -1593 MWt                         200 10  -                                                                              100/ /OLTP      -1593WIt I l100%
Care low - 4.0 Mlbbhr 0                                                                                               I          I         I         -O 0        10           20       30  40      50        60            70          80           90         100      110          120 Core Flow (%)
Figure SRXB-A-16-2

Attachment 10 BVY 05-072 Docket No. 50-271 Page 11 of 126 RAI SRXB-A-17 In Supplement 4, Attachment 5, Matrix 8, page 13, note for SE Section 2.8.5.4.1, there is an explanation for uncontrolled control rod withdrawal from a subcritical or low power startup
'condition. In this explanatory section, this event is considered as an accident and a fuel enthalpy of 170 calories/gram is given as the acceptance criterion. However, in SRP Section 15.4.1, this event is considered as a transient, not as an accident, and hence specified acceptable fuel design limit criteria is applied. Why is this event considered as an accident rather than a transient?
Response to RAI SRXB-A-17 This event is indeed considered to be a transient consistent with the SRP. For these low power conditions, the acceptance criterion for transients is the cladding failure threshold of 170 cal/gram, which is much lower than the acceptance criterion of 280 cal/gram for accidents such as a Control Rod Drop event. The approved GESTAR fuel Licensing Topical Report (LTR)
NEDE-24011-P-A-14-US refers to NEDO-10527 (GESTAR Reference S-12), which documents both the criteria and methods used in the RWE transient. The analysis of the RWE transient is documented in NEDO-23842.
RAI SRXB-A-18 Review Standard RS-001, BWR Template SE for Sections 2.8.5.1, 2.8.5.2.1, 2.8.5.2.2, 2.8.5.2.3, 2.8.5.3.1, 2.8.5.4.3, 2.8.5.5 and 2.8.5.6.1, guides the NRC staff to reach a conclusion regarding reactor coolant pressure boundary (RCPB) pressure limits not being exceeded.
However, the revised template reflecting the VYNPS licensing basis (provided in Supplement Nos. 4 and 8) does not include any acceptance criteria in the 'Regulatory Evaluation" portion of each of these SE sections related to the RCPB. Please confirm that draft GDC-9, "Reactor Coolant Pressure Boundary," is applicable to these sections and provide a markup of the SE template accordingly.
Response to RAI SRXB-A-18 Draft GDC-9, "Reactor Coolant Pressure Boundary," is applicable to Review Standard RS-001, BWR Template SE for Sections 2.8.5.1, 2.8.5.2.1, 2.8.5.2.2, 2.8.5.2.3, 2.8.5.3.1, 2.8.5.4.3, 2.8.5.5 and 2.8.5.6.1. See the SE template markups for these sections as Exhibit SRXB-A-18-1 in Attachment 11.
RAI SRXB-A-19 The following question relates to the review for SE template Section 2.8.5.4.3, 'Startup of a Recirculation Loop at an Incorrect Temperature and Flow Controller Malfunction Causing an Increase in Core Flow Rate."
VYNPS UFSAR Section 14.5.6 states that: "[flow dependent operating limits, MCPR(F)
[minimum critical power ratio (MCPR) flow-dependent limit], LHGRFAC(F) [linear heat

Attachment 10 BVY 05-072 Docket No. 50-271 Page 12 of 126 generation rate (LHGR) flow-dependent multiplication factor] and MAPFAC(F) [maximum average planar linear heat generation rate (MAPLHGR) flow-dependent multiplication factor] are developed to ensure that core thermal limits are not violated for the limiting flow increase transients." This UFSAR section also states that "[t]hese flow-dependent limits are generic ARTS [APRM and rod block monitor TS] program limits and are derived from a conservative treatment of a two recirculation pump slow flow runout event. The validity of the flow-dependent limits for the core flow increase transients was reconfirmed for the GE14 fuel introduction."
Confirm that validity of the flow-dependent limits were verified for the EPU operating conditions.
Response to RAI SRXB-A-19 It has been shown that the operating limit MCPR is not significantly affected and that the power and flow dependent limits are unaffected by CPPU 5 . The licensed core flow and the conservative control rod line are unchanged with EPU. Consistent with the CPPU methodology, the rated and off-rated MCPR operating limits are established for each fuel cycle as part of the reload analysis.
RAI SRXB-A-20 With respect to PUSAR Section 6.5, "Standby Liquid Control System":
a)    The results of the licensee's anticipated transients without scram (ATWS) analyses at EPU conditions determined that the calculated peak vessel bottom pressure is 1490 psig as shown in PUSAR Table 9-5. However, the standby liquid control system (SLCS) pump discharge pressure value proposed for the surveillance test is only 1325 psig (reference proposed revision to Surveillance Requirement (SR) 4.4.A.1). Clarify why this test pressure is acceptable.
b)    PUSAR Section 6.5 states that because of the increase in SLCS pump discharge pressure under EPU conditions, the surveillance test pressure in SR 4.4.A.1 will be increased from 1320 psig to 1325 psig. What is the SLCS discharge relief valve setpoint under EPU conditions? Taking relief valve setpoint tolerance into consideration, how much margin is there to prevent the relief valve from lifting?
Response to RAI SRXB-A-20 a)       The peak pressure referenced above (1490 psig at the vessel bottom) occurs very early in the transient event. Depending on the event, the first peak generally occurs within the first 10 to 50 seconds (see Figure SRXB-A-20-1). The SLCS pump discharge test pressure of 1325 psig is based on the peak reactor pressure that occurs during SLCS operation. This pressure is 1292 psia (1277 psig), occurring during the PRFO BOC event.
b)       The minimum SLCS pump relief valve nominal setpoint for EPU is 1400 psig. Based on the 1325 psig discharge test pressure, there is a minimum of 75-psi margin. This margin provides allowance for SLCS pump relief valve setpoint drift and for SLCS pump 5"Constant    Pressure Power Uprate", NEDC-33004P-A, Revision 4, July 2003

Attachment 10 BVY 05-072 Docket No. 50-271 Page 13 of 126 pressure pulsations. The GE recommendation for relief setpoint margin is 75 psi.
Oh      wOD    4WD                    OW              D                    2Ad "JD     woo
                                                ~Tirm
("C) woo           "MD     "MO Figure SRXB-A-20-1 Vermont Yankee Long-Term PRFO Transient Response at LPU and BOC RAI SRXB-A-21 With respect to PUSAR Section 9.1, "Anticipated Operational Occurrences," identify the staff approved evaluation model used for the plant-specific loss of feedwater flow event analysis.
Response to RAI SRXB-A-21 The SAFER code was used in the analysis of long-term Loss-of-Feedwater events.
RAI SRXB-A-22 With respect to PUSAR Section 9.3.1, "Anticipated Transients Without Scram:"
a)  Identify the staff-approved evaluation model used for the plant-specific ATWS analysis.
b)  Confirm that operator actions specified in the VYNPS emergency operating procedures are consistent with the generic Emergency Procedure Guidelines/Severe Accident Guidelines insofar as they apply to the operator actions for ATWS. Specify the time delay used in the ATWS analysis for starting of the SLCS pumps.
Response to RAI SRXB-A-22 a)     The ATWS reactor transient analysis was performed using version 10 of the ODYN code (Ref. 1), which is applicable to plants that use variable speed pumps for recirculation

Attachment 10 BVY 05-072 Docket No. 50-271 Page 14 of 126 flow control. The GE computer model STEMP04 was used for the suppression pool heatup calculation. The analytical models of STEMP have been accepted by the NRC in previous applications (Ref. 2) and other ATWS analyses.
b) The VYNPS Emergency Operating Procedures (EOPs) are consistent with Revision 2 of the Boiling Water Reactor Owners' Group (BWROG) Emergency Procedure and Severe Accident Guidelines (EPGs/SAGs), insofar as they apply to the operator actions for ATWS. With respect to the EPGs, plant specific setpoints, limits, equipment, and operating characteristics are substituted, as necessary, to ensure fidelity to the plant.
As discussed in the VYNPS PUSAR section 9.3, plant-specific analyses were performed to ensure the ATWS acceptance criteria of peak vessel bottom pressure less than 1500 psig (the ASME Service Level C limit), peak suppression pool temperature less than 281 degrees F (the wetwell shell design temperature), and peak containment pressure less than 62 psig (110% of drywell design pressure), were met. The limiting events with respect to these criteria are the Pressure Regulator Failure - Open (PRFO) and Main Steam Isolation Valve Closure (MSIVC) events. Each event was analyzed at beginning-of-cycle and end-of-cycle conditions.
As noted in the NRC SER to the Constant Pressure Power Uprate (CPPU) Licensing Topical Report (CLTR), NEDC-33004P-A, boron injection from SLC is assumed to start at the later of either (1) reaching the boron injection initiation temperature (BIIT) or (2) two minutes after the ATWS recirculation pump trip on either low reactor level or high reactor pressure. In the ATWS analyses of VYNPS at EPU conditions, the SLC initiation is assumed to occur two minutes after the recirculation pump trip on high reactor pressure.
In addition to boron injection, a number of other operator actions (consistent with the EPGs) are assumed in the VYNPS EPU ATWS analyses. These operator actions are assumed to occur at the same time or later than the timing assumed for the pre-uprate ATWS analyses, consistent with the NRC SER for NEDC-33004P-A.
The ATWS analyses methodology assumes operator action to reduce feedwater flow to the reactor in order to decrease reactor water level. This action occurs at the later of either reaching the BIIT or 90 seconds after the MSIV closure. In the ATWS analyses of VYNPS at EPU, this event was assumed to be initiated by operator action at the BIIT.
Finally, the ATWS analytical methodology assumes operator action to initiate torus cooling. For VYNPS the time at which operators initiate torus cooling was increased from the 10 minutes assumed in the pre-uprate ATWS analysis to 15 minutes at EPU conditions. This assumption, while increasing margin for operator action, is analytically more conservative because it allows additional torus water heat-up prior to initiating cooling.
In summary, operator actions assumed in the VYNPS EPU ATWS analyses are consistent with the operator actions in the EPGs and documented in the NRC SER for the CLTR.

Attachment 10 BVY 05-072 Docket No. 50-271 Page 15 of 126 RAI SRXB-A-23 Supplement No. 24, Attachment 3, Table 6-4, 'Metric Summary for VYNPS (120%)" presents the predicted maximum bundle powers and bundle power-to-flow ratios with exposure for the projected uprated conditions. In support of the staffs review of the LOCA analyses, please provide the following information specific to VYNPS:
a)   For the peak power fuel assemblies, provide the limiting axial power distributions and radial peaking factors. For different exposures, select bundles with limiting axial power peaking operating with bottom peaked, double-hump or mid-peaked, and top peaked axial power distributions. Please assure that the axial power distribution corresponding to the exposure with the highest hot bundle exit void fraction is also provided.
b)   Include in the selected bundles, the power distribution and peaking corresponding to the maximum powered bundle selected for the cycle state point of 13.184 gigawatt days per standard ton. Table 6-4 also shows that the bundle is operating at 7.51 MW. Please provide the corresponding predicted bundle operating conditions, including axial power distribution, void fraction distribution and bundle nodal exposure.
c)   Please also include the bundle inlet mass flow rate and inlet temperature.
Response to RAI SRXB-A-23 The requested data are provided in the following tables at selected points during the cycle as well as 13.184 GWd/ST. A range of bottom-peaked, mid-peaked, and top-peaked axial shapes are included.

Attachment 10 BVY 05-072 Docket No. 50-271 Page 16 of 126

Attachment 10 BVY 05-072 Docket No. 50-271 Page 17 of 126

Attachment 10 BVY 05-072 Docket No. 50-271 Page 18 of 126 I r T

Attachment 10 BVY 05-072 Docket No. 50-271 Page 19 of 126

Attachment 10 BVY 05-072 Docket No. 50-271 Page 20 of 126 RAI SRXB-A-24 Supplement No. 24, Attachment 3, Figure 6-6, presents the peak LHGR ((
                                                                                    )) Are bundles (pins) setting the peak limit for non-GE14 fuel? What uncertainties are applied to the peak LHGR to account for the calculational uncertainties?
Response to RAI SRXB-A-24 The peak LHGR values for Cycle 18 of Plant A are indeed close and sometimes ((
                   )) These peak LHGR values came from a simulation of Cycle 18 operation and are a result of the over-calculation of peak LHGR by the predictive off line nuclear methods.
The limits in question are for GE14 fuel. In actual operation, the peak LHGR never exceeded the limit value (see response to RAI-5 of MFN 05-029 TAC NO. MC5780). The uncertainties applied to the peak LHGR value are discussed in the alternate approach and the response to RAI SRXB-A-41.
RAI SRXB-A-25 Explain the reason for the increase in the ((
                                 )) The RAI response proposed void fraction weighting
                      )) Justify why the nodal uncertainties for ((
                                                                       )) in order to establish the uncertainties that should be applied to VYNPS bundle powers and thermal limits.
Response to RAI SRXB-A-25 It is assumed that in the first part of the question, by 'nodal uncertainties," is meant pin power peaking uncertainties. The data referred to are not used to establish the power allocation factor uncertainty, but were used to confirm the pin power uncertainty obtained from Monte Carlo comparisons, manufacturing uncertainties, and channel bow uncertainties. The overall derived uncertainty is ((
               )) The model differences have been evaluated and are correctly accounted for in NEDE-32601-P-A. The influence of pin power peaking on the R-factor uncertainty is discussed further in the response to RAI SRXB-A-41. The second part of the question refers to application of nodal uncertainties to VYNPS bundle power uncertainties. The process computer application of the PANAC core simulator model adapts the solution such that the axial shape is consistent with the measured TIP distribution. The radial component of the error is used to evaluate the uncertainty in the process computer bundle power. The practice of using the average RMS over a number of plants was approved by the staff in NEDE-32601 and NEDE-32694. This average is a best estimate of the uncertainty in the bundle power. In most of the US applications, the

Attachment 10 BVY 05-072 Docket No. 50-271 Page 21 of 126 bundle RMS is ((
                  )) assumed in the SLMCPR evaluation. This procedure, along with periodic updates to assure the accuracy is not changing, is adequate to ensure the plant operates within limits.
RAI SRXB-A-26
               )) Demonstrate that for the 20% uprate condition for the entire operating domain, VYNPS would not operate with core power/flow ratio greater than ((.
Response to RAI SRXB-A-26 The uprated power for VYNPS is 1912 Mwt, and the 100% flow value is 48 Mlbm/hr. Hence at the 100% power/flow point, the power/flow metric has a value of 39.8. The vast majority of operating history will be accumulated at this point.
With regard to the nodal and axial uncertainties, they are relevant only for the predictive capability of the 3-D Simulator. In monitoring mode, the solution is adapted to the measured TIP and LPRM readings (see response to RAI SRXB-A-28). The monitor nodal uncertainty depends on the radial uncertainty, (See NEDE-32694, response to question 11.5) which is shown to ((                                                           )). The nodal uncertainty in the monitoring mode is [
RAI SRXB-A-27 State what criteria are used to establish that the axial and nodal uncertainties are acceptable and do not reflect degradation of the neutronic methods predictions of the nodal and axial power distribution and peaking.
Response to RAI SRXB-A-27 Axial power distributions have an impact on both the expected thermal margins for a particular nuclear design, as well as the accuracy of the cycle length prediction. The alternate approach discusses the relationship between methods accuracy and design margin. If the methods errors fall outside the design margins, the design will not meet its requirements. Also, if the methods uncertainty of the process computer model is significantly larger than those assumed in the SLMCPR and LHGR limits, the safety limits have to be revised.
Inaccuracy in axial power distribution may also indicate a fundamental problem in the fuel or poison depletion model. Most of the time, these effects show up when a significant change is

Attachment 10 BVY 05-072 Docket No. 50-271 Page 22 of 126 made either in the power density, discharge exposure, or cycle length for a particular plant.
Core tracking calculations are performed in the predictive mode in order to continually assess the accuracy of the nuclear methods and to verify the design margin used in the design of future operating cycles. Further, the RMS difference of all predictive to measured TIP responses results in an average difference of less than ((          )) for the reference BWRs, indicating that axial power distributions are also predicted adequately. The radial RMS is checked to see if the average over a number of plants exceeds the ((                 )) assumed in the SLMCPR analysis.
Criteria for the nodal RMS are not reflected in any licensing analysis, but generally any nodal RMS values over ((        )) observed consistently require further explanation and review of the nuclear methods accuracy.
RAI SRXB-A-28 State where the axial and nodal uncertainties are accounted for in the thermal limits calculations and safety limit analyses.
Response to RAI SRXB-A-28 The relationship between thermal limits, safety limits and methods accuracy is discussed in the alternate approach section. A number of questions have related to the impact of axial or nodal uncertainties on thermal limits. In order to clarify all of the answers, it is important to emphasize the dual role of the nuclear methods. The 3-D simulator, PANACEA, is used in both the predictive and monitoring mode.
In the monitoring mode, the simulator is part of the plant process computer. The plant process computer accepts core operating conditions, and in-core instrument reading from TIPs and local power range monitors (LPRM) and converts them into a three dimensional map of linear heat generation rates and critical power ratios. The reactor operator relies on the process computer output to ensure that the LHGR and Safety Limits are not violated during operation. The uncertainties input to the SLMCPR and limiting LHGR are those applied in the monitoring mode.
In order to increase the accuracy of the process computer, GE developed an adaptive mode for the 3-D simulator. The adaptive mode for the original model (PANAC10) is described in NEDE-32694-PA and NEDC-32773 for the current model (PANAC11). In the adaptive mode, adjustments are made to the local diffusion parameters such that the axial distribution of TIP and LPRM readings calculated by the simulator are identical to the measured values. Hence in the monitoring mode, the axial RMS difference is zero, and the nodal TIP RMS is equal to the radial TIP RMS. The SLMCPR involves uncertainties in axially integrated quantities and is a function of the radial TIP RMS. The LHGR uncertainty is related to the nodal RMS, which is the same as the radial RMS. The LHGR uncertainty also includes the uncertainty in the TIP and LPRM signals.
The predictive mode is used in fuel and operating strategy design and as input to transient analysis. In the predictive mode there are no in-core instruments to normalize to, so there may be differences between the real and calculated axial power distributions. In the predictive mode, differences between the projected operating results and the actual operating history is accounted for by using design margin. In operation, the ratio between operating MCPR or

Attachment 10 BVY 05-072 Docket No. 50-271 Page 23 of 126 LHGR must never exceed 1.0. The designer is constrained to design the core such that there is a given amount of margin between the expected limiting value and the safety limit. Core tracking calculations, such as those presented in the answers to RAI-5, RAI-25, and RAI-27 of MFN 05-029 TAC NO. MC5780, are performed in the predictive mode in order to continually assess the accuracy of the nuclear methods and to verify the design margin used in the design of future operating cycles. The radial RMS is checked to see if the average over a number of plants exceeds the ((        )) assumed in the SLMCPR analysis. Criteria for the nodal RMS are not reflected in any licensing analysis, but generally any nodal RMS values over ((              ))
observed consistently for several cycles require further explanation and review of the nuclear methods accuracy.
RAI SRXB-A-29 State whether or not the axial and nodal uncertainties are accounted for in the initial steady-state conditions calculations for the transients and accident conditions (ECCS-LOCA). For example, the axial power peaking and distribution affect the response to LOCA and transients.
Therefore any under-predictions in the axial nodal powers could change the plants response. If nodal and axial uncertainties are not applied, justify how potential under-predictions in the axial and nodal powers are accounted for.
Response to RAI SRXB-A-29 The conservatism of the inputs to transient evaluations is assured by assuming an operating strategy yielding the most severe transient impact. The primary mechanism for shutting down a transient event is control blade scram. Control blade scram effectiveness is minimized when the power distribution is peaked at the top of the core. GE design procedures require a bounding analysis, which minimizes the effectiveness of the scram function for input to transient analyses. These bounding analyses result in greater penalties to the transient behavior than any caused by misprediction of axial power by the nuclear methods. This bounding analysis is further described in the alternate approach section and in the response to RAI SRXB-A-48.
RAI SRXB-A-30 Explain how the axial and nodal uncertainties are applied to the nodal exposures, the operating MAPLHGR (MAPRAT), and the operating LHGR (PKLHGR), to account for the nodal inaccuracies of the steady-state neutronic method and code systems.
Response to RAI SRXB-A-30 The alternate approach section contains a discussion of the relationship of nuclear and other uncertainties on the determination of LHGR and MAPLHGR limits. The LHGR limits include an allowance for monitoring uncertainties. In addition the fuel rod in question is assumed to operate on limits for its entire history, thereby maximizing the generation of fission product gasses and susceptibility to transient events. The peak clad temperature (MAPLHGR)

Attachment 10 BVY 05-072 Docket No. 50-271 Page 24 of 126 evaluation conservatively ((                                                     )). For VYNPS, the axial power distribution is assumed to be a cosine shape. (See reply to RAI SRXB-A-23)
RAI SRXB-A-31
((
                                                                        )) Comparing against a fixed limit value assumes that the selected maximum powered bundles are limited to low exposures bundles. State whether the ratio of the peak LHGR to the LHGR limit monitored in the core simulator (e.g. 3D MONICORE) includes the decrease of the LHGR limit with exposure.
Response to RAI SRXB-A-31 The metrics tabulated in Table 6-2 and Figure 6-6 of the original response to RAI SRXB-A-6 are meant to illustrate the highest LHGR in the experience base and were not intended to be compared against the limits curve. For VYNPS, the thermal-mechanical limits are incorporated into the monitoring system directly as a function of rod type (product line, local Gd concentration, and maximum Gd concentration anywhere in the rod) and local (pin) exposure.
Therefore, the decrease of the LHGR limit with exposure is accounted for.
RAI SRXB-A-32 State if any uncertainty is applied to the LHGR limit in the core monitoring system and if the uncertainty value is increased with exposure as the limit decreases. This is important since the peak reactivity and nodal powers increase with exposure, when the Gadolinium (Gd) bums out
(>8 GWD/ST) and the thermal-mechanical limit decreases with exposure at approximately 15 GWD/ST.
Response to RAI SRXB-A-32 The treatment of LHGR uncertainties is discussed in the alternate approach section and the response to RAI SRXB-A-31.
RAI SRXB-A-33 Describe how the core monitoring system and offline calculations calculate the peak nodal LHGR and the corresponding pin-wise peak LHGR, MAPLHGR and the accumulated exposure.

Attachment 10 BVY 05-072 Docket No. 50-271 Page 25 of 126 Response to RAI SRXB-A-33 The three dimensional power distribution, adapted to the TIP and LPRM response, is determined by the 3-D simulator. The method for the PANAC11 system used at VYNPS is similar to the one described in NEDC-32694P-A and is described in specific detail in NEDC-32773P. The nodal power, Pnode, represents the average power in a roughly 6 inch cube at a particular place in the core. For a given fuel type, a set of infinite lattice pin peaking factors are stored as a function of exposure, history water density and water density. The infinite lattice peaking factors are then corrected for fast and thermal flux gradients across the node in question. Fluxes and powers from adjacent nodes are used to construct the correction factors.
Pin powers are then established for each pin segment in each node in the core. Similarly, a pin exposure distribution is established. Limiting pin power values are established for each pin in the node, depending on the type of fuel rod and the local pin exposure. The maximum calculated-to-limiting power ratio is established for the node in question and also for the entire core. The MAPLHGR is related to the average pin power in the node. The average pin power can also be evaluated and compared to a limit usually constant with exposure but is a function of bundle mechanical configuration.
RAI SRXB-A-34 Since no gamma scans or isotopic inventory measurements are available for the current fuel designs (GE14) as operated, justify why it is acceptable to base the assessment of the predictions of the nodal powers and the pin powers on code-to-code comparisons.
Response to RAI SRXB-A-34 The process computer nodal power can be thought of as the product of three quantities. 1) The average radial four bundle power associated with a given TIP string, 2) an axial distribution obtained from the TIP response, and 3) a bundle power allocation factor giving the power distribution for the four bundles surrounding the TIP string. Factor 1) has been extensively qualified for recent applications ((                                                                 ))
Factor 2) comes directly from plant instrumentation. Qualification of factor 3), the power allocation factor has been performed to this point with the results of bundle gamma scan data.
GE recognizes that the gamma scans were carried out on older designs and applications, and is currently acting to obtain additional gamma scan data. The new gamma scan data will form the basis for a revised value for the power allocation uncertainty. Until these new data are obtained it is proposed that the current standard deviation of ((          )) be increased to ((
                       )) For additional information on the proposed uncertainties for the SLMCPR, refer to the alternate approach section.
RAI SRXB-A-35 Provide VYNPS-specific core follow data for operation at the current licensed thermal power.
Include axial individual TIP measured versus calculated data for limiting conditions in terms of

Attachment 10 BVY 05-072 Docket No. 50-271 Page 26 of 126 four-bundle nodal powers. Provide the individual TIP data for different power peaking (top-peaked, bottom-peaked and mid-peaked or double-hump). Include the corresponding four bundle predicted conditions (e.g., void fractions, nodal exposures).
Response to RAI SRXB-A-35 It is requested that the measured and calculated TIP readings surrounding hot bundle locations be examined in more detail so as to exclude effects of statistical averaging on the comparisons using plant instrumentation. The following figures and tables provide the information requested which demonstrate the accuracy of the nuclear methods for VYNPS for Cycles 23 and 24.
These data are the off-line calculated instrument readings compared to the measured instrument readings from the plant. The power and void fraction data associated with these figures and tables are from the off-line calculations using TGBLA and PANAC.
For each plant cycle, the following information is provided for one or more points in each cycle.
: 1. A summary through the cycle of TIP radial, axial and nodal core average RMS differences between calculated and measured TIPs. In addition, the summary indicates the radial, axial and nodal differences for the TIP instrument that is in the hot channel. In this case, the "hot channel" is defined as the instrumented location with the highest power bundle.
Cycle points where more detailed data are provided is also indicated.
: 2.      For each selected cycle exposure point (noted as 'B' for BOC, 'M' for MOC, or 'E' for EOC), a map by TIP string of the radial and axial TIP RMS difference across the core is provided. If the TIP instrument is "Failed", the TIP string was declared non-operational by the plant process computer and may be ignored. The failed TIP strings are marked in these maps.
: 3.      For each cycle exposure point selected for more detailed study, a plot of each measured and calculated TIP instrument reading as a function of axial height is provided. The TIP plots are arranged on a core-wide map to give the relative position of the TIP within the core.
: 4.      For each cycle exposure point selected for more detailed study, a numerical table with the nodal powers, exposures and void fractions corresponding to the four bundles surrounding the instrument adjacent to the hot channel and one additional channel is provided. The choice of the additional channel is arbitrary. An attempt was made to find a string with a different axial shape than the one adjacent to the hot channel.
: 5.      All of the TIP comparisons exclude node 1, 24, and 25. The TIP strings are 144 inches long and hence, go to node 24. Node 24 has been eliminated because on some plants there is TIP string hardware that distorts the signal. The power is very low in this node and does not affect the overall power distribution. Node 1 is eliminated because it usually is loaded with natural uranium, the power is low, and it contains the effects of the lower tie plate.
The overall statistics for the VYNPS core follow shows good power distribution agreement between the measured and calculated data. The average Radial RMS is ((                    )) for Cycle 23 and ((            )) for Cycle 24. The nodal RMS agreement is excellent, ((            )) for Cycle 23 and ((           )) for Cycle 24.
Tables SRXB-A-35-2 and SRXB-A-35-14 contain eigenvalue and thermal limit tracking data for Cycles 23 and 24 respectively. The eigenvalue behavior is consistent with that observed in all of the benchmark plants. The only exception to this consistent eigenvalue behavior is Cycle 11

Attachment 10 BVY 05-072 Docket No. 50-271 Page 27 of 126 of Plant E. This exception is not due to power uprate, but is due to an initial reload with a very large batch fraction (close to 50%).

Attachment 10 BVY 05-072 Docket No. 50-271 Page 28 of 126 VYNPS Cycle 23 Table SRXB-A-35-1 VYNPS TIP Comparison Data as a Function of Cycle 23 Exposure
((

Attachment 10 BVY 05-072 Docket No. 50-271 Page 29 of 126 Table SRXB-A-35-2   VYNPS Cycle 23 Summary Data
((

Attachment 10 BVY 05-072 Docket No. 50-271 Page 30 of 126

Attachment 10 BVY 05-072 Docket No. 50-271 Page 31 of 126 Table SRXB-A-35-3 VYNPS Cycle 23 and Cycle 24 Cold Critical Data 1JI
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Attachment 10 BVY 05-072 Docket No. 50-271 Page 32 of 126 Table SRXB-A 4 VYNPS Four Bundle Powers and TIP String Comparisons at BOC
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Attachment 10 BVY 05-072 Docket No. 50-271 Page 33 of 126 Figure SRXB-A-35-1 VYNPS Cycle 23 TIP String Comparisons at BOC
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Attachment 10 BVY 05-072 Docket No. 50-271 Page 34 of 126 Table SRXB-A-35-5 VYNPS Four Bundle Nodal Power & Void Fraction Comparisons for Hot Channel Instrument [String 12] at BOC 4                                 4

Attachment 10 BVY 05-072 Docket No. 50-271 Page 35 of 126 Table SRXB-A-35-6 VYNPS Four Bundle Nodal Power & Void Fraction Comparisons for Hot Channel Instrument [String 5] at BOC t                  1-
                                                                                                                ))

Attachment 10 BVY 05-072 Docket No. 50-271 Page 36 of 126 Table SRXB-A-35-7 VYNPS Four Bundle Powers and TIP String Comparisons Cycle 23 MOC
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Attachment 10 BVY 05-072 Docket No. 50-271 Page 37 of 126 Figure SRXB-A-35-2 VYNPS Cycle 23 TIP String Comparisons at MOC
                                                                    ))

Attachment 10 BW 05-072 Docket No. 50-271 Page 38 of 126 Table SRXB-A-35-8 VYNPS Four Bundle Nodal Power & Void Fraction Comparisons for Hot Channel Instrument [String 16] at MOC
: 4.                                     4

Attachment 10 BVY 05-072 Docket No. 50-271 Page 39 of 126 Table SRXB-A-35-9 VYNPS Four Bundle Nodal Power & Void Fraction Comparisons for Hot Channel Instrument [String 12] at MOC

Attachment 10 BVY 05-072 Docket No. 50-271 Page 40 of 126 Table SRXB-A 0 VYNPS Four Bundle Powers and TIP String Comparisons Cycle 23 EOC
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Attachment 10 BVY 05-072 Docket No. 50-271 Page 41 of 126 Figure SRXB-A-35-3 VYNPS Cycle 23 TIP String Comparisons at EOC
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Attachment 10 BVY 05-072 Docket No. 50-271 Page 42 of 126 Table SRXB-A-35-11 VYNPS Four Bundle Nodal Power & Void Fraction Comparisons for Hot Channel Instrument [String 6] at EOC 4-                    +              4-

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Attachment 10 BVY 05-072 Docket No. 50-271 Page 45 of 126 Table SRXB-A-35-14 VYNPS Cycle 24 Summary Data

Attachment 10 BVY 05-072 Docket No. 50-271 Page 46 of 126

Attachment 10 BVY 05-072 Docket No. 50-271 Page 47 of 126 Table SRXB-A-35-15 VYNPS Four Bundle Powers & TIP String Comparisons at Cycle 24 BOC
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Attachment 10 BVY 05-072 Docket No. 50-271 Page 48 of 126 Figure SRXB-A-35-4 VYNPS Cycle 24 TIP String Comparisons at BOC
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Attachment 10 BVY 05-072 Docket No. 50-271 Page 49 of 126 Table SRXB-A-35-16 VYNPS Four Bundle Nodal Power & Void Fraction Comparisons for Hot Channel Instrument [String 11] at BOC 4                  4                   4             4-

Attachment 10 BVY 05-072 Docket No. 50-271 Page 50 of 126 Table SRXB-A-35-17 VYNPS Four Bundle Nodal Power & Void Fraction Comparisons for Hot Channel Instrument [String 5] at BOC

Attachment 10 BVY 05-072 Docket No. 50-271 Page 51 of 126 Table SRXB-A-35-18 VYNPS Four Bundle Powers & TIP String Comparisons Cycle 24 MOCI
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Attachment 10 BVY 05-072 Docket No. 50-271 Page 52 of 126 Figure SRXB-A-35-5 VYNPS Cycle 24 TIP String Comparisons at MOC1
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Attachment 10 BVY 05-072 Docket No. 50-271
                                                                                                               - Page 53 of 126 Table SRXB-A-35-19 VYNPS Four Bundle Nodal Power & Void Fraction Comparisons for Hot Channel Instrument[String 11] at MOCI
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Attachment 10 BVY 05-072 Docket No. 50-271 Page 54 of 126 Table SRXB-A-35-20 VYNPS Four Bundle Nodal Power & Void Fraction Comparisons for Hot Channel Instrument [String 8] at MOCI 1]

Attachment 10 BVY 05-072 Docket No. 50-271 Page 55 of 126 Figure SRXB-A-35-21 VYNPS Four Bundle Powers & TIP String Comparisons Cycle 24 MOC2
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Attachment 10 BVY 05-072 Docket No. 50-271 Page 56 of 126 Figure SRXB-A-35-6 VYNPS Cycle 24 TIP String Comparisons at MOC2
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Attachment 10 BVY 05-072 Docket No. 50-271 Page 57 of 126 Table SRXB-A-35-22 VYNPS Four Bundle Nodal Power & Void Fraction Comparisons for Hot Channel Instrument [String 7] Cycle 24 MOC2 4                  +                  +             +

Attachment 10 BVY 05-072 Docket No. 50-271 Page 58 of 126 Table 35-23 SRXB-A-VYNPS Four Bundle Nodal Power & Void Fraction Comparisons for Hot Channel Instrument [String 16] Cycle 24 MOC2

Attachment 10 BVY 05-072 Docket No. 50-271 Page 59 of 126 RAI SRXB-A-36 The individual axial TIP data shows that the ((
                                                                           )) Explain how it can be ascertained that the PCTIP peaking being closer to specific bundle power peaking is not due to variability in the positions of the TIPs which may reflect the response of specific bundles and not necessarily the maximum powered bundles. Provide an evaluation that establishes the reason that the measured PCTIP reflects the response of specific bundles.
Response to RAI SRXB-A-36 It is important to point out that what is measured in these comparisons is the gamma TIP response and not the axial power. The gamma TIP axial response is close to the axial power distribution, but it is not an exact match. The gamma TIP response is proportional to the gamma flux at the instrument location. Most of the gamma rays come from the fission process, but some come from capture gamma rays emitted when a neutron is captured in the Gadolinium poison, so the source of gamma rays is not directly proportional to the fission energy, and varies with bundle exposure. Also, the TIP response varies with void fraction, because the gamma rays interact with the electrons in the water. This variation of TIP response with void fraction and exposure is not large, but will cause the axial TIP response to be different that the axial power. The 3-D simulator contains a gamma tip response function derived from explicit Monte Carlo gamma transport calculations, providing an accurate estimate of the relationship between the instrument response and the bundle power density adjacent to the TIP string.
RAI SRXB-A-37 Section 3.1.1 Model Uncertainty of NEDC-32601P-A discusses the method used to determine the accuracy of TGBLA in computing the fuel pin peaking factors. The accuracy of the TGBLA model was established by comparing its peaking factor distributions with Monte Carlo (MCNP) benchmark results. Table 3-1 presents the RMS differences between the MCNP and TGBLA rod differences for 8x8, 9x9 and 10x1O lattice designs
                                                                              )) Confirm that this is the case. Specifically, MCNP is not a depletion code and there are no biases assumed for MCNP calculations to demonstrate that the impact of the infinite pin power peaking with exposure is accounted for. Justify why it is acceptable to establish infinite lattice pin power peaking at BOL conditions to determine the uncertainties associated with pin power peaking as the bundles deplete under hard spectrums with fuel designs and loading that do not reflect the current core conditions.

Attachment 10 BVY 05-072 Docket No. 50-271 Page 60 of 126 Response to RAI SRXB-A-37 The comparisons presented in NEDE-32601P all came from TGBLAIMCNP comparisons performed at beginning of life. Since that time, a number of studies, using a 'snapshot' approach have shown that the maximum TGBLAIMCNP differences occur at zero exposure. In the snapshot approach, the TGBLA model is used to evaluate the isotopic abundances for each pin in the lattice, which are transferred into the MCNP model. The pin power differences between TGBLA and MCNP are then compared for an isotopic snapshot for a given exposure.
The results for a GE14 lattice design with a very high enrichment and Gadolinium content is shown in Figure 37-1. The snapshot process is conservative for evaluating RMS vs exposure because higher power pins will deplete faster than lower power pins resulting in lower U235 inventory, resulting in lower power for the next exposure. Hence power differences tend to even out with increasing exposure. This reduction in RMS with exposure was observed in the TGBLAILANCER studies presented to the staff. The peak RMS at beginning of life conditions is not surprising because the presence of poison rods makes the transport calculation more difficult. Fewer poison rods simplify the TGBLA evaluations. Studies on low enrichment lattices with no poison rods show TGBLAIMCNP RMS values of less than ((                     )) for all void fractions.
Figure SRXB-A-37-1 Pin RMS vs Lattice Exposure The average RMS obtained by the snapshot approach decreases with exposure. The 90%
RMS has been included up to an exposure of 13 GWd/st. Design studies show that at these high void fractions, there is very little burnup because of the low powers. ((
                                               )) The GEXL correlation relates the occurrence of dry out to the quality for a given flow condition. The quality is proportional to the axially integrated power. The R-factor relates the local quality near a given pin to the average quality over the entire lattice.

Attachment 10 BVY 05-072 Docket No. 50-271 Page 61 of 126 RAI SRXB-A-38 Table 3-1, "Summary of TGBLAI MCNP Pin Power Comparisons," of NEDC-32601P-A combines the RMS differences in rod power using historical GE fuel designs (8x8, 8x8 GE9, 9x9 GE11, 10x1O GE12 and 10x1O SVEA96) ((
                                                                                   )) Justify why the uncertainty calculation should not be limited to fuel designs with similar characteristics and operated conditions, since pin power peaking is dependent on the fuel design, loading (enrichment, Gd loading, etc.) and operation (e.g., high void conditions).
Response to RAI SRXB-A-38 GE agrees that the infinite lattice peaking factor uncertainties should be based on up to date product lines (GE14) and designs. To this end, the TGBLAIMCNP snapshot burnup study was carried out on an aggressive application of the GE14 product. The lattice forming the basis for Figure SRXB-A-37-1 is a GE14 application with a lattice enrichment of ((
                                                                                      )) Gadolinium concentration. As illustrated in the reply to RAI SRXB-A-37, the peak RMS ((
      )). A series of TGBLNMCNP comparisons have been carried out on GE14 lattices representative of EPU applications. The updated data set consists of 13 beginning of life lattices, and has an average pin power RMS of ((            )). The average value of ((           )) is less than the ((       )) uncertainty derived from an extensive and conservative review of pin-by-pin gamma scan data obtained by GE. This ((                   )) is referred to in the Alternate Approach.
RAI SRXB-A-39
((:
                                                                                           )) The fuel designs used to establish the gradient peaking ((         )) is not discussed in the SLMCPR LTR.

Attachment 10 BVY 05-072 Docket No. 50-271 Page 62 of 126 a)  State if similar ((                                  )) analyses were performed for cores loaded with GE14 fuel and core designs that reflect the current and proposed operating conditions.
b)  Demonstrate the validity of the currently used gradient peaking of ((         )) for the current operating strategies and fuel designs (GE14). State if similar [l
                                                                  )) to establish the influence of neighboring bundles on the local peaking distribution.
c)  State whether the core simulations calculations are performed at different exposures. If not, establish a methodology that accounts for the exposure dependence.
Response to RAI SRXB-A-39 a)  One of the four DIF3D studies documented in NEDE-32601P-A (BWR/6 800 Bundle Equilibrium Core) was performed using the GE12 lattice configuration, which is identical to GE14. The average gradient peaking RMS obtained for this application was ((                    j],
which is quite consistent with the RMS values obtained with the other cores. This application was based on an equilibrium design for a large BWR/6 with an 18 month cycle, consistent with current power uprate strategies.
b)  Currently, a value of ((       )) is used to represent the gradient uncertainty. As pointed out in part a), the 800 bundle BWR/6 example is quite representative of current applications. No studies have been carried out at void fractions larger than [l          ))- The results in Table 3-2 of NEDE-32601P-A show ((
c)  The core simulations were performed on beginning of cycle configurations. A set of bundles with different exposures were used make up the core configuration. High and medium exposure lattices were placed adjacent to each other as one might see in a typical beginning of cycle core configuration. The BWR/6 core is shown below in Table SRXB-A-39-1.

Attachment 10 BVY 05-072 Docket No. 50-271 Page 63 of 126 Table SRXB-A-39-1 Quarter Core Map for DIF3D Benchmark rr 11

Attachment 10 BVY 05-072 Docket No. 50-271 Page 64 of 126 RAI SRXB-A-40
                              )) State if the current plant-specific SLMCPR calculations ((
             )) uncertainty was included in the ((                   )).
Response to RAI SRXB-A-40 The ((                    )) uncertainty is included in the plant specific SLMCPR calculations and in the determination of additional margin discussed in the alternate approach section.
RAI SRXB-A-41 In the review of the SLMCPR methodology, the staff asked in RAI 5 the following, " The process computer monitors peak kw/ft and MAPLHGR. While MCPR depends primarily on the radial bundle power distribution, peak kw/ft and MAPLHGR depend on the bundle axial power distribution and, consequently are significantly more sensitive to the 3-D MONICORE replacement of the TIP/LPRM axial power distribution. Provide an uncertainty analysis for the 3-D MONICORE prediction of peak kw/ft and MAPLHGR." Update this RAI response. Justify why the peak Kw/ft uncertainty (see RAls addressing changes in the peak and nodal uncertainties) would not change for the current operating strategies and fuel designs (GE14).
Describe how the uncertainty is applied in 3D MONICORE.
Response to RAI SRXB-A-41 Table SRXB-A-41-1 is a summary of all the components included in the total pin LHGR uncertainty for the process computer. Listed are the values originally approved in NEDE-32601 P-A and those proposed for application to EPU designs.
Table SRXB-A-41-1 Summary of Uncertainty Components for Process Computer LHGR Evaluations
((
                                                  .1               4

Attachment 10 BVY 05-072 Docket No. 50-271 Page 65 of 126 1))
The revised infinite lattice and power allocation (PAL) uncertainties are discussed in the alternate approach section. Since the publication of NEDE-32601, the ((
                                                    )). The PANAC11 version of the 3-D simulator contains an improved model for acting for the impacts of flux gradients on the pin power. The DIF3D studies described in RAI SRXB-A-39 show PANAC11 Gradient effects uncertainty to be
((       )). This ((        )) was representative of both the entire data set and the single 10x1O lattice core. The revised total uncertainty is ((           )). A local uncertainty of ((
is assumed in the generation of the thermal mechanical limit values.
RAI SRXB-A-42 Section 3.2, 'Conversion of Peaking Uncertainty to R-factor Uncertainty" of NEDC-32601P-A discusses that the R-factor represents the influence of the rod power peaking on the critical power. In addition, the R-factor methodology is described in NEDC-32505P, "An R-Factor Calculation Method for GE11, GE12, and GE13 Fuel," dated July 1999. The bundle R-factor is an input to the GEXL correlation. ((
                                                        )) Tables 3.4a, b, and c of NEDC-32601-P-A provide the basis for the((                                                        )). Explain how the uncertainties in the lattice physics pin power data are accounted for in ((
                                                                           )) Explain how the void and exposure dependency of the uncertainties of pin peaking factors is incorporated in the R-factor methodology. Specifically, explain how the exposure dependence of pin powers uncertainties are established if the infinite lattice pin power uncertainties are established using MCNP/TGBLA code to code benchmarking based in BOL or the standard GE MCNP/TGBLA comparisons with exposure is used.
Response to RAI SRXB-A-42 Based on questions raised by the staff, the procedure for evaluating the R-factor uncertainty was modified. The modified procedure is documented in NEDE-32601-PQC and is the basis for the currently used R-factor uncertainty of ((          )). This approach takes into account the fact that the ff                                                                         )) The lattice model uncertainties are based on the beginning of life MCNP/TGBLA comparisons based on the

Attachment 10 BVY 05-072 Docket No. 50-271 Page 66 of 126 arguments presented in the response to RAI SRXB-A-37. ((
                                             )) The overall distribution of R-factor uncertainties is illustrated in NEDE-32601-PQC.            This procedure is used to evaluate all new R-factor uncertainties based on updated model and channel bow uncertainties described in the alternate approach.
RAI SRXB-A-43 The subcooling increases for the EPU conditions in comparison with operation at rated power and lower domains. Justify why the uncertainty due to core inlet temperature (Section 2.5 NEDC-32601 P-A) would not change.
Response to RAI SRXB-A-43 The basis of the core inlet temperature uncertainty is discussed in Section 2.5 of NEDC-32601P-A. In this discussion, it states that "It is concluded that the core inlet temperature uncertainty of ((        )) specified in the (GETAB Analysis Basis) is adequately conservative to accommodate as much as a factor of ((
                                     )) stated in that reference. This level of conservatism is judged to be adequate to accommodate all plant specific variations."
RAI SRXB-A-44 State if VYNPS EPU conditions would result in any bypass voiding due to the high bundle power conditions. The LPRM uncertainty increases with increasing void, ((
                                                )) If any non-solid bypass voiding would occur during steady state, evaluate the LPRM and TIP uncertainties and justify why the current uncertainty based on zero bypass voiding remain applicable. Consider an increase in the random noise. If bypass voiding does occur during transient events (e.g., RPT) and plant maneuvers in the off-rated high power/low conditions, provide an evaluation of the impact of non-solid bypass voiding on the reliability and accuracy of the instrumentation.
Response to RAI SRXB-A-44 The design specification for limited local voiding in the bypass region at the uppermost LPRM detector for the highest power four-bundle configuration assures that the effects of voiding are insignificant for the limiting fuel operating conditions. The presence of bypass voids at the uppermost LPRM detector can result in a small decrease in measured power by thermal neutron detectors for that location due to reduced moderation of the neutron flux. However, since the power is divided among the four fuel assemblies and averaged with the larger measured power in the lower, un-voided fuel nodes, the total effect of the maximum voids was determined to be ((                                                      )) for the MCPR calculation.

Attachment 10 BVY 05-072 Docket No. 50-271 Page 67 of 126 Furthermore, it should be noted that VYNPS has installed gamma TIPS which will normalize the LPRM reading to the correct nodal power level since they are not sensitive to bypass voids, and this would essentially eliminate the impact of bypass voids for steady state operation.
The evaluation of the impact of non-solid bypass voiding on the reliability and accuracy of the instrumentation is based on a calculation of the MCPR overprediction by the process computer when the LPRM reading is not properly normalized to the correct power. The TIP system is used to both normalize the LPRM readings throughout the core, and establish the axial shape between the LPRM readings for the process computer. The TIP system is utilized periodically after core power and power distribution changes, and after LPRM sensitivity changes. ((
         ))
The bypass voiding could increase the random noise at the LPRM due to passage of the bubbles around the LPRM detector as explained in the response to RAI SRXB-A-55. However, since the void level is less than 5%, the extra noise magnitude is small, and since this extra noise would add to the existing noise as the square root of the sum of the squares, the impact of

Attachment 10 BVY 05-072 Docket No. 50-271 Page 68 of 126 the extra noise is negligible. Moreover, since the voiding is only at the D level, its effect on the bundle power measurement is insignificant.
There is no reliance on the process computer calculation of MCPR during transient conditions, and therefore no requirement exists for monitoring MCPR under local bypass voiding for conditions outside the normal steady state operating map (such as after a two pump trip). The impact of bypass voids after a two pump trip, and other effects of bypass voiding during transient conditions are addressed in the response to RAI SRXB-A-55.
RAI SRXB-A-45 Explain why the axial TIP RMS is not included in the calculation of the SLMCPR limit, when the SLMCPR limit calculation is non-adaptive. The staff understands that the core ((
                                                                       )). Therefore, explain how the axial power distribution is accounted for in the (offline) calculation that establishes the SLMCPR value.
Response to RAI SRXB-A-45 The purpose of the use of the 3-D simulator in the evaluation of the SLMCPR is twofold:
: 1. The SLMCPR methodology sums up contributions of the probability of boiling transition from every node in the core, and hence is affected by the relative power distribution. A flat power distribution results in a higher SLMCPR because more nodes are near the limiting node and contribute materially to the probability of boiling transition. The 3-D simulator is used to establish a limiting power distribution, usually one that is as flat as possible and then the power is increased until one of the bundles is on limits. This search for the limiting power distribution leads to possible operating distributions, and since the limiting bundle is arbitrarily placed on limits, has little to do with the uncertainty of the power in that bundle. The axial power is calculated to be consistent with the core state. For a given core state, the critical power is a very weak function of small changes in the axial power.
Further, the RMS difference of all predictive to measured TIP response results in an average difference of less than ((         )) for the reference BWR's, indicating that axial power distributions are also predicted adequately. Changes in TIP and LPRM signals affect the axial power and are modeled in the SLMCPR process.
: 2. The 3-D simulator is used to evaluate the impact of changes in operating conditions such as operating pressure, core flow, instrument response, pressure drop and a variety of other inputs. Since the simulator is used to evaluate changes relative to a base state, the absolute accuracy of the model does not contribute to the magnitude of the change.

Attachment 10 BVY 05-072 Docket No. 50-271 Page 69 of 126 Since the intended purpose of the 3-D simulator in the SLMCPR evaluation does not depend on the axial power accuracy and only to set the initial conditions and to evaluate changes, the predictive mode is used to set the initial power distribution.
RAI SRXB-A-46 Describe GE's standard approach for determining the cross-sections for operation at high void conditions (fit/interpolate/extrapolate). Discuss why the cross-sections are not directly generated for void fractions above 70%, using TGBLA.                Describe any code-to-code comparisons that were performed in order to assess this method at high void fractions.
Response to RAI SRXB-A-46 The extrapolation accuracy beyond 70% voids has been discussed extensively in Section 2.1 of MFN-04-026. In these studies the extrapolated inputs to the 3-D simulator were compared to MCNP Monte Carlo values. Despite the fact that extended operation at 90% voids is rare, even for the highest power density plants, the extrapolation errors were shown to be small and acceptable for use in the 3-D simulator. Further, the application of the current 3-D simulator to high power density applications show little or no degradation in methods accuracy. Finally, the margin proposed in the Alternate Approach is designed to provide assurance that the accuracy and margin standards are upheld.
RAI SRXB-A-47 Provide plots of the isotopic concentrations and fission fractions of U-235, U-236, U-238, Pu-239, Pu-240, and Pu-241 as functions of burnup. Use lattices that are limiting in terms of enrichment and the number of hot pins and Gd concentration. Present the isotopic concentration vs. exposure plots for depletion at 40%, 70%, and 90% void conditions. The objectives are to baseline potential changes due to spectral hardening for operation at different void conditions and to determine both the accuracy of TGBLA and how TGBLA's accuracy changes with void fraction. Provide plots similar to the plots in Figures 3-8 and 3-9 of NEDE-20944-P for depletion at different void conditions and for different lattices.
Response to RAI SRXB-A-47 Plots of the atomic density of the identified isotope as a function of exposure and void fraction are provide in the attached figures. U-234 is included to identify that an initial fraction of U-234 is present in the initial isotopic inventory.
Five lattices designed for use in Vermont Yankee Cycle 25 are provided for use in observing the effect of void fraction on the transmutation of U-238 into higher atomic number actinide

Attachment 10 BVY 05-072 Docket No. 50-271 Page 70 of 126 isotopes. The range of data provided is 0, 40, 70, and 90% void fractions and the exposure range from 0 to 65 GWd/st. The lattices provided are as follows:
Lattice   Number Lattice            Lattice Nuclear Name Designation         Type 696Dominant     P1 ODNAL453-16G6.0-10OT-T6-6996                D6996 6997    6997Domiant Dominant        P1 ODNAL453-12G6.0-10OT-T6-6997 6999    6999Vanihed Vanished        P1 ODNAL448-12G6.0-10OT-V-T6-6999 P1ODNAL413-14G6.0-10OT-T6-7007                Dominant        7007 7009    7009        Vanished Vnishe          P1 ODNAL403-14G6.0-10OT-V-T6-7009 For each lattice analyzed there are 7 plots that contain the isotopic inventory in the units of atom density for the required isotopes. The enrichment and gadolinium are given in Table 1 of each section. The plots of isotopic concentration in order of presentation are U-234, U-235, U-236, U-238, Pu-239, Pu-240, and Pu-241. A list figures and tables is provided below for plot and table identification.
Table SRXB-A-47.1-1 Enrichment and Gadolinium Distribution for Lattice 6996 Figure SRXB-A-47.1-1         Lattice 6996 U234 Isotopic Concentration Figure SRXB-A-47.1-2         Lattice 6996 U235 Isotopic Concentration Figure SRXB-A-47.1-3         Lattice 6996 U236 Isotopic Concentration Figure SRXB-A-47.1-4         Lattice 6996 U238 Isotopic Concentration Figure SRXB-A-47.1-5         Lattice 6996 Pu239 Isotopic Concentration Figure SRXB-A-47.1-6         Lattice 6996 Pu240 Isotopic Concentration Figure SRXB-A-47.1-7         Lattice 6996 Pu241 Isotopic Concentration Table SRXB-A-47.2-1 Enrichment and Gadolinium Distribution for Lattice 6997 Figure SRXB-A-47.2-1         Lattice 6997 U234 Isotopic Concentration Figure SRXB-A-47.2-2         Lattice 6997 U235 Isotopic Concentration Figure SRXB-A-47.2-3         Lattice 6997 U236 Isotopic Concentration Figure SRXB-A-47.2-4         Lattice 6997 U238 Isotopic Concentration Figure SRXB-A-47.2-5         Lattice 6997 Pu239 Isotopic Concentration Figure SRXB-A-47.2-6         Lattice 6997 Pu240 Isotopic Concentration Figure SRXB-A-47.2-7         Lattice 6997 Pu241 Isotopic Concentration Table SRXB-A-47.3-1 Enrichment and Gadolinium Distribution for Lattice 6999

Attachment 10 BVY 05-072 Docket No. 50-271 Page 71 of 126 Figure SRXB-A-47.3-1       Lattice 6999 U234 Isotopic Concentration Figure SRXB-A-47.3-2       Lattice 6999 U235 Isotopic Concentration Figure SRXB-A-47.3-3       Lattice 6999 U236 Isotopic Concentration Figure SRXB-A-47.3-4       Lattice 6999 U238 Isotopic Concentration Figure SRXB-A-47.3-5       Lattice 6999 Pu239 Isotopic Concentration Figure SRXB-A-47.3-6       Lattice 6999 Pu240 Isotopic Concentration Figure SRXB-A-47.3-7       Lattice 6999 Pu241 Isotopic Concentration Table SRXB-A-47.4-1 Enrichment and Gadolinium Distribution for Lattice 7007 Figure SRXB-A-47.4-1       Lattice 7007 U234 Isotopic Concentration Figure SRXB-A-47.4-2       Lattice 7007 U235 Isotopic Concentration Figure SRXB-A-47.4-3       Lattice 7007 U236 Isotopic Concentration Figure SRXB-A-47.4-4       Lattice 7007 U238 Isotopic Concentration Figure SRXB-A-47.4-5       Lattice 7007 Pu239 Isotopic Concentration Figure SRXB-A-47.4-6       Lattice 7007 Pu240 Isotopic Concentration Figure SRXB-A-47.4-7       Lattice 7007 Pu241 Isotopic Concentration Table SRXB-A-47.5-1 Enrichment and Gadolinium Distribution for Lattice 7009 Figure SRXB-A-47.5-1       Lattice 7009 U234 Isotopic Concentration Figure SRXB-A-47.5-2       Lattice 7009 U235 Isotopic Concentration Figure SRXB-A-47.5-3       Lattice 7009 U236 Isotopic Concentration Figure SRXB-A-47.5-4       Lattice 7009 U238 Isotopic Concentration Figure SRXB-A-47.5-5       Lattice 7009 Pu239 Isotopic Concentration Figure SRXB-A-47.5-6       Lattice 7009 Pu240 Isotopic Concentration Figure SRXB-A-47.5-7       Lattice 7009 Pu241 Isotopic Concentration

Attachment 10 BVY 05-072 Docket No. 50-271 Page 72 of 126 47.1 Lattice PIODNAL453-16G6.0-IOOT-T6-6996 Isotopic Results ER[
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Attachment 10 BVY 05-072 Docket No. 50-271 Page 112 of 126 RAI SRXB-A-48 With respect to the accuracy of the steady state neutronic methods for current fuel design (GE14) and operating strategies, demonstrate that possible variations in isotopic content, under applicable EPU conditions, have been accounted for including consideration of void reactivity coefficient.
Response to RAI SRXB-A-48 In the response to RAI-25, RAI-27, and RAI-29 of MFN 05-029 TAC No. 5780, data have been presented which show that under power uprate conditions of 20%, the exit void fraction increases by about 3% on average. Figure 29-5 of RAI-29 shows that under high power uprate and long cycle conditions, exposure weighted void histories do not exceed 85% and those nodes with voids greater than 83% rarely exceed 20 GWD/st.
At this point in exposure, the lattice reactivity is changing and the void coefficient is dependent on the inventory of four major isotopes: Gd157, Gd155, U235, and PU239 (U238 is important, but the percentage change with exposure is small). The total change in void reactivity over this range from BOL to 20 GWD/st is about a factor of two based on a number of factors, including Gd depletion, Pu buildup and U235 depletion. Changes in depletion history can change the detail of the overall void coefficient behavior, but these effects are much smaller than the overall change. The major driver for transient response is not the effect of isotopics on void coefficient, but the influence of lattice depletion and operating strategy on the axial power distribution at the end of an operating cycle. For pressurization transients, which are the most severe transients for most BWR's, including VYNPS, the single most important parameter is the scram reactivity, which influences how effectively the transient is mitigated. Operating history can have a significant effect on scram reactivity, as shown in Figure SRXB-A-48-1.

Attachment 10 BVY 05-072 Docket No. 50-271 Page 113 of 126 Figure SRXB-A-48-1 Axial Power and Exposure 1.I                                                                                            -  35000 1.2    -                                                                  000CtS I
                 *1                                                                                                 5000 i   0.8      _  t* i    .                                           '         -   4   -,-           - 21'7000 i   0.6
* I,,/-
                          .,,...,,,.1.l
                          .,,, '.1             :.                                   ',.....                         5000 ,
I 0.4 I 000  I
_   *      ''1lj                >                 ,j i,* ,    I' , HR',\,,._1
                                                                                 ,4;,,>   ; b_.r 0.2                                                                                                  ;0004 0
0                    5               10           15                    20             25 Axial Node (I=bottom, 25-top)
                     -    Nominal Power              HBB Power -Nominal      Exposure -              HBB Exposure l Note that the end of cycle HBB exposure shape is peaked to the bottom of the core relative to the nominal exposure and has a significant impact on the EOC power distribution. The HBB power shape is peaked to the top of the core. This power shape is conservatively used to determine the end of cycle transient response and is significantly different from the nominal power shape. As the control blades enter the bottom of the core, the scram strength is sensitive to the power level near the bottom. This difference in power shape is an important nuclear contributor to the transient response. It is influenced by the exposure and core isotopics, in that they play a role in the overall power shape. Their influence on void coefficient is secondary.
For this reason, the HBB strategy has become the primary mechanism for ensuring conservative nuclear inputs to the transient evaluation process.
RAI SRXB-A-49 With respect to the GEXL correlation, for the VYNPS EPU conditions, state if any double-hump power shape is projected. Describe the methods and criteria used to determine that no additional SLMCPR penalty should apply. If any double hump power shape is predicted irrespective of the corresponding bundle power level, justify why it is acceptable to predict the MCPR performance of a bundle using a correlation that was developed with out the specific power shape, without adding a bias.
Response to RAI SRXB-A-49 The GEXL correlation for double humped axial power shape has been evaluated by comparison to data trends from GETAB (Ref. 1) and KTH (Ref. 2) and by benchmarking to detailed

Attachment 10 BVY 05-072 Docket No. 50-271 Page 114 of 126 subchannel calculations. These comparisons all showed a consistent trend that the critical power was overpredicted by approximately ((                                    )) for most GNF-A fuel products. These comparisons were then used to develop a bounding penalty at the ((                    ))
confidence level to the bias and to the uncertainty of the GEXL correlation for the double humped axial power shape. When the SLMCPR is evaluated, the axial power shape is examined for all bundles in the core, and the above bias is applied to those bundles that have a double humped axial power shape. Note, the SLMCPR is determined on a cycle specific basis.
This process was described in detail to the NRC at the 2002 technology update meeting (Ref.
3). There were no double humped axial power shapes for the VYNPS EPU cycle 25.
RAI SRXB-A-50 For the transient and LOCA events analyzed along the proposed operating domain statepoint, state if the power/flow ranges fall within the data base for which the GEXL correlation was developed. Give specific examples.
Response to RAI SRXB-A-50 The critical power correlation for GNF-A fuel is developed from critical power test data. The critical power data are obtained for bundle mass fluxes ranging from ((
                                )), inlet subcoolings ((
                             )) and pressures from ((                      )). These data cover flow ranges from less than natural circulation to well beyond rated flow and include the flow ranges for EPU/MELLLA+ applications. As these data cover power levels up to the critical power for each flow, void fractions up to [f          )) are included in the data. These parameter ranges also cover the expected ranges for LOCA and transient events. The process follows the NRC approved process as described in GESTAR II (Ref. 1). An example of this process is shown in Figure SRXB-A-50-1 that shows the GE14 application range together with the expected range for typical operational transients. For LOCA application, the GEXL correlation is used for the calculation of the early boiling transition during the flow coast down immediately following the break. This typically occurs when the flow has dropped to 30-50% of the initial value. This is well within the application range for the GEXL correlation.
((

Attachment 10 BVY 05-072 Docket No. 50-271 Page 115 of 126 RAI SRXB-A-51 Provide an evaluation that demonstrates that the void reactivity coefficients are applicable and were developed for the ranges of core thermal-hydraulic conditions expected for the transient and accident conditions, including ATWS and SBO.
Response to RAI SRXB-A-51 The Alternative Approach section summarizes the use of the 3-D simulator as the basis for all of the nuclear input to transient evaluations. The 3-D simulator has a bypass void model and the nuclear inputs are accounted for. The discussion in the answer to SRXB RAI-48 lists the reactivity response to axial power shapes changes and the physical reason that for all transients (including ATWS), the most important feedback from void reactivity comes from the power at the bottom of the core. The nuclear input is supplied on a nominal basis, whether it is supplied to the ODYN, TRACG, or stability evaluation tools. Conservatism is accounted for by operating assumptions, and in the case of the TRACG code, extensive statistical analysis base on variations of input parameters.

Attachment 10 BVY 05-072 Docket No. 50-271 Page 116 of 126 The VYNPS SBO analyses do not specifically model any particular transient or accident conditions with respect to core performance. Therefore, void reactivity coefficient changes do not influence the results.
RAI SRXB-A-52 Demonstrate that the pressure drop measurement database include GEI4 fuel designs and the test were performed to cover the ranges expected for operation at EPU conditions. Include in your response, the maximum range of bundle power to flow ratios and bundle mass flux over which the data were taken. With respect to bundle delta pressure validation, the information provided to date indicates an underprediction in delta pressure for high bundle power to flow ratios. How are the delta pressure uncertainties applied?
Response to RAI SRXB-A-52 The range of mass flux and power to flow ratio for the GE14 pressure drop measurements are given in Section 2.2 Figures 2-41 and 2-42 of MFN 04-026. The pressure drop measurements cover from 30% of rated flow up to 120% of rated flow. The maximum mass flux in Figure 2-41 of 1.5 Mlb/hr/ft 2 corresponds to a bundle flow of 148,000 lb/hr and covers power to flow ratios up to 1.52(MW/(lb/hr))*10E-4, compared with a maximum of 0.9(MW/(lb/hr))*10E-4 observed in the compilation of bundle power to flow ratios (Figure 6-2 of BVY05-024). This maximum power to flow ratio is therefore 50% larger than expected in EPU applications. All predictions are within
[l        )) of the measured values, with a maximum of ((           )) at high power to flow values.
The standard deviation of all the errors is ((               )). An uncertainty of ff            )) is assumed in the SLMCPR analysis documented in NEDC-32601P-A. A change in channel pressure drop of ((        )) has a negligibly small impact on the SLMCPR.
RAI SRXB-A-53 The document states that the ASEA-813 and ASEA-513 tests, which varied rod power distributions and side-to-side or in-out skew, were not included in the correlation development because of concerns over bias in the measurement. Enclosure 3 adds that there are many points of consistent overlap with ASEA-713 and the data serves well to validate the correlation developed using the most accurate data sources. Provide a tabulation or other means of justification that shows the data points supporting the validation of the correlation to the high void ranges. Identify the corresponding power profiles for the specific data set, the applicable test conditions /ranges (e.g., the bundle power/flow ratios.) and fuel designs and characteristics.

Attachment 10 BVY 05-072 Docket No. 50-271 Page 117 of 126 Response to RAI SRXB-A-53 Consideration of uncertainties in the void correlation and where they are actively considered in the GE licensing methodology are addressed in the response to RAI SRXB-A-54. Justification of the void-quality correlation in use by GE is contained in Attachment A (page A-39) of NEDC-32601 P-A, "Methodology and Uncertainties for Safety Limit MCPR Evaluations". This attachment addresses sensitivity of the void fraction to bundle type, water rods, part length rods, spacers, and pin peaking.
RAI SRXB-A-54 An EPU or a high density plant can have an exit void fraction of ((
                                                )) in the corresponding water density calculations?
Response to RAI SRXB-A-54 The void fraction prediction from the core simulator is best-estimate and further utilized for identification of appropriate cross sections for the state. The uncertainty in the void fraction impacts the flow and power distributions. As a result, any inaccuracy in the void fraction prediction is quantified as a part of the bundle & nodal TIP comparisons. (e.g. see the response to RAI SRXB-A-35). It is to be noted that such comparisons also include the direct effects of core instrumentation & measurement uncertainty. Therefore, the total uncertainties in these parameters and therefore also the effect of the void fraction uncertainty are accounted for in the SLMCPR.
VY EPU void fractions for realistic core designs (Cycle 24 and Cycle 25) are not expected to exceed 87% while the database for the void-quality correlation extends to 95% void fraction.
RAI SRXB-A-55 The VYNPS response to RAI SRXB-A-6 states that the review of the steady state calculations at natural circulation indicates that the ((
                                                                        )). The RAI response then justifies applying the ((
                                     )) would not have significant impact on the thermal-hydraulic conditions and, therefore, the coupled reactivity feedback would not be affected.
As an Option 1D plant, VYNPS relies on APRM scram to provide SLMCPR protection for operation in the high powered/low flow zone of the power/flow map. For void fractions greater

Attachment 10 BVY 05-072 Docket No. 50-271 Page 118 of 126 than ((                                                     )), issues that had not been addressed in the response are whether, (1) the SLMCPR would be protected before an ARPM scram occurs in order to meet GDC-12, and (2) would the reliability and effectiveness of the APRMs for stability protection be impacted by the high void conditions. Specifically, with 2RPT event conditions, [
                                                                       )) the instrument noise, and potentially the temperature, could increase, affecting the reliability of the APRMs relied upon for instability protection. Right after a 2RPT event, the core thermal-hydraulic conditions may result in an SLMCPR value that is higher than for the cycle-specific SLMCPR value. In addition, the coupled neutronic and thermal-hydraulic response at the cited void ranges may change depending on the response of the void reactivity coefficient with void fraction. Therefore, the RAI response is inadequate.
Provide an evaluation that considers all impacts discussed above, including the any increases in the uncertainties of the ((
                                    )). For any conclusions or assessment made, please provide the supporting bases.
Response to RAI SRXB-A-55 The impact of high in-channel voids (such as those that could occur under 2 RPT conditions) on the methods and on SLMCPR determination, is described in the alternate approach. This response will focus on the impact of the high bypass voids on the protection provided by the APRM scram in mitigating stability events under 2 RPT conditions.
The primary effect of voiding in the bypass region on the neutron detectors (LPRMs and TIPs) is to reduce the detector response, assuming the same power in the adjacent fuel. This reduction is due to a decrease in the moderation which decreases the thermal neutron flux incident on the detectors for the same neutron flux generated in the adjacent fuel. There is also the potential for some additional noise in the neutron flux signal, but that has minor impact on steady state operation. These impacts are greatest for the highest elevation LPRM (D level) where the highest bypass voiding occurs, and are discussed quantitatively below.
: 1. Impact of Bypass Voids on LPRM and TIP
((

Attachment 10 BVY 05-072 Docket No. 50-271 Page 119 of 126
                                                                        ))
The impact of bypass voids on the APRM can be estimated by determining the bypass voids
((     )) at each of the ((         )) 4 LPRM elevations (A, B, C and D) and each of [         1]
LPRM strings in that APRM, reducing the LPRM signals ((                       )) and averaging them, and then comparing this average to average of the LPRM signals Rid without the reduction due to bypass voids.
2 Impact of High Bypass Voids on VY Option 1D Stability Solution For VY Option 1D stability solution, the APRM flow biased scram is used to mitigate stability transients. The analytical limit for the scram setpoint is based on assuring that scram occurs before stability oscillations get large enough to cause the MCPR to approach the SLMCPR.
High bypass voids can potentially reduce the APRM reading, and so the margin to scram would increase and this could be non-conservative from the stability mitigation point of view since it would take higher amplitude oscillations to initiate APRM scram.

Attachment 10 BVY 05-072 Docket No. 50-271 Page 120 of 126 1))
: 3. Bases of the VY Option 1D Stability Setpoint Analytical Limit Validity The detect and suppress evaluation for the VYNPS EPU is performed under the approved LTR basis (NEDO-32465-A, General Electric Company, "Reactor Stability Detect and Suppress Solutions Licensing Basis Methodology for Reload Applications," August 1996.). The flow biased APRM scram setpoint analytical limits are initially established with conservative margin such that they are found applicable to future fuel cycles during reload confirmation calculations.
The calculation of the scram setpoint analytical limits is based on the limiting fuel bundle being at the Operating Limit MCPR and the Safety Limit MCPR not being exceeded during the instability oscillation.
The detect and suppress evaluation for Cycle 24 under EPU conditions has been re-performed to demonstrate the adequacy of safety margins should a thermal-hydraulic instability event develop at VYNPS.
The detect and suppress calculation assumes a flow runback along the rated licensing rodline to natural circulation flow. The APRM flow-biased flux scram setpoint analytical limit has been rescaled for EPU operation. The flow-biased APRM trip analytical limit at natural circulation is 53.7% of rated power. The best-estimate power level on the rated rodline at natural circulation is 49.4% of rated power. [
1))
The discussion in item 2 of this response shows that [
                                             )) because of the conservatism in the stability evaluation, the margin to current setpoint analytical limit (AL) is adequate to include the effect of bypass voids, and that no downward adjustment of the setpoint is required.

Attachment 10 BVY 05-072 Docket No. 50-271 Page 121 of 126
                )) Hence, the SLMCPR is fully protected for VY EPU 2RPT runback conditions.
: 4. Impact on Noise due to Bypass Voidinq The increased voiding in the bypass region could potentially affect (increase) the LPRM noise because of the steam bubbles going by the LPRM instrument assembly in the water gap. The increase in noise depends upon the bubble dynamics as described below. Note that the discussion in this section refers only to the extra noise caused by the bubbles in the bypass region, and not the normal noise (-2% for APRM) that is present because of the flow induced vibration of the LPRM assembly in the water gap and because of other thermal-hydraulic phenomena inside the fuel channels. Note that there is an additional LPRM detector noise component due to the random nature of the process that produces neutron flux, which is proportional to the square-root of the neutron flux and is the source of the signal used for the IRM detectors. However the magnitude of that noise for LPRMs is small in the low frequency range of interest in this measurement, and does not need to be considered for this evaluation.

Attachment 10 BVY 05-072 Docket No. 50-271 Page 122 of 126 The noise due to bypass voids slightly increases the overall APRM neutron noise at off rated conditions where the voids may be significant. However the impact of this noise on the APRM scram setpoint is negligible because the setpoint (derived from the analytical limit by considering noise and other instrument errors) is based on the normal (no void) noise at rated conditions (-2% of rated power), and this bounds the increased noise at off-rated conditions since the decrease in normal noise at off-rated conditions is more than the increase due to bypass voiding.
: 5. Impact of LPRM Temperature Increase due to Bypass Voiding There is no significant increase in detector temperature due to an increase in bypass flow voids.
The basis for this conclusion is that bypass flow temperature external to the LPRM tubes is limited to a range from approximately 40 degrees Fahrenheit subcooled to saturation and that the heat transfer coefficient on the LPRM tube surface is not sensitive to the void content.
Additionally, LPRM tubes for BWRI2 to BWRI5 plants have internal cooling flow that maintains even lower temperatures at the LPRM detectors and in internal tube component surfaces. Note that LPRM and TIP detectors are ionization chambers made of high temperature materials and are capable of operating reliably at higher temperatures. For example, the BWR-6 LPRM detectors operate in a dry tube, where the temperatures are higher.

Attachment 10 BVY 05-072 Docket No. 50-271 Page 123 of 126

==References:==
: 1. NEDO-32465-A, Reactor Stability Detect and Suppress Solutions Licensing Basis Methodology for Reload Applications, August 1996
: 2. OG01-0228-001, Additional information - Determination of Figure of Merit for Stability DIVOM Curve Applicability, July 16, 2001 (updated at NC and 45% flow, 1/16/2004)
RAI SRXB-A-56 Provide an evaluation of the impact on instrument random noise and reliability for operation at high power/low flow conditions during plant maneuvers and SLO operation.
Response to RAI SRXB-A-56 This discussion addresses the treatment of noise in the instrument readings for both high and low power and flow. The EPU includes a new operating region in the high power high flow operating conditions only. The operation below - 75% flow in terms of absolute power and flow, which includes single loop operation (SLO), is the same as pre-EPU MELLLA, so there are no new low flow noise effects with EPU for VYNPS.
The effect of process noise on both the power and flow instrumentation is accounted for as uncertainties in the calculation of the SLMCPR, and in the calculation of APRM setpoints and these are different for two and single loop operation. Also, for SLO there is an additional bias that needs to be considered in the drive flow to core flow relationship because of the effects of the reverse flow through the inactive jet pumps. However, this is not a noise source and is accounted for properly in the SLO APRM setpoint calculation. There is some additional uncertainty in power and flow measurements in SLO and that is accounted for in the licensing analysis through the use of approved uncertainties for the calculation of the SLMCPR (Reference 1) for both dual and single loop operation. These uncertainties are applicable to EPU operation because they reflect normal BWR noise and are given in terms of % of power.
Also, reactor power and flow noise during steady state operation is significantly reduced by BWR instrumentation because of the multiple measurements used both in power (multiple LPRM detectors) and flow (multiple jet pump DP transmitters). Additionally, noise filters are used in the instrumentation system to improve the measurement of average power and flow during steady state operation. The noise in the power signal is approximately proportional to the absolute power, but the noise at the APRM output at rated EPU power is approximately the same as at rated pre-EPU power because the APRM instruments read normalized power. The change in noise for EPU has no impact on the ability to perform normal plant maneuvers and surveillance tests. The effect of noise on instrumentation reliability during transient conditions is addressed in response to RAI SRXB-A- 55.

Attachment 10 BVY 05-072 Docket No. 50-271 Page 124 of 126

==Reference:==
: 1. Letter F. Akstulewicz (NRC) to G. A. Watford (GE), Acceptance for Referencing of Licensing Topical Reports NEDC-32601P, "Methodology and Uncertainties for Safety Limit MCPR Evaluations"; NEDC-32694P, "Power Distribution Uncertainties for Safety Limit MCPR Evaluation"; and Amendment 25 to NEDE-24011-P-A on Cycle-Specific Safety Limit MCPR (TAC NOS. M97490, M99069 and M97491), March 11, 1999 RAI SRXB-A-57 In the response to RAI SRXB-A-6, the licensee states "[t]he reactivity events are analyzed with the steady state tools and the results presented regarding steady-state methods in this response are directly applicable. There are some increases in power, which are significant but remain within the comparisons between the above plants for corresponding events." This RAI response does not provide sufficient detail.
a)   State the specific reactivity event being referred to (e.g., control rod drop accident, rod withdrawal error).
b)   State what steady state methods evaluation were described in the response to the RAI SRXB-A-6 response. The ((                                                       )) would not serve to demonstrate that impact of local reactivity event on the fuel enthalpy and performance. Revise the RAI response and provide an explicit discussion of the event.
Response to RAI SRXB-A-57 The two transient reactivity events are the control rod withdrawal error (RWE) and the mislocated bundle error (MLE).
RWE The RWE analysis is performed by first finding a limiting condition during the plant cycle with high rod worth conditions. The event represents the normal withdrawal of a control rod by the operator, however an error is made in the rod selection such that the operating fuel limits may be exceeded. The core simulator is then used to determine the change on MCPR for the condition with the maximum worth rod from a fully withdrawn condition. This MCPR, from the change in conditions at the limiting bundle plus a statistical adder to ensure it is bounding, is added to the safety limit MCPR (SLMCPR) to compare with the plant operating limit MCPR (OLMCPR) from other events to determine the most limiting value. In addition, the RWE analysis confirms compliance to the 1% plastic strain criteria. The type of hardware installed at the plant determines the type of rod withdrawal analysis that has to be performed. The RBM is not credited at VYNPS, so the highest worth rod is chosen from a fully withdrawn condition. The

Attachment 10 BVY 05-072 Docket No. 50-271 Page 125 of 126 BWR steady-state simulator is used for the analysis throughout the core cycle evaluation, determination of the highest worth rod and the MCPR analysis since this is a slow event where analyzing the neutronic and thermal-hydraulic conditions in equilibrium is appropriate. The conditions for this analysis lie within the range of the methods applications defined in the RAI SRXB-A-6 response. The following are the bases for the analysis:
: 1. The core is operating in a high worth nominal control rod pattern. The operating state is characterized by a control rod pattern, which is at high power and in an equilibrium xenon condition.
: 2. The RWE transient occurs over sufficient time to allow void re-distribution and heat transfer to reach equilibrium while leaving xenon concentrations and core inlet conditions unchanged.
: 3. The highest worth control rod when withdrawn will insert the maximum reactivity into the core and cause the greatest local bundle power response.
MLE This is not specifically a transient event in that it is due to a bundle being mislocated in the core.
The analysis is to determine the ?CPR for a bundle which is mislocated to a location, which would increase the power of such a bundle to a value greater than it would be in its planned location. Since core instrumentation does not monitor some fuel bundles, no analysis credit is taken for higher local power which may be detected by the instrumentation. The BWR simulator is applied for the fuel cycle and the specific bundle analysis. The selection process involves high reactivity worth bundles being placed in high power regions. The design limit calculated is the MCPR. The change in MCPR is added to the SLMCPR to determine the limiting value to compare with other plant transient events to provide the core operating limit MCPR.
RAI SRXB-A-58 For the transients, LOCA and ATWS, during the initial condition, the axial power peaking and distributions and the accuracy of the steady state neutronic methods affect the plants response.
State whether uncertainties are applied to the axial nodal powers, and the calculated void fraction. If uncertainties are not applied, please justify.
Response to RAI SRXB-A-58 All power distributions input to the transient process are based on nominal 3-D simulator calculations. As stated in the Alternate Approach section, the operating strategy and burn inputs are structured to obtain power distributions giving the maximum transient, or conservative

Attachment 10 BVY 05-072 Docket No. 50-271 Page 126 of 126 response. The hard bottom burn assumptions ensure that the control blade scram reactivity response is minimized, yielding conservative results.
In Section 5.6 of NEDC-24154P-A, (Supplement 1-Volume 4), it is stated that the ODYN comparison to TRACG and TRAC-BE1 demonstrates that it qualifies as a best estimate code for ATWS analysis, which is conservative in most cases. It is also stated in this section that for ATWS applications, prior regulatory approval has been granted for best-estimate code application based on the low probability of the event, conservatisms in key inputs and the acceptance criteria.
The axial power profile ((
                                                )) The average and hot bundle void profile is determined by SAFER at the initial and post-LOCA conditions. Uncertainties in predictions of void reactivity have no impact in the SAFER/GESTR methodology.

BVY 05-072 Docket No. 50-271 Attachment 11 Vermont Yankee Nuclear Power Station Proposed Technical Specification Change No. 263 - Supplement No. 30 Extended Power Uprate Response to Request for Additional Information Exhibits Exhibit SPSB-C-52-1 Exhibit SPSB-C-52-2 Exhibit SPSB-C-52-3 Exhibit SPSB-C-52-4 Exhibit SPSB-C-52-5 Exhibit SPSB-C-52-6 Exhibit EMEB-B-138-1 Exhibit EMEB-B-138-2 Exhibit EMEB-B-138-3 Exhibit SRXB-A-1 8-1

BVY 05-072 Docket No. 50-271 Exhibit SPSB-C-52-1 Vermont Yankee Nuclear Power Station Proposed Technical Specification Change No. 263 - Supplement No. 30 Extended Power Uprate Response to Request for Additional Information Calculation VYC-0886, Rev. 2 Total number of pages in this Exhibit (excludina this cover sheet) is 13. l

VY CALCULATION CHANGE NOTICE (CCN)
CCN Number:         04        Calculation Number:       VYC-0886          Rev. No.      2 Calculation

==Title:==
Station Blackout Documentation Analysis Initiating Document:       EPU VYDC/MMfrM/Spec. No.J other Safety Evaluation Numnber:         N/A Superseded Calculation:             N/A                          Superceded by        N/A Implementation Required: 0 Yes El No Computer Codes: N/A Reason for Change:
The VYC-0886 Rev 2 is updated to assess the effect of Extended Power Uprate (EPU) on this calculation.
Description of Change:
This CCN updates VYC-886 Rev 2, for EPU.
Technical Justification for Change:
See Attachment A

== Conclusions:==

The results of Reference I were addressed at EPU conditions. The effect of EPU on VIYC-886Rev2 are summarized in Attachment A.
Are there any open items in this CCN?       E Yes [ No Prepared By/Date             Interdiscipline Review By/Date     Independent Review By/Date      Approved By/Date 02/ill1ch                         jt7l             -           7        a           {nGo Liliane Schor        /lff/                                               S9 Alan L. Roberts aw Jaes G. Rogers Final Turnover to DCC (Section 2):
: 1)       All open items, if any, have been closed.
: 2)       Implementation Confirmation (Section 2.3.4)
QCalculation accurately reflects existing plant configuration, (confirmation method indicated below) 5 Walkdown          El As-Build input review      a Discussion with OR                                              (print name) 0 N/A, calculation does not reflect existing plant configuration
: 3) Resolution of documents identified in the Design Output Documents Section of VYAPF 0017.07 has been initiated as required (Section 2.3.6, 2.3.7)1 J_                                     _
Print Name                              Signature                           Date Total number of pages in package including all attachments:      13 pages Note: VYAPF 0017.07 should be included immediately following this form.                                   VYAPF 0017.08 AP 0017 Rev. 8 Page I of I

VY CALCULATION DATABASE INPUT FORM Place this form in the calculation package immediately following the Title page or CCN form.
VYC-0886/CCN04                        2                                N/A                                        N/A VY Calculation/CCN Number          Revision Number                 Vendor Calculation Number                    Revision Number Vendor Name:               N/A                          PO Number:               N/A Originating Department: DesignEngineering Critical References Impacted: E UFSAR Z DBD El Reload. "Check" the appropriate box if any critical document is identified in the tables below.
EMPAC Asset/Equipment ID Number(s): N/A EMPAC Asset/System ID Number(s): N/A Keywords: Decay Heat. SBO. Torus Temperature. Condensate Storage Tank. Ventilation For Revision/CCN only: Are deletions to General References, Design Input Documents or Design Output Documents required? a Yest Z No Design Input Documents and General References - The following documents provide design input or supporting information to this calculation. (Refer to Appendix A, sections 3.2.7 and section 4)
Significant                Critical Difference    Affected   Reference
* Reference #      ** DOC #     REV #                        ***Document Title (including Date, if applicable)                  Review tt     Program        (/)
1          VYC-0886        2        Station Blackout Documentation Analysis, 01/03/2001 2         TE 2003-064               Station Blackout PUSAR Input                                                                    l 3         GE-VYNPS-       N/A       Letter, Michael Dick (GE) to Craig Nichols (ENOI), VYNPS EPU Task T0400: Decay Heat AEP-146                 for Containment Analysis dated March 10,2003.
4          VYC-2282        0        Vessel Stored Energy wth GE14 Fuel at 20% Power Uprate, dated 5/9/03 5         NUMARC 87-      N/A       NUMARC 87-00, dated 11/20/87, including NRC accepted errata and 0 &A's from 00                   MUMARC seminars and Topical Report F.
6          NVY 91-98      N/A       Letter, USNRC to VYNPC, NVY 91-98, 'Vermont Yankee Station Blackout Analysis,'
June 5, 1991 7                         N/A       ASME Steam Tables 8                         N/A       VY Technical Specification 9          VYC-2270        0        VY GE 14 Appendix Rat 20% Power Uprate, dated 05/09/2003 10           VYC-415         0       Appendix RIRCIC, HPCI &ECCS Room Cooling, dated 4/29/1986 11           VYC-415         0       Appendix R/RCIC, HPCI &ECCS Room Cooling, dated 9/04/2002 CCN 01 VYAPF0017.07 AP 0017 Rev. 8 Page 1 of 4

Significant              Critical Difference  Affected    Reference
* Reference # **  DOC # REV #                  ***Document Title (including Date, if applicable)              Review tt   Program        (1) 12       VYC-886    2   Station Blackout Documentation Analysis, dated 9104/2002 CCN3 13       VYC-2279   0   Evaluation of EPU Impact on Ambient Space Temperatures During Normal Operation, dated 04/1112003 14       VYC-1347   0   Main Steam Tunnel Heatup Calculation, dated 11/1/96 15        OT-3122   19  Loss of Normal Power, 04/18/00 16      VYC-1628D   0   Torus Temperature Response to Appendix Rand Station Blackout Scenarios, dated CCN02 VYC-0886 Rev 2 CCN04, Page 3 of 7 VYAPF0017.07 AP 0017 Rev. 8 Page 2 of 4

VY CALCULATION DATABASE INPUT FORM (Continued)
Design Output Documents - This calculation provides output to the following documents. (Refer to Appendix A, section 5) tttCritical
                                                                                                                         **** Affected   Reference
* Reference #    ** DOC #       REV #                          Document Title (including Date, if applicable)              Program         (1)
VYC-1432         4       VY Vessel Level for Appendix R, 05/17/1996 VYC-1458         0       Station Blackout Load Capacity Analysis, 10/15/1996
_               VYC-1628         0       Torus Temperature and Pressure Response to Large Break LOCA and MSLB Accident CCN3      Scenarios, 3/21/2002 VYC-1628D         0       Torus Temperature Response to Appendix R and Station Blackout Scenarios-dated November 5, 1998.
VYC-2159         0       VY-Cycle Independent Decay Heat-Comparison Between ORIGEN-2 and ANSI/ANS 5.1-1979 Standard, 2/27/2001 VYC-2314          0      Minimum Containment Overpressure for Non-LOCA Events, 9/03/2003 DBD SADBD                                                                     DBD .
DBD CPS                                                                       DBD/
DBD HPCI                                                                      DBD DBD HVAC                                                                      DBD                 V DBD MS                                                                        DBD DBD NBVI                                                                      DBD                  /
DBD RCIC                                                                      DBD                 V DBD RIR                                                                       DBD SSCA Vol I                                                                    Appendix R VYC-0886 Rev 2 CCN04, Page 4 of 7 VYAPF00 17.07 AP 0017 Rev. 8 Page 3 of 4 LPC# 6

VY CALCULATION DATABASE INPUT FORM (Continued)
* Reference # -         Assigned by preparer to identify the reference in the body of the calculation.
**Doc#-                 Identifying number on the document, if any (e.g., 5920-0264, G191172, VYC-1286)
*** Document Title -    List the specific documentation in this column. "See attached list" is not acceptable. Design Input/Output Documents should identify the specific design input document used in the calculation or the specific document affected by the calculation and not simply reference the document (e.g., VYDC, MM that the calculation was written to support.
**** Affected Program - List the affected program or the program that reference is related to or part of.
t                       If "yes," attach a copy of"VY Calculation Data" marked-up to reflect deletion (See Section 3.1.8 for Revision and 5.2.3.18 for CCNs).
tt                      If the listed input is a calculation listed in the calculation database that is not a calculation of record (see definition), place a check mark in this space to indicate completion of the required significant difference review. (see Appendix A, section 4.1.4.4.3). Otherwise, enter "N/A."
ti.t                    If the reference is UFSAR, DBD or Reload (IASD or OPL), check Critical Reference column and check UFSAR, DBD or Reload, as appropriate, on this form (above).
Note: All calculations in the Design Output list were reviewed. No revision required.
Other Design Output were reviewed. The following revisions are required:
: 1) DBDs referencing VYC-886 Rev2 need to be addressed.
: 2) Calculation VYC-1347 should be addressed for EPU.
VYC-0886 Rev 2 CCN04, Page 5 of 7 VYAP0017.07 AP 0017 Rev. 8 Page 4 of 4 LPC#6

Page    6    of 15 VY CALCULATION REVIEW FORM Calculation Number:      VYC-0886          Revision Number:         2             CCN Number:              04

==Title:==
-Station Blackout Documentation Analysis Reviewer Assigned:.         Alan Robertshaw                                               Required Date:    February 2004 l:] Interdiscipline Review        `7 Independent Review Comments*                                                                         Resolution
-1. Assumptions on page I of Att. A need Reference.                               1. Added
: 2. Need Reference for Table I in Att. A.                                         2. Done
: 3. On page 5 of Att. A please state the TS CST Inventorv.                       3. Added TS CST inventory OA
          *ReviewerSignature Q               abbe Date /
avatf                           ,~-~-
Calculation Preparer (Comments Resolved)
                                                                                                                                 /}oVA Date
                                                                                                                                                 /zco     -
I-.'    -
Mcthod of Review:            l Calculation/Analysis Review                                            ultin/ ay El Alternative  Calculation El Qualification Testing                           Reviewer Signature (Comments Resolved)                     Date /
*Commrents shall be specific, not general. Do not list questions or suggestions unless suggesting wording to ensure the correct interpretation of issues.
Questions should be asked of the preparer directly.
VYC-0886 Rev 2 CCN04, Page 6 of 7 VYAPF0017.04 AP 0017 Rev. 8 Page 1 of I

Calculation VYC-0886 Rev2 CCN04                                                 page 7 of 13 Page     7   of 7 VY CALCULATION OPEN iTEM LIST Calculation Number:     VYC-886                  Revision Number:    2                     CCN Number:     4 Open Item                                        Resolution                                Method of 01 Tracking or Date Closed DEDs referencing VYC-886 Rev2 need to be verified for changes to torus temperatures.
VYC-1347 needs to be CCN for EPU VYC-0886 Rev 2 CCN04, Page 7 of 7 VYAPF0017.05 AP 0017 Rev. 8 Page 1 of 1

Calculation VYC-0886 Rev 2 CCN 04                                                        Page 1 of 6 Attachment A Reason for Revision Revision 2 of VYC-0886 is updated to incorporate the EPU changes This CCN incorporates:
: 1) Condensate Inventory Requirements at EPU incorporating:
                  - The decay heat at EPU from Reference 3.
                  - Vessel stored energy at EPU from Reference 4 (VYC-2282).
: 2) Loss of ventilation
: 3) Torus Temperature Assumptions (same as in reference 1)
: 1. No off-site power available (SBO)
: 2. The reactor depressurizes from 1095 psia to 100 psia during the SBO scenario. The 1095 psia is assumed to be an average SRV setpoint. The 100 psia is assumed a low pressure setpoint where RHR system is deployed for shutdown cooling.
: 3. It is assumed that at about 8 hours, the vessel pressure decreased to about 100 psia.
: 4. It is assumed that at 100 psia the fluid and solids in the reactor vessel are at the same temperature.
This is a reasonable assumption, since at 8 hours, most of the metal in the vessel will be at the fluid saturation temperature.
Condensate Inventory Reouirements The inventory required for decay heat removal will be calculated using a formula given in NUMARC 87-00 and also using the decay heat calculated in Reference 3.
Condensate Inventory to Remove Decay Heat From Reference 5 V = 35.55 galIMWt = 35.55
* 1912
* 1.02 = 69331 gallons for 8 hours From the decay heat calculation Q decay at 8 hours (interpolated in the integrated decay heat table - Table 1, Reference 3, next page) 20000 4.68951E+08 28800 6.11079E+08 40000 7.91978E+08

Calculation VYC-0886 Rev 2 CCN 04                                          Page 2 of 6 Attachinent A Table 1 - Integrated Decay Heat for 20% power Uprate Time (sec)             GE 2 sigma P/Po  Integrated     Integrated Kwsec Integrated, BTU 0.00000            1.00000           0.00000          0.00000         0.00000 0.10000            0.99210           0.09961       1.94254E+05     1.89980E+05 0.15000           0.96250            0.14847       2.89552E+05     2.83182E+05 0.20000           0.93280            0.19585       3.81959E+05     3.73556E+05 0.40000           0.74710            0.36384       7.09580E+05     6.93969E+05 0.60000           0.59080            0.49763       9.70503E+05     9.49152E+05 0.80000           0.49380            0.60609       1.18203E+06     1.15602E+06 1.00000          0.33880            0.68935       1.34440E+06     1.31483E+06 2.00000           0.15480            0.93615       1.82572E+06     1.78556E+06 4.00000           0.06073            1.15168       2.24606E+06     2.19664E+06 10.00000          0.05234            1.49089       2.90760E+06     2.84363E+06 20.00000           0.04546           1.97989       3.86127E+06     3.77632E+06 40.00000          0.03986            2.83309       5.52521E+06     5.40366E+06 60.00000          0.03687            3.60039       7.02163E+06     6.86715E+06 80.00000          0.03466            4.31569       8.41664E+06     8.23147E+06 100.00000          0.03321            4.99439       9.74026E+06     9.52598E+06 150.00000          0.03073            6.59289       1.28577E+07     1.25749E+07 200.00000           0.02909           8.08839       1.57743E+07     1.54273E+07 400.00000          0.02550            13.54739      2.64207E+07     2.58394E+07 600.00000          0.02346            18.44339      3.59690E+07     3.51777E+07 800.00000          0.02197           22.98639       4.48290E+07     4.38427E+07 1000.00000          0.02079           27.26239       5.31682E+07     5.19985E+07 1500.00000         0.01861           37.11239       7.2378 1E+07    7.07858E+07 2000.00000          0.01707           46.03239       8.97742E+07     8.77992E+07 4000.00000           0.01370          76.80239       1.49783E+08     1.46488E+08 6000.00000           0.01209          102.59239      2.00080E+08     1.95678E+08 8000.00000           0.01114          125.82239      2.45384E+08     2.39985E+08 10000.00000          0.01047          147.43239      2.87529E+08     2.81203E+08 15000.00000          0.00986          198.25739      3.86649E+08     3.78143E+08 20000.00000          0.00918           245.86739      4.79500E+08     4.68951E+08 40000.00000           0.00775          415.22739      8.09793E+08     7.91978E+08 60000.00000           0.00699          562.63739      1.09728E+09     1.07314E+09 80000.00000          0.00647          697.19739      1.35970E+09      1.32979E+09 86400.00000          0.00633          738.15099      1.43957E+09      1.40790E+09 100000.00000          0.00608          822.54579      1.60416E+09      1.56887E+09 150000.00000          0.00539         1109.27079      2.16334E+09     2.11575E+09 172800.00000          0.00515         1229.46099      2.39774E+09     2.34499E+09 200000.00000          0.00492         1366.41299      2.66483E+09     2.60621E+09 259200.00000          0.00451         1645.27459      3.20868E+09     3.13809E+09 345600.00000          0.00406         2015.28259      3.93028E+09     3.84382E+09 400000.00000          0.00384         2230.16259      4.34935E+09     4.25367E+09 432000.00000          0.00373         2351.26659      4.58553E+09     4.48465E+09 600000.00000          0.00327         2939.18259      5.7321 lE+09    5.60601E+09 800000.00000          0.00290         3556.48259      6.93599E+09     6.78340E+09 864000.00000          0.00281         3739.39459      7.29272E+09     7.13228E+09 1000000.00000          0.00265         4110.81059      8.01707E+09     7.84069E+09

               . .. .. - .. Z @ . @ . "
Calculation VYC-0886 Rev 2 CCN 04                                                  Page 3of 6 Attachment A Q = M (h(g) - h(l)) to calculate the inventory requirement Where:
h(g) (Reference 1) = 1187 Btu/lbm (average between 1095 and 100 psia) [see note on page 22 of Reference 1) h(l) = 118 Btullbm (150 'F conservative temperature of CST) v(l) = 0.01634 ft3/lbm @ 150'F All properties are from Reference 7.
6.11079 E8
* 0.01634
* 7.48
                                                        = 69867 gal (1187 -118)
This inventory matches very well the NUMARC formula and it will be used.
Therefore the inventory requirement for 8 hours of decay heat is 69867 gallons.
Condensate needed to remove the vessel stored energy to depressurize from 1095 psia to 100 psia.
Stored Energy in Fluid The fluid energy at full power (EPU conditions is) (Reference 4)
Fluid (EPU, t=0.0)                     Mass              Enthalpy         Total Energy (Ibm)             (Btu/lbm)             (BTU)
Liquid                        386,971              525.54           2.03369E8 Steam                       13,186.12             1191.05          0.15705E8 Total                                                             2.19074E8 The fluid energy at 100 psia is not changed from Reference 4. The level will be the same after depressurization for current licensed power (CLP) as for EPU. Hence the volumes of steam and liquid will be the same, as well as the enthalpy.
Fluid (depressurized at                      Mass            Enthalpy      Total Energy 100 psia)                          (lbm)           (Btuflbm)          (BTU)
Liquid                         510322.4             298.4          1.5228E8 Steam                          703.13              1187.2          0.008E8 Total                                                             1.5311E8 Thus, the difference in fluid energy:        AEflUid = 2.19074e8-1.5311E8 = 0.65964E8 Btu

Calculation YYC-0886 Rev 2 CCN 04                                               Page 4 of 6 Attachment A Stored Energy in Solid (From Reference 4)
Solid                  Total Solid Energy          Heat Conductor Effective (EPU, time= 0.0)                    (BTU)                     Temperature (&deg;F)
Liquid Exposed                 0.9604399E8                          601.24 Steam Exposed                  0.25155507E8                         609.18 Total                     1.211995E8                          602.83 Q = MCp AT = MCp (602.83 - 32) = MCp 570.83 MCp = 1.211995E8 /570.83 = 212321.53 Btu /OF Tsat @ P= 100 psia = 328&deg;F (ASME Steam Tables- Reference 7)
At 100 psia:
Q= 212321.53 * (328-32) = 0.628472E8 Btu Total Energy removed from structures: AE stuctures 1.211995E8 - 0.628472E8 =
0.58352E8 Btu Total energy removed from the vessel during depressurization =
AEfluid + AE stuctms =    0.65964E8    + 0.58352E8 = 1.24316E8 Btu The inventory needed to remove this heat =
V = 1.24316E8 * (.01634)
* 7.48
                                            = 14,214gallons (1187-118)
Hence, the total inventory requirements = 69867 gallons +14,214 gallons = 84081 gallons

Calculation VYC-0886 Rev 2 CCN 04                                                     Page 5 of 6 Attachment A For the CLP (Reference 1) the total CST inventory requirements for removing the decay heat and vessel stored energy = 75,837 gallons.
The TS (Reference 8) CST inventory of 75000 gallons is exceeded at both EPU and CLP for the 8 hours coping duration.                                                        /
The HPCI'RCIC taking suction from CST has to make up for the vessel leakage (TS allowable and pump seal leakage of 61 gpm - Reference 6, page 17 of TER, also used in both Appendix R analysis (Reference 9, VYC-2270) and in Reference 1. The leakage amount does not change for EPU.
The needed CST inventory to account for leakage; V= 61 gpm
* 8 hours *60 min/hour = 29, 280 gallons.
This inventory, added to that already calculated for decay heat and depressurization would total:
V = 29280 + 84081 = 113361 gallons, which would normally be available from CST but, if not could easily be made available from the torus. Therefore, the conclusions of VYC-886 Rev2 that the Technical Specifications CST inventory requirement of 75000 gallons is not adequate for an 8 hour duration is valid at EPU. However, with Alternate AC (Vernon Tie) and low pressure systems available, sufficient inventory is available from the torus for the required 8 hours.
The power uprate results in a need for more inventory, 113361 gallons versus 105117 gallons at CLP.
Reactor Coolant Inventorv The depletion of the available inventory in CST will not jeopardize reactor coolant inventory because makeup inventory can be provided from the torus. When the torus temperature exceeds 140 OF and RCIC and HPCI can not be used with suction from the torus, reactor inventory can be provided from the torus via low pressure pumps. Since VY is an Alternate AC plant, crediting use of the low pressure pumps is acceptable. This conclusion is unaffected by power uprate.
Loss of Ventilation The heat-up due to the loss of ventilation due to an SBO event for RCIC Room, HPCI Room, Main Steam Tunnel, Control Room, Switchgear Room, and Intake Structure is addressed in Reference 1.
RCIC Room The heat-up calculation is based on VYC415 RevO (Reference 10 modified by CCN 1 (Reference 11)). The heatup is based on the piping temperature, RCIC turbine and Switch Heat Loss. The RCIC Room Temperature calculated in Reference 12 is unaffected by EPU.

Calculation VYC-0886 Rev 2 CCN 04                                                   Page 6 of 6 Attachment A HPCI Room.
The heat-up calculation is based on heat loads from VYC-0415 Rev 0 (Reference 10). The heat loads are from the piping and the HPCI turbine. The heat loads are unaffected by power uprate. Therefore the calculation for HPCI room heat-up is not affected by EPU.
Main Steam Tunnel The issue is isolation of HPCI and RCIC on high steam tunnel temperature.
Reference 13 calculated an increase of 0.6 0 F in the normal temperature of the steam tunnel, at EPU.
The conclusion of VYC-886 Rev 2, that the main steam tunnel heat-up is slow on loss of ventilation and the reactor will already be in the process of cool-down, is valid at EPU.
Furthermore, Reference 14 (VYC-1347) concluded that the heat-up in the main steam tunnel is less than that required to isolate HPCI and RCIC. For the case when the feedwater and main steam isolates (SBO conditions), the peak room temperature from Reference 14 is 174 0 F (isolation temperature assuming loop accuracy) at approximately 18 hours. Based on the results of Reference 13, the change in Main Steam Tunnel Heatup will be very small at EPU. Furthermore, procedure OT-3122 (Reference 15) limits HPCI & RCIC operation to 2 hours; hence the reactor pressure after 2 hours should be low enough to permit operation of the Low Pressure Pumps (CS and RHR).
Therefore the impact of power uprate on heatup of the Main Steam Tunnel is negligible. It is recommended that calculation VYC-1347 be updated for EPU conditions.
Control Room Restoration of ventilation in the Control Room is governed by Procedure OT-3122 and is unaffected by Power Uprate. Control Room Heatup for loss of ventilation is unaffected by power uprate.
Switchacar Room The heat loads in the switchgear room are unaffected by the power uprate.
Intake Structure The heatup of the Intake structure on loss of ventilation with only 2 Service Water available is unaffected by the power uprate since the heat loads in the intake structure are unaffected by power uprate.
Torus Temperature The Torus Temperature calculation for SBO at power uprate was performed in Reference 16 (VYC-1628D CCNO2). The peak suppression pool temperature is 187.9 'F.

BVY 05-072 Docket No. 50-271 Exhibit SPSB-C-52-2 Vermont Yankee Nuclear Power Station Proposed Technical Specification Change No. 263 - Supplement No. 30 Extended Power Uprate Response to Request for Additional Information Calculation VYC-1347, Rev.0 Total number of pages in this Exhibit (excludina this cover sheet) is 29.

t
                       / 17 7,,4f x
ORIGINAL,     PAGE 1 of  / 17 PAGES Rev. 1:   PAGE 1 of       PAGES
* Rev. 2:   PAGE 1 of       PAGES Rev. 3:   PAGE 1 of       PAGES OA RECORD?                                                             IMS NO. M02.01O05 I/  YES                                                        RECORD TYPE     09.C16.004 NO                                                          W.OJP.O. NO. 4055 YANKEE NUCLEAR SERVICES DIVISION CALCULATIONIANALYSIS FOR TITLE   Maln steam Tunnel Heatup Calculation PLANT   Vermont Yankee                                     CYCLE   _ NIA CALCULATION NUMBER            VYC-1 347 KEYWORDS      GOTHIC: Room: Heat-up: RRU G         htiUCLEAR SERVICES DIVISION YANKEEATOMIC ELECTRIC COMPANY 580 MAIN STREET BOLTON, MASSACHUSETTS 01 740

YANKEE NUCLEAR SERVICES DIVISION                    CALC. NO. VYG1347                         REV.        _       DATE 11--9 TITLE Main Steam Tunnel Heatup Calculation PREPARED BY     Jim Pappas               REVIEWED BY                               -                 PAGE           2       OF TABLE OF CONTENTS Section                                                Description                                                              Pane LIST OF TABLES..                                                                                                            3 LIST OF FIGURES..                                                                                                           4 1.0    PROBLEM DESCRIPTION.                                                                                                        5 1.1   Objective.                                                                                                            5 1.2   Method of Solution ....................................... ;                                                          6 1.3. Design Inputs.                                                                                                        6 1.4   Assumptions...................................................................................................... 9 2.0    PROBLEM ANALYSIS                              ..                                                                           10 2.1   GOTHIC Model Input ...................                                                                               10 2.1.1   Control Volumes .10 2.1.2   Thermal Conductors .11

====2.1.3 Heaters====
Main Steam Isolation Valves ..................................                                :      11 2.2   GOTHIC Runs                          ..                                                                              20 2.2.1   Run MST .20 2.2.2   Run MST2 .24

==3.0    CONCLUSION==
                                     .                   .                   .29

==4.0     REFERENCES==
 .........                                                                                                      30 Appendices A          GOTHIC Run MSTI .35 B          GOTHIC Run MST2 .81 C          Computer Code Evaluation ....................                                                            112 D          Reviewers Comments .114

1111111-
                            -- -                                                                                          -- I YANKEE NUCLEAR SERVICES DIVISION              CAMC. NO. VYrC-1347                REV.     -           DATE -1.-19 TITLE MaTn Steam Tunnel Heatup Calculation PREPARED BY Jim Pappas                REVI IEWED BY                                      PAGE -O3              OF_
LIST OF TABLES Table No.                                      Description                                                           Pane 1.3-1             Thermo-Physical Properties..................................................................... 7 1.3-2             Outer Environment Temperatures ...........................................................          a 2.1-la            MvST2 Dimensional Data..
MST2'i DmninlDt............................      ...................................
2.1-1b.           West Wall Dimensional Data........................... ...... ............................          16 2.1-2a            Piping Design Data................    .......................................................... 17 2.1-2b            Piping Heat Transfer Surface Areas..........................................................       18

w
* n:  N   o {
* I YANKEE NUCLEAR SERVICES DIVISION                CAMC. NO. VYC-G1347 REV.       DATE  11-1-96 TITLE Main Steam Tunnel Heatup Calculation PP.PARED BY    Jim Pappas            REVIEWED BY,                        PAGE    4  OF LIST OF FIGURES Figure No.                                       Descriotion                               Pape 2.1.3-1            MSIV Valve Outline   Drawing .19 2.2-la             MST1 Schematic                                                          22 2.2-lb             MST1 Temperature     Profile .23 2.2-2a             MST2 Schematic .                                                        26 2.2-2b             MST2 Temperature     Profile to Seven Days .27 2.2-2c             MST2 Temperature     Profile to Four Hours .28 A-1                MST1  Schematic.                                                        38 A-2                MST1 Input Tables .39 A-3                MST1  Graphical Results.55 A-4                MST1  Output Verifying the MSIV Model .70 A-S                MST1  Output Showing Condensation Heat Transfer Fluctuations.           73 A-1               MST2 Schematic .82 A-2               MST2 Input Tables .............                                          83
. .I    A-3               MST2 Graphical Results .98 1

YANKEE NUCLEAR SERVICES DIVISION                CALC. NO.      Y   LQ:1347 REV.           DATE 11-1-96 TITLE Main Steam Tunnel Heatup Calculation PREPARED BY Am Pannas                     REVIEWED BY                            PAGE      5    OF 1.0     PROBLEM DESCRIPTON This calculation determines the temperature rise in the main steam tunnel with a loss of HVAC to document whether this rise will result in an automatic isolation of the HPCI and RCIC systems. The HPCI/RCIC excess air temperature switches, within the main steam tunnel, provide automatic 8; isolation of the HPC1 and RCIC lines if a temperature of 185'F i 5*F is sustained for longer than 30 minutes (References' 23 and 24). The loop accuracy is 6F (Reference 36). Therefore HPCI and RCIC isolation can occur at a steam tunnel temperature as low as 174'F. This high temperature isolation scheme' is for line break protection and it is not intended for non-line break events, such as loss of main steam' tunnel cooling under loss of normal power.
Normal ventilation in the main steam tunnel is supplied by the Reactor Building Ventilation System and by RRUs 17A and 17B, located in the tunnel. A Reactor Transfer Fan (RTF-IA/ IB) exhausts air from the main steam' tunnel at approximately 4200 cfm. Each fan has a total caprcity of 14,400 cfm and takes inlet air from'various locations in the Reactor Building. The RRUs circulate and cool air inside the main steam tunnel. The fan capacity of each RRU is 5000 cfm and service water, supplied to coils within the RRUs, provides the cooling.
RTF-lA and both RRUs are powered from 480v MCC 6A. This MCC is NNS and is supplied from 4160v Bus 1. RTF-I1 is powered from 480v MCC 7A, which is also NNS, and is supplied from 4160v Bus 2.
Controls are located on the Turbine Building HVAC control panel, with auxiliary indications on the Control Room 9-25 panel. Typically, one fan and both RRUs are operating, with the second fan in stand-by. The operating fan and the RRUs maintain the main steam tunnel environment temperature at a yearly average of 125F, as described in the Vermont Yankee Environmental Qualification Program (Reference 25). A review of temperature data for the main steam tunnel indicates that the air temperature can peak at about 150F during the summer months (Reference 30).
1.1    Oblective The objective of this calculation is to determine the temperature rise of the air in the main steam tunnel, during a loss of normal power and under the following conditions:
* Loss of HVAC In the main steam tunnel.
             . Summer peak temperature for initial and boundary conditions.
* Main steam and feedwater lines in the tunnel are:
                     -   Isolated (steam and water are not flowing, GOTHIC run MSTI).
                     -   Not isolated (steam and water are flowing, GOTHIC run MST21.

YANKEE NUCLEAR SERVICES DMSION                 CALC. NO. WC-1347          REV.        DATE 11-1-96 TITLE Main Steam Tunnel Heatup Calculation PREPARED BY       Jim1Paoas             REVIEWED BY                            PAGE    6    OF 1.2     Method of Solution A lumped parameter GOTHIC model of the main steam tunnel is used to calculate the air temperiture rise on loss of HVAC. GOTHICl33A34) is a general purpose thermal-hydraulic computer program for design, licensing3 safety, and operating analysis of nuclear power plant containment and other confinement buildings. See Appendix C for verification of the GOTHIC version used.
The models consist of volumes, flow pcths, & thermal conductors arranged and connected to represent the thermal-hydraulic response of the main steam tunnel. The thermal mass of each conductor is included in the GOTHIC computation.
1.3    Design Inputs 1.3.1   The thermo-physical properties for the materials used are shown in Table 1.3-1..
1.3.2 The boundary temperatures for spaces surrounding the main steam tunnel are shown in Table 1.3-2.

YANKEE NUCLEAR SERVICES DIVISION                  CAMC. NO. .Y-1347 -    REV. -   DATE       i-.
TITLE Maln Steam Tunnel Heatup Calculation PREAE B.       m , A,                 R.
PREPARED BY. ,Jim Pappas                  REMIEWED BY                        PAGE  7      OF__
:                                                                                    . _~
Tab. 1 .31 Thermo-Physical Properties
                   ..                        .W
* 0    .
ConcreteP1 ')'         75                       1.05               142            0.156 Steel Plpel'.          68                      25.0                487             0.11
:1000                     0.073 900.                    0.067 800                     0.060 700                     0.055 Insulation0            600                     0.050        15O                   o.io&#xa2;'
500                     0.046.
400                     0.042
                        -300                     0.038 200                     0.036 100                     0.033 a) Superscript numbers refer to References In Section 4.0.
b) Assumed.

YANKEE NUCLEAR SERVICES DIVISION              CALC. NO. MCf1347 .         REV.        DATE 11 TITLE -:-n t      TUnnel Heatup Calculation PREPARED BY    Jim Paopas              REVI ,EWED BY                    -      PAGE    8     OF Table 1.3-2 Outer Environment Temperatures
         .. r        M     aK-                             -Z'AA           _    ir Normal            150"F              125"F             100&deg;F             83eF0" Operation           160F               125F              100&#xb6;              83F&deg; Peak Summer            1604F             1300F             1200 F            90OF~b Notes a) Vermont Yankee FSAR (Reference 32), Table 2.3.2, Highest mean daily maxirr Jm for'summer months.
b) Vermont Yankee FSAR (Reference 32), SecUon 10.12, Summer design temperature.

I      YANKEE NUCLEAR SERVICES DIVISION TITLE Main Steam Tunnel Heatul Calculation CALC. NO.       C13347      REV.        DATE _1:i PREPARED BY Jim Pappas                    REVIEWED BY                              PAGE    9     OF 1.4     Assumptions The critical assumptions used in the GOTHIC models are as follows:
1.4.1   Initial main steam tunnel air temperature is 1507F. This is based on Reference 30 and is considered conservative. Reference 30, describes that this value is derived from a temperature element that is close to hot process lines. Therefore, the corresponding ambient room temperature should be lower.
1.4.2   Initial main steam thermodynamic statepoint is saturated steam at 985 psia based on the heat balance shown on Figure 1.6-1 in the PSAR (Reference 32). Therefore, the temperature is 543'F.
1.4.3   Initial feedwater thermodynamic statepoint is saturated water at 373F based on the heat balance shown on Figure 1.6-1 in the FSAR (Reference 32). Therefore, the pressure is 179.8 psia.
1.4.4   HPCI and RCIC turbine steam supply temperatures are 543'F.
1.4.5   Both RRUs are inoperative for the analysis.
1.4.6 The air temperatures in the spaces surrounding the main steam tunnel are listed in Table 1.3-2 and are assumed to be constant throughout the transients.
1.4.7   In model MST1, where the main steam lines are isolated and the feedwater pumps are off, the four main steam lines (MS-1A through D) and the feedwater lines (FDW-14/15/16/17) dissipate the heat in the line volume, cooling down as they do so.
All other lines contain fluid at their respective constant temperatures, as listed in Table 2.1-2a.
1.4.8    In m6del MST2, where the main steam lines do not isolate, all lines contain fluid at their respective constant temperatures, as listed in Table 2.1-2a.
1.4.9    Miscellaneous piping, steel, and equipment are left out of the models.
1.4.10 The floor is left out of the models to add conservatism to the room heat-up.
I .*
1.4.11 The west wall contains a metal section through which the main steam lines pass and which two blowout panels are installed. This metal section is modeled in the GOTHIC runs. However, other non-concrete wrall sections are not. They include:
a) ventilation duct with blowout damper in the north wall, b) a blowout ranel and a blowout door in the south wall, and c) various duct work and pipe sleeves.
This is assumed to be conservative since it inhibits natural circulation that would normally exist in the room.

M                        --i YANKEE NUCLEAR SERVICES DIVISION                  CALC. NO. Wrc1347          REV.     -   DATE 11-1-96 TITLE Main Steam Tunnel Heatup CaDculation PREPARED BY       Jim Pan               REV,IEWED BY        _I                    PAGE     10    OF 2.0     PROBLEM ANALYSIS 2.1     GOTHIC Model Input The following sections describe the major input that was calculated for the GOTHIC models.
2.1.1    Control Volumes Main Steam Tunnel The main steam tunnel is Volume 1 in the models. The relevant dimensional data of the main steam tunnel for construction of the models are shown in Table 2.1-1. From these data, the wall surface areas and the room overall volume are obtained. The volume of the steam tunnel is:
VuST   = (North Wall) x (East Wall) x (Height)
VMST   = 36.25 x 24' x 25.5 = 22,185 1t5 The hydraulic diameter is:
Dh = 4A PI where A is the cross sectional area of the volume (i.e., the ceiling or floor area) and P, is the wetted perimeter. P,, is defined by GOTHIC as S/h or the surface area of all structures divided by the height of the volume. S would, therefore, be the total surface area of all the walls and the ceiling. The floor is not modeled.
4Ah Dh        s D             _-.=       '"
bAnoh'+AM         + Aeut + Aest + Aceng 4(870X25.5)
Dh     97438 + 924.38 + 612 + 612 + 870 Dt = 225 ft In the run where the main steam and feedwater lines are isolated, those lines are modeled as separate control volumes. The four main steam lines are lumped into one volume as are the two feedwater

                                                                                              .. I,      -, .
o YANKEE NUCLEAR SERVICES DMSION                     CAL. NO. VYQ=1347       REV.      DATE         11-1-96 TITLE Main Steam Tunnel HeatUp Calculation PREPARED BY       Jim Pangas             REVIEWED BY                              PAGE  I11 OF lines. The volumes are obtained from the pipe data in Table 2.1-2b, and follow. The hydraulic diameters are simply the pipe diameters.
Main Steam Lines VMSL       4 X       )2 x L VMSL     4 x n(8062 I12)2 x 46.5 VMSL = 263.75 f 3 Feedwater Lines VFCIW  2 a    R) 2 14 116 +  115,17)
VFDw   =  it(6.781 1 12)2 (45.7 + 47.9)
VFDW = 93.90 ft 2.1.2    Thermal Conductors The input for the thermal conductors that represent the steam tunnel walls and ceiling is taken from Table 2.1-1. The floor is left out of the model to add conservatism. Typically there is little heat transfer through the floor of a heated room.
The thermally significant piping found in the main steam tunnel are described in Table 2.1-2. The GOTHIC input for these conductors is also shown in the table.

====2.1.3 Heaters====
Main Steam Isolation Valves The main steam isolation valves (MSIV) in the steam tunnel have a substantial amount of un-insulated structure that makes up the yoke and actuator. Figure 2.1.3-1 (Reference 39) shows the outline of the valve. Heat will conduct through and out of this structure into the main steam tunnel. There are four such valves in the tunnel.
The yoke of each MSIV consists of four 3' solid rods attached to the bonnet (Reference 40). The yoke acts as a support for the actuator and ns a spring guide. Through the center of the yoke, the valve stem travels.

YANKEE NUCLEAR SERVICES DIVISION                      CALC. NO. YC1A347   REV.        DATE .11J-9 TITLE Main Steam Tunnel Heatup Calculation PREPARED BY        Jim Pappas                 REVIEWED BY                          PAGE    12   OF The method of modeing the heat transferred to the main steam tunnel by this assembly will be to treat each 3" yoke rod and the stem as a fin. The stem will be modeled as though it were a fifth yoke rod.
Therefore, each MSIV will be modeled as having five 3" solid rod fins heated at one end. The heated end is that attached to the body of the valve. It is assumed that the actuator is far enough away from the valve body that any heat conducted to it is negligible.
GOTHIC cannot model this situation because it involves two-dimensional conductive heat transfer.
GOTHIC can only model one-dimensional conduction. Therefore, a formulation of the heat rate provided by the yokes will be derived here and input into GOTHIC as a theater".
The general equation for such a fin is (Reference 29):
q =    Jhk(r,    - T,.)tanh(mL) where:           q = heat rate (Btulhr) h = convective heat transfer coefficient (Btulhrft 2 F)
P = perimeter of the fin (7td ) (ft) d = diameter of the rod (ft) k c thermal conductivity of the rod material (Btu/hrftt.F)
A c cross sectional area of the fin (*1)
T. temperature of the heated end ("F)
T    = ambient room temperature ("F) m= W (ft')
L=length of the fin (ft)
The values in the following table will be used. The value for h is taken from Reference 38 and is considered to be conservative. In a transient calculation, it would be expected to vary around a value of 0.5 Btu/hr-ft 2 IF to 1.0 Btu/hr ft2-@F. The length, L, Is taken as the 'AC' dimension from Figure 2.1.3-1. This is clearly much longer than the actual length of the yoke. However, the yoke dimension is not given. So, the more conservative, longer length is arbitrarily used. This presents hardly any penalty In heat rate to the room because the value is used in the tanhO function which is barely sensitive to the length. For example, using the 9 ft value tanh(9) - 0.9999 and using half that value tanh(4.5) - 0.9998.

m                                                    =   --
YANKEE NUCLEAR SERVICES DIVISION                        CALC. NO. WC-1347      REV.      -  DATE 11-1-96 TITLE Main Steam Tunnel Heatun Calculation PREPARED BY Jim Pasnas                       REV VIEWED BY                           PAGE    13  OF h                    1.65 Btudhrfe-F                Rp erence 38
                                           = 3 in d                                                   Reference 40
                                           =  0.25   ft
                                           =rtd P                    =,n (0.25)
                                           = 0.785 ft k                   25 Btu/hrfrF                   Table 1.3-1 A                    =             2 4
                                            = 0.049 ft2 m
m                =  l-  165 x   0.785-.
25 x 0.049
                                             =1.028fr              .
                                             = 108 in                     Figure 2.1.3-1 L                     = 9 It                       (See discussion above)
The temperatures will be taken from the GOTHIC run using control variables. This will allow the temperature difference to vary with time .to more accurately represent the changing heat transfer rate.
The source temperature, T., will be taken as the temperature inside the main steam line. This is highly conservative since the more appropriate value would be that rf the bonnet. Calculation VYC-660 (Reference 40) is a state-point calculation of the heat conduction through the same MSIV structure. For the state-point modeled in VYC-660, the steam inside the pipe is modeled at 545'F and the bonnet temperature is calculated to be about 375'F. So, as expected, the bonnet is cooler than the steam inside the pipe. However, the assumptions in VYC-660 are not all consistent with those of the present calculation and a definitive correlation between these two temperatures is not readily derivable. Therefore, using the steam temperature is certainly conservative since.it is clearly bounding
  - the bonnet can never be hotter than the steam.

                                                                                              --V "
YANKEE NUCLEAR SERVICES DMSION                  CALC. NO. VC1347           REV.    -  DATE 11-1-96 iTITLE Main Steam Tunnel 1eatuo Calculation
* PREPARED BY Jim Pappas                  REVIEWED BY                             PAGE   j4      OF The ambient temperature, T,,, is the bulk room temperature calculated by GOTHIC.
So, the heat rate generated by a single yoke rod is:
q    .f    x 0.785 x 25 x     .049)(T - T.)tanh(1028 x 9) q =  126(T; - T,) Btulhr Each of the four MSIVs is to be modeled as having flive wirh rods, and GOTHIC requires input in units of Btu/sec. So, the final input to GOTH(C is:
q =4 MSIVs x 5 Rods x 360 hrsec xxt6T  126(T   T) q a= 0.007(r. - T.,) Btu/s In GOTHIC this will be represented as a heater with a heat rate of 0.007 Btu/s multiplied by a forcing function. The forcing function is in turn equated to a control variable. And, the control variable represents the temperature difference between the main steam line and the room average of the tunnel.

YANKEE NUCLEAR SERVICES DIVISION              CALC. NO. -C Y=-l347      REV.      DATE 11-1-96 TITLE Main Steam Tunnel Heatup Calculation PREPARED BY Jim Pappas                REV IEWED BY                           PAGE  15  OF _
Table a.iLli Main Steam Tunnel DimeunQalData I.
Superscript numbers refer to References in Section 4.0.

YANKEE NUCLEAR SERVICES DMSION                  CALC. NO. VYA1347           REV.      DATE A1-1-96 TITLE Main Steam Tunnel Heatup Calculation PREPARED BY    Jim Pappas               REVI EWED BY    -
PAGE 16    OF Table 2.t -1kb Msal imnil~a     V Length                    2(           ~2           2            ~2 200)                     424 .002'24 Height I Wkdth           25.5(12                     8.0t'       ,        25.5 Area.                     612.0                       192                 420 Thickness                 3.0(t)                   0.25 in""               3.0 Superscript numbers refer to References in Section 4.0.

YANKEE NUCLEAR SERVICES DMSION                    CALC. NO. WYC       1347 REV.           DATE _11-1-96 PREPARED BY Jim Papoas                     REV IEWED BY                              PAGE 17     OF Table 2. I*2a eipg~LnD~
MSIA                 IBM                     80             Z           _
MS-1B                  is                   an                               543
* MS-IC               181M                    ea MS-1D               18("m                   80              3.5&n             543")
MS-4A0)              10'9)                   80              3.5(19)           543(1I MS-4Bb')             10(9)                                   2.009             5430' FDW-1414i        6     1   (      _)           120            20(va)(w           373a)
FDW-1 5117              16(18)                 120            2 0(40)a.)         373    )
MS-5A0)                3(fl)                 160             2.0(22)           543
                .RCIC-2")               4(21)                 120            None2"1            140")
RCIC-813                4(21)                 120              2.0(21)          14 00 )
RCIC4AW                 4(21)                 120.                              1 40ch)
HPCI-1 SB()             14(2o)                 120              2.5'"2           1 40(d)
HPC-14Aw)              14(20)                 120              2.5(27)          14 a         Notes a)   Reference 27 says 2.5S b)   HPCI Steam Supply c)   RCIC Steam Supply d)   FSAR (Reference. 32), Figure 6.4-1 (Highest temperature at Location 2) e)   Supe~rsipt numbers refer to References in Section 4.0.
f)   RCIC Discharge g)   HPCI Discharge h)   FSAR (Reference 32), Figure 4.7-3 (Highest temperature at Location 3) i)   Assumption 1.4.2 j) Assumption 1.4.3

YANKEE NUCLEAR SERVICES DIVISION                   CALC. NO. WC-i347     REV.          DATE   11-1-96 TITLE           team   TUnnel Heatuo Calculation
_Main PREPARED BY Jim Panpas                      REV IEWED BY    .
PAGE - 18    OF Table 21e-2b PiDins HeatTafer Surface Areas I
I          MS-iA
                -,. . .                  8.062                9.000          46.56.T)            304.3 MS-lB               8.062               9.000          46.5t7              304.3 MSIC               8.062               9.000          46.5'"               304.3 m S-ID             8.062               9.000          46.5(-7              304.3 MS4Ao'               4.781               6.375           17.6(19)             81.8 MS-4BP'              4.781               5.375           35.2'9)             135.9 FDW-14116              6.781               8.000           45.7('e)            239.3
                 ; FDW-16117              6.781               8.000           47 .913)            250.8 MS.SA'c)            1.312               1.750           46.7P)               91.7 II RCIC-I('            1.812               2.250           28.712')             33.8 RCIC-2')            1.812               2.250           29.7(21              35.0 RCIC-8B&deg;             1.812               2.250            655(21)             14.5 RCIC-8A')            1.812                2.250.          11.42')             25.4 HPCI-15B()             5.906                7.000          40.6()              202.0 0
HPCI-15A( '            5.906                7.000           9.2m                45.8 Notes a)   Area 2ic(Outer Radius + Insulation) x (1 ft12 In) x Length b)   HPCI Steam Supply c)   RCIC Steam Supply .
e)   Superscript numb-3rs refer to References in Section 4.0.
n) RCIC Discharge g) HPCI Discharge

I i
I !

YANKEE NUCLEAR SERVICES DMSION                 . CALC. NO.. VC1347 L.       REV.    -   DATE 11-1-96
                      .TITLE MainSteamTunnel HeatupCalculation PREPARED BY JinmPappas                   REV IEWED BY _                          PAGE   20     OF
                   .       I ..      .              .
                     .2.2       GOTHIC Runs 2.2.1    Run MST1 I"
This run of the main steam tunnel heat-up represents a typical loss-of-normal power event. On a loss-i .I tI                of-power, the HVAC system trips and the MSIVs and feedwater pumps isolate. The room then heats up because of the heat gain from the pipes within it. However, the major loads are from the main 1I   7?, ,     .
: 1. '.            steam and feedwater lines and the fluids in those lines are not flowing. Therefore, their heat gain to i   !
i,    .
                      -the room diminishes as the transient progresses and the room eventually peaks out and then begins to 1:-
:1 , ... . .
drop in temperature.'
This run represents a typical heat-up of the tunnel following loss of ventilation, however many conservatisms are included so that the results are assured to bound a true event. These conservatisms include:
* The initial main steam tunnel temperature of 150'F is based on the Reference 30 data and represents a peak room temperature as opposed to an average room temperature. A more representative average (initial) room temperature would be something lower.
Miscellaneous structures and equipment in the room are not modeled. They would act as heat sinks resulting in a temperature rise that is slower than that predicted by GOTHIC.
* Miscellaneous 'cold' piping, such as service water piping, is not modeled as heat sinks.
* Wall openings such as ventilation ducts/dampers or pipe sleeves are not modeled inhibiting cooling by natural circulation.
* Natural circulation through the RRUs is not modeled. The RRUs trip on loss of power however they continue to receive cool servkle water and would contribute a small'amount of cooling.
* On a loss of power, HPCI and RCIC would automatically start resulting in flushing some of the 373'F water from the feedwater lines and replacing it with 140"F water. This is not accounted for.
* The MSIV. heat gain is conservative as described in Section 2.1.3. Most notably, the source temperature of the yoke, modeled as a series of fins, is the steam temperature itself as opposed to the bonnet temperature of the valve.
* The floor as a heat sink is not modeled.
The non-conservatisms in the model are:
* Main'steam line drains are not modeled. They would add heat to the room but only a small amount because the lines are about 2Y2' NPS. This is believed to be counteracted by the lack of 'cold' piping being modeled as well.
* There is no account for MSIV leakage that would continuously add a slight heat load to the main steam lines. (Run MST2 in Section 2.2.2 accounts for this).

i YANKEE NUCLEAR SERVICES DIVISION                  CAC. NO. V YrQ347           REV.       DATE Jjj1-9 TITLE MTunnel                Hp o    Ci;-tatidn PREPARED BY         Jim Pappas             REVIEWED BY                             PAGE   21    OF There Is no spatial definition in the model, therefore an axial temperature gradient is not calculated. Because the HPCI/RCIC temperature switches are in the upper, ceiling, area they may experience a higher temperature than that of the bulk room. This is believed to be counteracted by the initial room temperature of 150'F which represents a temperature friom' a hot area of the steam tunnel.' (See Assumption 1.4.1).
It is believed that the conservatisms listed above far outweigh the non-conservatisms. Therefore, it can be assured that the true heat-up profile of the main steam tunnel will be a curve that is below -
and therefore bounded by - the GOTHIC result.
The GOTHIC model MSTI is shown schematically in Figure 2.2-la. It consists of the following:
* Volume i representing the main steam tunnel.
* Volume 2 representing the main steam lines isolated at t       0 seconds.
* Volunme 3 representing the feedwater lines with pumps off and no flow at t      0 seconds.
          . Flow path 1 connecting Volume 1 with a pressure boundary condition I P. This flow path and boundary condition are used to maintain the pressure within'the main steam tunnel at 14.7 psia as the air heats up.
* Thermal Conductors 1-6, 8, 9, 11-16 connecting the heat sources and sinks to Volume 1.
Thermal Conductor 7 connecting Volume 2 to Volume 1.
* Thermal Conductor 10 connects Volume 3 to Volume 1.
* Heater I representing the MSIVs.
Appendix A contains the detailed listing of the GOTHIC input for this run. Included are graphical results and calculations validating the run. The model is run for 7 days to determnine the temperature rise profile of the air in the main steam tunnel. The heat-up of the main steam tunnel is shown in Figure 2.2-lb for the full 7 days.
The graph shows that the steam tunnel reaches a peak average temperature of 174F after approximately 3i of a day. It then drops during the remainder of the transient. The peak is considered to bound the actual peak that would result during a true loss-of-ventilation scenario because of the conservatism discussed above.

Smwn Tuwbl USTI        :
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GOTHIC 4.1 (Alc 10/29/96 15:56:41                                               i ICALCULATIO YANKEE ATOMIC ELECTRIC COMPANY ]
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YANKEE NUCLEAR SERVICES DIVISION                 CAMC. NO. VYC-1347          REV.         DATE Jj-19 TItLE 'IMain Steam Tunnel Heatup Calculation PREPARED BY. Jim Pappas                   REVIEWED BY                  :          PAGE     24   OF 2.2.2    Run MST2 This run of the main steam tunnel heat-up represents an extreme loss-of-normal power event. As stated in Section 2.2.1, on a loss-of-power, the MSIVs and feedwater pumps normally isolate.
However, the poissibility exists that one or more isolation function fails - suck as a MSIV not closing.
Fufthermore, as stated in the non-conservatisms of run MSTI, it is more than likely that some leakage would exist past the MSIVs.
Justifably uantifying such conditions is not stiaighit forward. However, the situation can be bounded. The most bounding scenario is that nonelof the main isteam anc. feedwater lines isolate and the steam/water continues to flow. The initial temperature of the fluids within these lines is, therefore, constant throughout the transient resulting in a much higher heat gain to the room. This modeling technique also clearly bounds any postulated leakage past isolated MSIVs.
The model, itself, is identical to MST1 except:
* the volumes representing the main steam and feedwater lines are removed.
* the conductors (7 and 10) that connected those volumes are moved into the main steam tunnel volume as internal conductors.
* the boundary heat transfer coefficients on conductors 7 and 10 are fixed temperatures representing the steam and feedwater temperatures.
All other conservatisms and non-conservatisms listed'for run MST1 remain in this run.
The GOTHIC model MST2 is shown schematically in Figure 2.2-2a. It consists of the following:
* Volume I representing the main steam tunnel.
* Flow Path 1 connecting Volume 1 with a pressure boundary condition I P. This flow path and boundary condition are used to maintain the pressure within the main steam tunnel at 14.7 psia as the air heats up.
* Thermal Conductors I - 16 connecting the heat sources and sinks to Volume 1.
* Heater 1 representing the MSIVs.
Appendix B contains the detailed listing of the GOTHIC input for this run and graphical results. The model is run for 7 days to determine the temperature rise profilel of the air in the main steam tunnel.
The heat-up of the main steam tunnel is shown in Figures 2.2-2b for the full 7 days and 2.2-2c for the first four hours. The first four hours is of particular interest for Appendix R scenarios.

YANKEE NUCLEAR SERVICES DIVISION                 CAMC. NO. WC-1 347     REV._ DATE 11--9 TITLE Ma            unnel Heatun Calcution PREPARED BY Jim PaDoas                   REVI IEWED BY                      PAGE. 25    OF The graphs show that the steam tunnel reaches an average temperature of 172?F after four hours. It continues to rise until it is about 207'F at 7 days and still rising.
As with MST1, these results are considered to bound the actual temperature rise that would result during such a scenario. Furthermore, because the scenario itself is extreme by nature, the results greatly bound any possible steam tunnel heat rise that may bc postulated.

0  N I I
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                                                                            -P/C      OF _~c l

Stun Ttu   -alMS72 Wed Od 30 07422 1996 FILvvc              Z.Z-Z? GOTHIC Verion 4.1(QA)-c - May 1995 Steam Tunnel Bulk Room Temperature CD
                              -         1 2                 3       4       5   6 Time (days)
GOTHIC 4.1 (QA)-c 10/29196 16:07:43                             __-
IANKEE ATACHMEfHr ATOMIC ELECTRIC COMPANY
                                                                            'CALCULtATION F1W.ia.kyt~j pAGf  Z? cf  .
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YANKEE NUCLEAR SERVICES DIVISION                 CAMC. NO. -VYb1347       REV.         DATE    11--9 TITLE Main Steam Tunnel Heatuo Calculation PREPARED BY      JimPaoas                 REVIEWED BY                          PAGE     29    OF

==3.0     CONCLUSION==

The results of the two G3THIC runs show the air temperature in the steam tunnel to approach about 172PF in the first four hours of each transient. Both transients show almost the same profile for that time'frame because the cooldown of the main steam and feedwater lines is not large enough in the first four hours to' have a significant effect on the room temperature rise.
In the case where the main steam lines and feedwater lines isolate, the peak room temperature is about 174F at approximately 18 hours. In the case where these lines do not isolate, the room temperature rises to 174"F at about 6'h hours. It continues to rise and is about 2077 at the end of the seven day +ransient and still rising.
As discussed in Section 2.2, the many conservatism included in the model offer a high degree of confidence that the GOTHIC results envelop any true heat-up profile of the steam tunnel. Therefore, the actual room heat-up is expected to be something less and it is concluded that HPCI and RCIC would not isolate under the conditions modeled.
The results of this calculation do not affect the FSAR, Technical Specifications, Technical Programs, or controlled drawings.

BVY 05-072 Docket No. 50-271 Exhibit SPSB-C-52-3 Vermont Yankee Nuclear Power Station Proposed Technical Specification Change No. 263 - Supplement No. 30 Extended Power Uprate Response to Request for Additional Information Calculation VYC-1502, Rev.0 l Total number of pages In this Exhibit I (excludina this cover sheet) is 43.

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Control ]Room Heatup due to Loss of HVAC; VYC-1502 Rev. 0                                                                                                           Pagpa 1.0           Objective The objective of this cWcilation is to determine the tempera ire of the control room as a function of time as a resilit of a los:; of ventilation of the co.Atrol room due to an Appendix R fire                io. The suicess criteria 1,r this analysis is to have the control room maintained in a habit ble'condition and one which ensures equipment operability. Based on published guideiis, both are safied by maintaiig control room temperature at or below 120&deg;F [1]
2.0 -         Method The 6opltol room is an inclosed room vith concrete vAalls, floor and roof. The roof; north all S . aid are expo~d to -the outdoor environment while the west wall is a common wall ith. te turbine builiigad the south. allc is a omnnn wal with the reactor buildin.:
                         -. rote;
:         aAlsilin       -       ia tile    g '. blowtle         ligthticng       contrl iael                hldb room roof  abole   the which sseddciigsncue ventilation duing is t    .-riid.
                         .. .                 i         it                       sand lig                                    fanes"ar']i by he sdsped ceiling strtktue.
                         ; he.,.cobtrl 6 mias ebicd h itas                                                      nyCiontr6l panels and cabinets with a variety of
                         . .elt ico                         nsuc has ra, indicating lights, powlrsupplies and control systems.
e~           R: .-.                  e. as Well as 'other snintdes (computers, refrigerator, copy cheat Tis beat is iorially remnoved by an HVAC system.
Bu .it               postulatethatan AppendixR-event in specifin locations may render the HVAC lg
* a .aint of the contol room.
                       . .c                                                      %en                                               bioi      4iieof key parameters. The first keyj..                                   i
                                                 '--'.'         'o                                                       :                         musSt be removed by convection-
:';;          ..;.no'            tuiii~r.-mGiven''alossofventaion mass transfer is not available to -
r                reov -toh     the heit throhest-.
: : i74St~                           iheiviormidentnn.:Since hit removal by, convecton and ...
   .;                     ci~aaui'Ofl                     ,
* 0            to               ietCDJ~i{5rCWICTel                  across lewl,-ina Hetinn oti        en "'-ed  ldfiecsjc wihn eniaCi aoeirr                                                              i            a ieoe  t hmtr~o' ie        t w          t ar t m e.
 .:.                  ,              ctin an Mlln-                                           ofno ifiie- s                                    d           ugb the surrounhdinlg walls to thei.;
g          w                 i~             ai                          jfheat rEnmoval is deedei on the thermaln conduciiy of.i th . . . .:ile~wiudl.lfeconete.on of heat from teaiine room to the inside wll.
ffi:uuf: . .;and                             frrntieod w~all srae to ft e ouiitsde ai.' .These three facors mnake up.the
&  .:-.                   t: rmi r iuthfe valLtThs lira r istaS iceis the second key pareter and tle
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Contrbl Room Heaup due to Loss of HVAQ WC-1502 Rev. o                                                                                                        Pace respobseby acting as'a beat sink (as                                               oomheats up, some of the beat is absorbed by the interrWal conddctoras it heats up) or a heat souce (if the heat sink started out hotter than the surrojnmiig air, it will give up its beat to the surroundings.)
In this problem, Al tee key ptaraeters need to be detemined in order to solve the transient heatup of the control room followu g a loss of vehtiladon. These parameters were determined usiacombinat on oftest and aaysis. lhe physical configuration of the outer walls of the
                      .Con*im ca eemied foim drawings and allows for determining the heat removal
                                       .du4convection and ondton. Th.e inte                                             confgdumion of the room with respect to interil heait sources                           4ida si1 is more problematic and the method of determining these two pb~urieiei 4`6lned-m he subseqtimt -seaos YThetnsient                     control room heatup will be calculated using the GOTHIC computer code L2].
                       -boun gscenaio isdetermined in order to irovide conservative input for the code.. In didn o;t:& conse                                             :ve:,imrdary-and i Ucondit a                         ofofitrol                  tomperatre i             ersus teare -chdoe        wich can    Thethen GOTHIC    in .:
be used6ode        't
:e                            .eatr atinn eas                                                    ventilation, to the control room.
t.eGOTHIC 64np4t& code is capale-of solin:g g theemo-bydraulic problems and bas.
many           Ii eful tools (ihtcral conductrs, heat so , conro mechanims, air properes a a
:s-:idn ofpsurne nrmperature,.d). GOTHIC has bee validated for this type of iplication [2]..
24.
g::        .      2.1- . Heat Sources.:
e ea sources in icontrol room can be classfed as follows:
l.cti-i.0 - so                                     of biat due to iR losses in elecicdal equipment A variety of electrical
                     ;q. -- :tthh          ns.                                         e conrfbl room (computrs, monitors, relay; ambient lighting, Wconitwlaand mdica4Znssonveniencessuch as :refrigeratiss .cpy -machines, microwave ovens, A
tc Ambieth-if-        i           the                                r1igher             is
                                                                                               `convi~.e...urh             than the inside air temperatu                       eat is a                                                                 -d         nn           ito the room.
. .            -                  .nel       -          ohet                                    Andthat beat L, given off to the surroundings. The total heat
                     ;eit               by                 peo pe dson the numer of eople in the control room at any one time.
ai4 t             type-,-                                    eing       romed(the more strenuou the activity, the greater thiheat '
:-by' Y.
24r:              ba' beeno compile                                    ;Rvwof HVi!C docuents reveail at-least one source of inforaion                                                    : :.
M'fht-______IS 6                                  _____o___p______.c                   tat     o   w      nj 2'','~~~~._                           .,'.'''
                                                  '*'.-'\-...........
                                                                 '''.                                                                    *..     ,..-t.   ..         .
__  __  __   __    __      ;    '    .............             . 't___
                                                                          -S.ist                                  ,      .'.'                                    of               s.

   .,I i
ow RomHe.l                 dto     11S      of HVAC;           VYC-1502 Rev. 0                                                                                                             pase<
I alcule tht totalkieat generauion in the control room. This data insth ofE data teiod .he HVAC system IncldingjII1l wate flows and tenpeuratre. In ldihbn~a ;
tet          as.pro.7 id r f whic ixas mthe ifIowtrough the control room irAteiiircs.                                                can be used to calculate the totadbeat rate xThisdt
~ihfiecon~trol room;:.                                         ......               ,       .                -
smin T:is combinaton~ol ttest da a                                   nalysis an adequate heat source can be detesmined m-ent  transe heatu                        . .ditin                  **.*.                        K 2l,5;} -.t ms             s, . .
.          - .F-.,..,,.r, at          tranifeir: ht               fr.76m t               .inside.ar to the outside ambientair. Ihe deteiinedbywldion'ifomaion-ondrawigs. How6eve.
ib-oe tbt-asx r                                      6ffbit t6 the room lddonot trasii
                                      *-more         dificxlt.t6                        ;jj ;mmhee ties s.scan bv in. many                                                                                     .-
M-aiw Imr'. >9-m soshe   u                               "tc.                              i        u                                           e.'.       *..         .
aM t            was c                        sicteoalos               diafficqits                          ofdatathat cn b A,!.~I; , -1               +             ... SZ ,; 0 ;,
oution                                                            -oow
                                                                                                                           -fd             o                                        Xcn1:.
si          e      o ia                      test T-hiunng A plained outge of thee HVAC 3XS~,hi                             -A,                                                                                   ii-:
                                               =0 = .......... 2.citr
                                                                     -".-:                       oo                                   qeuordi             e                         ifii.
L-canto    $                                ndiddcoittime of-teit" ieteted inihe '                                                                                                    .
i~sintm~clug th tiasiet ten-e-ure                                                    ddata        s clo~sel aY-0 iile ca  hnb sd intecnrlro                                                                    oelfreaiigte                                                      .:
* 4f -;f-                                it -nta etidios                                                                .                     *      .                        .  .-
usuedi efointii ac1d.                                                     Tefort
                                                                                                                                           -                  mcdIh9~t n          hntett minks. .eco-s                                        model assesd thie &,ntrol wrno                                                          ..
jj                  j            .
o4-mmg          ad~~sof& . AC~to the cont o rpom.'' 1,
                                                                                                       -c ottoSw~i                                                           osbe optosIat can mitigate-<
t.v Winror a.o                                                                                       uo.                                            .-             ';.
                                      )io'r V;s
                                           'Montro I'*C HI 3i:.tiiositem ral     t IM .'.
befhecfrl6i'tnprtwreiow i          l AoAo e
i
                                                                                                                            &sect;S.a pM Ci 110 0ao           Fi (othis&            aroccurs> .
                                                                                                                                                                                          +
12............................r.....t..6 -I6
                                                                                          ....................w..............-''*' '.-**bel-*ow........
dh
                                                                                                                                                                                                        '                 'X'    S C........*,.-*, ............ ,.'..,........
                                                                                                                                                                                                     '              *f.^
:< '*:~.~
                                                       .,.        .                 o.                 .                 ;
s ;        i.;     .- .......
                                                                      ,:fs.,*
0..'_.....
DI.                           .. &*,:'.,,e                            I* to:N: .............................
a < ;'':~ ;-'''.'..f<. w                                    AH

=MM 1,:, " ,"
i I        . . I
                                  ,s of HVAC; VYC-1502 ReV. 0                                                                          Page al fm~inhe                         soom  iintrol Itis         exjiected that shortly after four c~ t~                 will r           and exceed:120 OF.
a *w.4 aclion can                          that th6 contiol room xemins
                         *. I'~-:
I cU1 t                 of mntigai the room heatup
                   -,&i;at-o-fo                our;  TiV      e             m            e
                                                         ~        lt     1id ln t~!controI room. It was foubid that ensurthat the         contOi room long-term
                  &yatleasld                                               ti uin ontrol -iooi h.at lohm.) it 'Wu in Vtda                 _q '
                                                         .alf                                   on                 a control room heat 52 perent' k l in,.                     t the contiol room                     Iong-term      temperare
                                                          '~~~~~~~~~~~              .. ,,.'..'.
s in this            ladOM Thse iude:
lratedtermned tby test- is iy1                                       0 pe                to accout for maYeat9~   ~leerwndby tesis decrased by 10 perrcentto
                                                                                 ,,~-'-'          aS Wdi 'o         il           c,s'.'*     e h                          tc i tps. fi.ti*    s -cou be troloS            .
HIM tr     ;~        -n*o* * ..9
                                                                                                    -c    ith' n     do our o t                h                                         t daxs ro.t otarsWlmth     jiot a'ccntdoi(hsoudnrae the eat t       -~61 - -t a               -               --                                                            w
                                         ...-<     -  r-    -      -   ~                 ;;     ~. 1; j..

I-----                                      -                                      MCMUMMM                                            --     M    .         - - -  --- --
Contol Room Reatump due 7 L=s of HVAC; VYC-1S502 Rev. 0                                                                                                           Page 7 3.0          Input and MAisption s The ipu and assiptons used for the GOTHIC model wlM be addressed step by step based
'                uponh GOTHC piit tables&#xa2;: lhis secton will describe and address the physical cofig ion of the""""                                              rom model.                                    aSton how fte model was used and 4 willsddiiss ly                  o caons      any to th mdel tht wemae in oider to aess specific variations
                -,:q~           "              s . ...             purp or' clcuilating ,te tansient control room heat up. For -intance,
: ."e                    thatn`,o                              duplicit fiet 6-ntro room heatup test will have specific initial conditiions. The n6l1:ati                                                                           fe tilin 1cm ring an A.iR                       f ivhe                         diernt initial codtions. This seon descri                                                 the co             stwiratw                                          *Ii:id moin  l         A    lisng     of t'=e   GOTHIC            input      tables     for the control o           em tr modelis                        f incldied s APpdix A. The input tables for the control room
      ;. -.- ts~ti~t~eis.is nclued 'as .di                                                               1.'
Figur4;3]E                         ][6i            npovides a -cbematic plan view;of the control room envelope.
3e.1     -.                                                                                Mode5                                s.GeThconoloIC Th.....e Conol;.                         is moled as a fw~o lemgd-paramer                 :                       volimes Thc GOTHIC model is on.:-                                       i Figur+2.Wle the Cintrol Room has a more complicated dtree-.
geomer                               iloyvby two ped- mariatio                     rae
                             . :k.:                                   into aecoiu857          Z From .j~ure -- th' Fkooi                                iz~sideplan'.dimensions           .        are: .80.063 feei(twest                        wls) x 48.5wls):
                  , feV                    t'.')(io~i;,c                                             -Conirol R              h           sparate, stfive                   smaller      rooms along
::ion                                                                      it was fd                    t  ihe  drs       o          ese               ooms are  norally
       .umed                                                                        t    S              tha tfie dors                       rooms are opin an.Oid the
                  '';~            or these :o~ w                                          up              6;s=e rate             & rest f the Control Room. Thus; the Ir              a;         f-orthContro Ro~o'm issi~mp                                        i -           -f W two side dimensons
                                       ' ' imd-      t       r.,>:
                                                                               ..-,                                                    . t.   ...........
                                                                                                                                               . ... r hefoor                    '        "on               ..            trol Room-is 272.5 ft andhth roof elevation is 290.0 ft [3]...                                   '      -
                                                            "i5467,,ftfor a tol insi                                   i ht of l5.833'. However the Cont.ol R66Matedc"                                                                           "cilig             teome*trd' dd w'.instale          ciling.. From         h oberwtiSn,<it is~i
                   "-                                                      ~t'              re.op cphings arei airVy,!fficip -in prevenfing coimingl g ..
s: E, ULAf.                                             eul g                     rwfwir  f         i 'iz. ,befmr.,                            re, ue          spae
                                                                                                                                                                              'ir
                   -:i beii                                                           idht                 illbemodeled           -as    separate           h:Apc            dper         r                 :
                                                    ?S;           ;jtoitiik;IciToisof Aaretw                      o iln Te st is a 34                             A     x  32 -f    (8  xi?
                   .::                    irld cb                                              x                  38 ines above trst                         ofte ceiling l7]. The
                   .second results an s. cefihlev                                         of 10 t aboehe Control Room 'loor and represeits tbe                                                       -.    :;;
baln. of         DthCorl                          R~iom             c     irgel.            tinl ie five     oficeuility spaces the drop ceiling
         ;--.C.                                     ...                           :ia
                                                                            .:.....                       ,              dies., be; on.-s p    =   ttt4~
Sai th.'
                                                                                                .:./.f.

oom tcanp due t9 L=s of HVAC VYC.I1502 1kv. 0                                                    PWgc EID                  S
.4..r'.
FI           272.
0~:'levcllo:.
1-. *.
                                                                                                                ~    -TZ 2r,  w.
4*  .
3 P-01 i m
                                      ..  *.*.4,J 3 6 V 8
m~ ~P   c~i aan V ew                ..
                                                                                      .     .-.                     Z, orh 3 6Becsnsejny
                                                  .i 2Rf-x te A 0e9Co     t~ox        flo yx3pcin ilxfl4S             A. 373or                                       '               *'
v 1 if4rCilp.7I f oo                                        c z e u fi i n o                i n
                                                                               --  **       4   ) ,    .
im i o ngy s i m n h ii v d a o u c a c ~ d x i i~1 1C O~ &e ~ f~ b~~                                            wh

                                                                                                                                          -   *   -P:           In
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-                                                                                                                                                                            9 C                                                                                                                                                                            -A.
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f Control ;tLHP                          '         ticqs of UVAC VYC-1502 Rv.O                                                                                                     Page
: ; .  ;   l,   X,,.,..:       s'_  ^.S4  :,.',  ,',;,          . .' ',                    '       "
                                                                                                                                                                                                                        *9 JCRbi*6Heaft,7omt ModS
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                                                             ,> -,"       :         ."      .;          ;      -, -   '"        ..     '.-'-.             .                                                                I q
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                                                                                                                                                                                          -                    .h
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* C*                             A adc iif 1'Y Co0n'l61 RO 6A N....
5.s.    .'-w
: i. m1          5                .
                                                                                                                                                                                    *  -,   . .9 9IC.b 9   $4.

:,'  -. ,.  , T. - , -                                                                                                 .      .    ... . . 1t v  . w.
Contoi Room Hmp due t%. Lss of HVAC; VYC-1502 Rcv. 0..                                                                                          Pagc
                                                                                                                                                                                                  . l shoi d be noetaed.;lie contol roon. hi liarge arrays of cabines and control panels. Most fIt                                                                                                                                                                             : *:
cabin Mo, i.en-backed nd .notf1y Senclosd. .Thus, the fre volume of the control
                                                                                                  ' i, that are enclosed, it is room is not dcd f-ithe             o        ii in               Foi                nl,ie                                                                                             .      .
veoitespt tie avnrage room ;to                                 t          eat                   the cabinets do not reduce the                                            ;
* freaiiS space     since          the      hat  load       in  the     ontrol        drc,             in    gener,        contaiiied within the                                     s .!
:        4 Th-:hydraulc diaim                       is ctalciite from 4*Ar                       I Wetted Peimcter                                                                                  .
4 -l*(80063 .:49.573)/ (2
* 80.063.+2
* 4.8573)      =. 6046        ft                                                               . .?.
                                                                                                                                                                                                  .e >
The volume above the,,dwp ceilings can b~e bund .b: dcida g the total volume of thes                                                                                                           j e       e by t              c           twazid subt                       tg e *olume below the drop ceilin cal                                                                          . :
* s;
:::              aQ ve Vbui A:i-vropCe'.iling                               80.063e'*-48.753'
* 15.833' -.41533.75 ff3 20267.4                                     fi.
:            *.o f.eoffie                                                                                                                 ceiling in tihe control room                                                  ;
              -Th el-oj                       eiin                  'rine-Is tin cicaua v-. Nolii 27            ft.         b05              85 L-                                                                                                                      N@ @           .      e
    ;        e:differee;d               ing                               i,.th                           betee          the ceiling el         on (rf                                :
a I
                              -Too                     ".                                           vlue lvaion:                           .,.              ,     -                             . 1,
                * ~-~~':Zy                     -~~>,         7--                                                        -.
                                                                                                                                                                                             -V wii:vltnneis Iihe                          aehi drauilic diamee of ihe control .                                          ..
i           wallsd aid.aWre -same plan dimen:.o: .
r                 usdto dc ondensaion                                 orvucration                                             .
ue is           t           A              aiWd P-h?1Ii              vaue calicuai      tedb e d..         W 'ITA s are       cludeI b t                    o               eArepressur bounry cpr rinore to                       econtrl room ind d                     cei>.g hei)pressure.
ent con ti                   f1                        90&deg;F in ccodance WIth
                                                             *.*a .                                    4                 n.
a-                              ,   S,
                                                                                           >*-                .: .~.                               .4.    -   5    .      ..    *U  r    ,I.*. -SJ

        .                             I        ,      -.....     .                                                                                         ..
I ICono        Heaup d                to Loss of HVAC; VYC-1502 Rev. 0                                                                 Page II
                               .Itis                         .dIhathuthe       control room. model .'ill-underoonly beating, no cooling, and will try tio p                   s`sA]rW Since-the 'bowday condition is set to maintain a constant prsure, w mass~rill    ftom Tfe control room to the bundary co:dition to maintain the constant
                                                                              . ';-finl      be no uf1w frm the bdixdary condition, the ternperature of the tTables                    'thru.uh                                                                             .3 F     ~ * '                         'l                 z'*
                                                           !                      t'             ,
Four flow p.h arespcified:in the Iodel. Two of these.flow ths connect the control rI',i. Sd cili :Volumes to hciirespe-htiv flu~d.boimdary condition and the othr.two 7                          cozhect thi6il'                                                 ilvolmnie With the contiol o                  o ie.i
                               .l::eilons^;aretarbflow pahs connecting               rly buitthe            voiimies tii 2witiii~evcifia                              conditions,    the connection The connection      height is set
                              .l :
                         ...                                                                                                                                              iig   i.S y           ,-2'fe              Te             -A1was      s-t t a ciendtly large area to ensure that
                               *t&#xa2;e        "     *o'vouie                                                                 o not ccUr.                 ckwi      the hydraulic diameter length :.eictto .ensure no pressrizion. Ihe loss coeffiient-w .ov                                    , e .soe                     oSS  1       U: not su.ffilcient to pprssurize the model volumes..
J     ~.~              -.      ot~                  o         :i       h
:.ivh1               p                             i W.he* .:t ............... roonmvolume bonirol           and the drop ceiling volume sre modeled ceiling~idov to .'a ow air circulinw* h p......................
6liijn>.';Ti ;.oW pgatiare~iedio model the                                                     nins as it allows bidire~cforial j;*
Sc O -
Uwe.;z
                                                         .               ob                                  C f
z zz,.
source.f;tint flow -trough,a junctonsl S tinwih            hdro control room isn the control TOOM'oieie
                                  .                                         g ihppanclilllo-i.the                      w         het an         r ise-d the nomially cooler air m
                                                                                - ." fall.: The                      onnecons-art Elevation andheights to allow
                                                 .h- 4contdl volume too inuncatwith he bt                                                                      of the dwp ceiling v              r0i .                                       ume:    -                             ,
E                   ; 3v. S8eight0.01 28ft                                ;
_                              .Vol:                           Eie- W83:= ; Height-0.01 AR                                         .-
     -                              ,      E                                                                   .    .i                        -                                                   :1-Th-lwae                 s40 ft2 ea                       based n A O p iebc nig re oved and the panelsare o~n
              . :            :Theiydraul eiametis based on the individuil 2 ft x 4 R opening:
10  Oc~ib~ebto d~tfffr     wWrWifr wee         Om Mwuol nokitwwb
                                                                               .          T.       *                                                                                             :    ;

w Con I RomHau 6ue to Lsf               HVAC; VYC-150! Ridv. 0 Th6 kdtaIA' -ij~   iAketo' bd                                                  e venter of the control romm thj~btien 4
v1dbet          in         '     `66       ',            -IF   u's   e, vbout 5I8-ibch (sue J. -   -
amonditioeinigtile           Aw,
                                                -,Z                           I MW                                                                                           ntslietf.;I9 44
                                                            .  .Mjhtb" w          =     ...
                                                   -.                 ej.
j,V-thcmaVoiutr                           hr ceeM~al-C  oe            Aft              c At.

Af~~-      .. ..w. - :Ale x   . P. -    - -, ---
                                                                                                                                                  ..C.,
                                                    , j *&#xa2;. . .    ;
4 Rev. 0                                          Pagp /3 Control Room Hcaw due to Loss of HVAC; VYC-1502
                   ;iohel                   th waii ofth drop ciig volume,
                     .i. i.        fthe-       of hie control room vown ix internairoom h                        tsirks.
                            - ~pn~ 12' Thasawc                  kulated
                  ..##3Est/WestW#                   .. : 8Q.063.ft 10.0fi.- &.63 f*
4r,.
A.i.           . 48.573 ft *10.0                 -. 485.73 f
     , a , -,, ..
l~gTle, Roof:                 48.73 ft*8O.63 ft              3888.9ft a              -   ;
                                               .i               :..80.63' i
* 5.833         467.0 it Val                  48."573 ft * . 33f*              283t3 ft2 en m heatsik               i        is;udas a          le to match.te test a Aallded to in thie rlA              descripton and detisii of ihe i b~ ion,                                    pct I d~' bi c'adleatse~rm
:                                                  film
                                                                                                          -.           hidgied to the a
thr ide~ot~eeofpemw                                                                        lArge
                                                                      =;,die.                               aa

Control PLoom Ieamp due to. iss of HVAC; VYC-1502 Rev. 0                                                                         Page/s'-
The eW surces           for (Volume A connee dons for themal conductors #1 - #5 & #7 - #11) use an assdmed i ou nv edion e trasfei coefficicnt represaita:ive of their orientation coupleA with a psepe.dtemperature bounay cdnditionh                                                    on the expected temperatures of the adjoining spa= le they the                                                     ffribilding tune buldg or outside air.
The spkific vilues                 'jj:forthe lheas              6nfer         ificients are cxplained further ia the next secin4                .         .:
_t      T       e C          cits            T.Ruis .t,,
Heat ti-ns fEents                        - 2 & 3 arethe nd";iconvection6rrelation coefficients co Idmg to                              a1, face up or'f dow                            m spectvly. These wre used .ith no Mo6difcaton.
         ,;i            S.                  ;                    +         S            ,            "          on 6he +at  tansfer:rcoefficients.
Th.,eseeic               sare                 i         desption of the outside surface}ha trafisfir coefficients or;           4 9 e.d ~ibnientemperatres ire=ir6 e. btC                                                   doei not!allow one to
             ,B5Se Ao-Xe.'IOoMfent wn                                               as.ecied                              The Vlues of these.
coci-b .                                     a geuerur1.O.F..tempn. .e diferece beten thie wall s i :
a2ndee
                  -                                               atrnfr                     tcoirelafons of interest are given by [10]:
                   .          :'                  ''   9t   '.t.for                             he h'.0.2VAf                                 oioia
:    ,o ;                     9p     tesfacing upward whenabated in the twi                      .'-ent V         - h:*-and          i-   r           UV           L     n ut
                                                                                                                                     . of;f
,.. '.:'                               . .'                                                              J zh'-                     a         for hiot           plateh              0.47.0fg wwdih                   in
                                              .                                               !     `- '     -:B7U-/'r'      .-
pleac';                            .t         sertcal
                                                                                      .                D., O .BoUa          2   r        c.fl Ffs o
tle', to *-v  *f ,
a                r           t ,
h vA                  -A

Con~o 6olP -Om Hcalup du to Loss of HVAC; VYC-1502 -Rev. 6                                                                      ae/
The boundary~&afidiins for the out+/-de waland roof and floor) surfame consist of specified tuansfdic~efcients 'with a constant abient teziqetature For those 'wajjs with o~n otorePosur&enaw-verage.-e                             rdur i chosen ba'ed on the SOLAIR, ietho of tk inoactbuislt rf.al..
fetig fth                                                                   hesuaunins 10 MWSLI                                     l]~m;.p.,   i Ta!'-~provid~                             _,~as     wAllandroofbhavbe~
th24 bour            ftrnsienft SOLAIR temperatr o d~r~jsurbace                                h     sta                b~en4rt6c~um is the :avera-ge tcmperat-m-.
i 1-"'.           Tble 1            -24jHqiir             SOL-AIRTempratre t~~t5.99~~900090                                            '72
      ~1V.
      ~..~          :2.~1S0;.-4.96/.                                             -84
* i
      .1   ~ -.                                                                               _ ____    ____   ___7 P4.   ". V                                                      ..
D5o4.                   J                         .,                     660
                   .*.,:.      2      Y.
                                       , V.-r4!.                   6S A-                       --
                                                      .-                       107.                     19.4
                        .O   ___   ____
                                 = .. ,
_   _  ___      ____ ___            AI     b.f     -                                           a___
          -.    .-               AbuIe         ii~                iub j tI sc                      inciorbatz"it              gmhlAf~           W
         .1         -,
                                                                                ~uUU~ammdiheA

I' .   ... . .            .   ..    . - .. . !E..    .. IM - -_. - - --
Contiol Room HeaWp due to Loss of HVAC; VYC-1502 Rev. 0                                                                                                                       Page LG Trie of fay.N&6 MOall                                               East WaJ                RZof Tc=n.
                                                       -                ~:rcii-p...*                   Temp.
l1200                                _ A4.:     *4.48                                   :.2023 1-136                                                                   1736                   261..2 i9 t                     .          -:-             89.28                      28            . 19I3 AVerge.;                        %        .... 76,4g ;                           .2               114.1.
Te verage                          SOLIR                                                     iathii Ahe transient temperature for ease of A :                                                         Dy6A1 thick                          concrete         lowel                condutivit) of the walls and roof the effect Tpof any bisi           sufficiely da                     ta                   cessenoso,               .space is, i
                                                                                                                         '                                                                                         * * *e;
   >.&#xa2;-        i'Al                 ;'-          fe                                .
e';                    them offa  -i                 .un tth            :u-d; i                       : :                        li                       s                                          buildics From the FSAR,                                            Fn 3 saximnu                       decne                          theturbie                egisc105F.                                          Hvr,             he.
n.amenltn-'euiej                                  .f                     tdhe i              bdi;dtt 1c0eThaway ottde facor
                                                                                                                            '.a                                                              m ratur    he                .
Biow'the nLnv1 room te aible l ad bttimon                                                                                                         is       tthe     outside
                                                           -s<b                                sthnle samiA syst- as the                                room nr rom husa,-emter-
                                                                                 'oI'oos       u . -rf This ni1               the cntto             room     i';to         c        o             r        na fite
:'           '5n+a'
                                 -1zit                   '^.                             6kimtrooteIs reqir                        d rthi                              udig.o. Fem the PSAR, the
                                    -10 ,;.;s                                                   .l        .t            l~~ftemj                                                    the.                    e:
                   .;&.W31,.                       :',:'.,                            uitf                                    I~bi6iid   -O e:                             Svcdofi tO.1-23]. .
Iofii.1i                                    .:-
      '.                        -BjexeeSalsst                                                             of i>:-         ,      - hE.abie                 l ibabteymhiS                                                    h
        --                          i9"W                         C~                       a 11166i  ffeh ibe4i~d6                                               ib:o n 7e hair thme iie vault ansrm                                                   aeh'e't                     a h cnrl om                                                                         thmE l                          f thsl~ i& ''.n.'             iiickcnrt~~b(oth&Wstells                                          itni~
ki            '    '
                                                                                                                                                                                                                               ;n    ;lo
 '2'.x                  ,'                                                -;       .- 'b                                          '             ':
jf                  2ftzh&MVCq:i; ':c 2ckcoucrolJoe    ).iitelab(Est orwla              ito i ll Ws                    ilsad                   ~
3a  ' -, I                                                      t                   '.'
                                                                                                                                                                             .'"d  . ro       -         .
f   'th               .    .' . _,..*
rxd-f-^.                                 -k          :uro.  -d                           -             lh                  -'. i*   6      room            -m-ta t:
is
                                                           '.                   ''               at         .......                                   I                                                                             ,
                    ..     'a. dM
 $gt_      '     .    '                     ..
             ',':            ';     ',             ':'.'                                            .(Nrt
                                                                                                 's.,'                  & Wes                                Mor .alin 2i              . 2......k-..                 coci6,'i                 Mis W...              ..     -:
SX"''~~~~~~~-Y.............w;',l'i.;.':,:"S ........,:-s

Codirol Room Heampduc to Loss of HVAC; VYC-1502 Rev. 0                                                                      Page 1;
: 4.             4 ft tlick conete slab(South Wall)
: 5.             5/8" tickceilingtile 6;-          12 1g                  s      r abstraiateral               tbrmal         conductor The                UC ai               atiec su           sionlrouti i describd in the GOTIWC User Manual. It
* .    . va                          gen'ai          me            point in .                    laonduc-or given constraLits (startig and
                                              ,Coe I                                       udres) and q$ualtative assessmets ohhe slab
              -g     il.eat .si i~ofi~          zents. Foral slb sides be qualitative assesent of the heat
: r. nsf-er codficiet isuVERY LOW. E regi is dvided into 10 subregions.
Ti
:te               b oiir                       o'thte control rm south Wall, 2 feet thick j?] and the adjacei.
rt              biildall,                            also; 2 feet thick:k (12]. This results in a total thickness of 4 feet' zeacW.                                                       -2:.          -    .
    ;....                                    r          .inthe(,3OT                 :model, concrete, acoustic tile and abstract therial      -:
b.ii ler                                           arcamdto have con taphysical properties and for the k(65 0F 200 t 0&deg; b is a ged assumption.
DedW 6 4 i~ibm/ft 2
           .1                                                                            {i3 b e 443
                                                                                       ;(3.                                               .;
Sp:          .        5.              TX fbmsl].Tab                                          e. 25:                                    ..
          -Aroustid     w' '   iei        t"    aeilTp    ;          2..(0 Chapter2.Tbe      3]               ,Mnrl iebad wetk feted'
                               *   .    ;          S  .      .       :      .* -cousti 2              ,.'
_4, Th.,.... r',,''' ' Co du.'                                              -. -
                                                                               .iy~0O9          .'-f ":    . -'. ;,.               .       .'    '
                                                            'W'Wm   'Can.

                                                 -    .. -    .                         .M!W'   Sw ~~     ..
I*....
Contol Room Heatup due to Loss of HVAC; VYC-15Q Rev. 0                                                                         PaWe /
Abs~ac Thermal Conductor is Material Type 3 (steel was used as the basis for this thermal conductor since a vis% in                                     on of the control room girres the impression that the majority of th6 internal conductor is steel in the fornmbf thick cabiziets.)
[.13, pg 6f.i SteeL:- APS Cl020]                                                                                                              4 Dcnst D 490.1 I'bm/if Ther . Coadciity     Tal                          27 BTU/hr-ft-vF:
SPCCITC Rest O.iO6TrU/Ibm-0 F F'                 S~glla,-
Two different heati are modeled. These are-
               . ;i- ;ecEricaleatIoad..
: 2.                 1   ci~ ts; 1.833 BTU/sec e         C.ltiil he            ai is determined 4ljiusti               d in Section 4.2.
T. -                             iQ                   cupantsie J         on61o               "ring-the Unient in question is
                                  ^::;obe12.       .e             load per person is taken as 550 BIU/br.
(C.r26,       .             0m]- ) " iab: . .. .                       w w- -
ivolumetoiisi&Waoethi                                      te included ii        modl.                         nei do    riot represent
           -                            Ila D                        used                      effiectof the operators e e `lat                                 g ihe ceiling iPa                            . 'l.                th.
                                                                                                                 -u,'t)mbi   ,   .                f.
        -. te ~onitpI oam.As uch the inu ausInhe: volumetri fAn tables ame chose bitr'niM Erow-td ensure, thtthe' air In th w o insIsmxdthorougly f              l       a      i       u             the model. They are msed to modelthe opening of the ceilig.
mquic     to control ti tiS that t occus. When the ciling
                                                                           ;ebyth&
e g'                             ivedc                    ths         oCpeniltes arck as o allow the exrbazhng of
   --        ,.,--        .i'      :     ''' :

   - - - - - - - -                                                                                               - __ .           2    pi:                                                             I ComW RPom Heip di-to Loss of HVAC; VYC-1502                                                            Rev. 0                                                                   Page*
Onevave, uick cose type, is includ d to prvent m
:.:-                                                                                                                        than one presure boundary voiue oh.ion t;t                   iib. *gdkve                 qu-heciCntl  etayon                  diyte,prolume WhA &j f                                uowpathsbe twinU1room
                                                              .i                                                              thw ow ptl to one of the boundary
                                                                               !rvn                                      '                o.           ;fifr C       kbati               o&
IN"                             -avsisi-dd I           a . oftiS::                                                 -asllkiid                        losss we in                           d in the flow paths losany ui                           iv            iez                 sos 'n          io pro                 any~moerestictive                           openig f
:l:res                s&            l                h            Low:
10-'~~~~P&                     -eitiS.:-
                      ..            l              M                            the valves and the volumetric fin. It is postuated that the operators
                            ;.         - .ctionto,                 on .theeil.tiles.'whea the conol room teperature reaches 100 F.
V,,.                      ;                  4;            d7ffncd                '            ;iable                                                     into a form which cm be used as
                      ~~~~~to.                                                                                *t'     .      Of.;,!t,,,
ThS~e d~fis u coixtrl variableis used to get the cIt rom                                                                                                 e so that it can be lo. - o                                  used a fing                Msfutoni for he components described previnouly.
1 -;*                        ..... *.:                                                   i,;:--;                .....        .,,     :.......:.-      ........
l vre                 arst.F                             initilly but can b modifed based upon the given scenario ie             ts T i                        ty           isfot           ieaiVe ay sOficant effect but is set to the esi     ;                                                          SGSecn .10223].
iy                                   ibort wi*thspect to di: rest o          -imt tr<.::            .lc                                          i~b            i,:                          iifile in tlz hednWia dMd                               ;B
                                                           ''                    S''.;*,                                                     ,'*,-J,,4v'is                                                     _
air, L'                                                     ~~~~~~~---W';-;                                                                                                                                               4'*-
* g ' f
         ,,t,>.                 wo^4
                                     '              ' "-'";*J"'-'
                                                * ** -         ,>. S,*;t-*? z's. v..-Q)-;*
                                                                                                '    '        '            '  -                "                -                                   "*           'L;

                                                                                                                             - - ---      ..   .--    .      -. f        ... - - ...
                                                                       --.  .- 1-.     . I     . .;_ .    - I.-   -
                   . . I. . - .        -   .   -.-     .. -  -   -
Coatrol Room Heatp di to Loss of HlAC; VYC-1502 Rev. 0
:                          ti:ihand ndfe intervals de Inc the trsieudperiod for dctemnimg the
                                                    - inihh roo                      Ib4j weich~ose5n 4 bound th period of ftaa (the                                                        m
            .i~ .                                           sa~                                                                          Y-SSR)taw                -                             .: ;..........
                  -ft           d   ,;   .iiPA.                       .    %-....:.
e anda
               . s o'A *ff s:e .Dt-4io7"A:rie 'the cioniiiion "' uations to cmieci ii~~~~t.e
: .,iiel #0.                                    ;Ows lihi                     e c                       Euiin 167 eaih cqiffitium mmne. Wbile.
i.':     i *o*n'r'te                                                             time do            ai,-t1,             i 1t
                                                     -V.i AI I'    p       -
Z~
'Vt.'."
,,,-i-.-.!.
                                                                                                                                                                                                 ~!7, W- z*~~&#xb6;~    .-. :t%..4t'          i~.iFa.,r          *- btt-
                                                                                              -;       Y-....    . ... .....         .       .            ..'-'*v*    ;
                                                                                                           '         <*-.;-,                          I^-  ,. ,
R;**-
                                         >           4          t       9          4       ""
t

Cotuut       Room Hea~up d&e to Loss of HVAC; VYC-1502 Rev. 0                                                                                Page?-!
32           Assumptions
                       ~easiptions sdi_..n this calculation are presented in se;arate sections as there ame some explct and iiplict; aptions th.t apply to the GOThC model and those that apply to the test dsta iteretionsa and calculations associated with the test data.
3.2.1 ;GOTH2IC ModelAssumpftio n desi~biig the basic. control room mod il GOTHIC input several explicit assumptions were made.                   a li                 ere for reference.
I.                hallway otide ithe                       control room and the c-ble vault/battety room lose HVAC at
                                       . sametime as the control room and experence the same temperature transient as
:tie conrol room.
: 2.                       th               r                  thie conro ro reacies 1100?                 Hhe the operators have removed W cufiit
                                     -. .:                t            es iii th ceiling .ush,Ihat:te two model volumes are now the
.g        :        :            .            lenttoe.,ylume.
t       RD2ofomiHeit oan Cfc                            lation'Assumptions a    of.. o.t6 room hiea loid tat follows, there are several explicit and a               on-,'                          '       h                ih"r                '
ti              In filtrtin ex    ti
                                                     '                     n~
itrion ad heiat liossestbro* the g are negligible or iginor o
s Thu:-:oniisl                           0         Ocessiwc by the lack'bf informabii n available t assess the
            .           ; -t    .       i               -                  exfll,
                                                                              .ion oreducti loses iniaton is conidered
                                                            .; .                      rom HVAC system ^isdcijed to be able torpride V        ,f              f      at e; .            - :;*                  vendldringvtrio 9US  -aCCdit conditdons wee excessi infi                            ion
     *j.      ;.          .                   ledk excesivediation 4oQs to th peraoa Sinn t.4i
                                                                    ..                           ,a              -le D  'PI
* e the pts for
                                                                                                                                            *=  ih ep - i   b-. -
cx-Ijrat.on ::eje ..a asme pai-s 20r~niuon ana -mitraion is asLnned to be:. .
               --.        - ne.l.               the-   'N=&#xa2;&sect; <-'6cnlain'a bZnae fori                                                     . Efj~yre ung .he ests ..w.. lw.,                                                'is w,..However, a l.0;pcr~tizi*r                           in control in                heat lonadwll be taken to account for
                                      ;on-co-Se n                              mpre=f -due to this astumpton. [14J
                                 'Whe.                 aiA66n
                  .:2             Wh         theU i             r           in te contal room ral                                100 F, the opertors take action to
      .              :         .        :      SmovC                                  ei&in-gtiles (      or     miori       JSasmd            in thin calcton)     . to           .
                                .allow akir!                                                        ir         .     -ihe66ntrb
                                                                                                                              ;                       mid vinc i'.
  .4    .   .               .-          ..                  ;                                                .                                                                                         . :
X:                 'k             ''                    '''                  ,.,                            .2s
                                       ~~~.                                                                                                                              e,.....
S.;.;

Contiroi RoomdHeax due to Loss of HVAC; VYC-1502 Rev. 0                                                          Paket_-
: 3.              fourhosfr     h                  th ]oss I    of Ii VAC, aciion is taken to mitigate the temperature rise.
I,Two posibiii ar postuated a.,               t                       atia,-is     Adiddthd brings in at least 6600 cf& of 6ztside air         .
jm tilaion                        is -provid 1hat brings i.at least 3        f of ou     e (90t *;      '     *heat  load ui t cot      rootmoii is rcduced by 52 per:ent T                    ibety-      -ee      chosen -        o theiri        toraintai n tfie contrl oom at.or.
              .         .              I i:t~
   ';v-.,';
                            '4..
                                                                                                                                      ,.              .. r .
2.*   ',,;.;-.,*.     '- *  '    .
         ,,,. or -.-
5S E
                  ;X0       t f
M L~~.                          .-   ..

* . ..                                I Control Room Hleatup due to Loss of HVAC; VYC-1502 Rev. 0                                                                                                      Page ? 3 4.0          ! Calcultion The calculation is divided into several .sections. he objecWive of the calculation is to calcuie he c.xrorom trasien' temperatre following a loss of ventilation due to an
                         -AppqcR eveintx 7pyi                                                      ch)flfigrtJon a'd tnbdel of the control room has been ol                       .eiD ion,3,i                      Twob                 pieces of i                on* are rquired to complete the jnodll contl rbom:b                                             loadand contol:rom in                            heat sink characteriation. Tests b
were             onduc              on theHVAC systei and. tie 'data th was taken can be used to chaacrimze both the heat lcadne                                           internalheat s                         .
Section 4.i                                           l          description of he data                        to calcudate the contzol room heat o                 Iad the i            td                                       tialo   thea1to f' sn                 4uses the beat load information
:aswe H thasthroolrc                            m                                                             Io- daacterize the control room
                       .het zt-riia s                                                                         -n4.3 uises;.boae                oo                        t  a          tices ontrOl rooI.
                          -.6tref6U                                   10g4an a loss                                    inoih ue o anem                                        Secl 4
                                                                   -- ::               . proun.osegiven p0             k        -.  :i        :.o t
                                                                                                                                                                          ..eij
                                                                                                                                                                          .;..I.t   ,,,-..>.
pe                                ,,;        ,N,.           .       M    :mitigatebe                              .cin longstermcorrective                  ai       o                 the.
:4.1                             'Contro R~~eat Load Calcultion T hdpw                                          :-stsre:.condu                       that obtained data that can be used to charactriei. c
              ;6&',  -     ~otroltrodrh -               ^    at16d ott h isettokI                           ,k        er~rreauc~ents
                                                                                                                              -txrx 6                           at the inlet a d"'oietf.Oweett df t               toj'- 011 M^ro                              onidu;r*                           i-rflow measaein the HVAC duct-ork df,, tio hcoli        ,A, roo,m                                                                     46        load.'-'-''
N-tei tei tas-.pe                            fm                 using           tiuxnenin              it    zonram                     norwerethey peronnidusi                                                     -       'iged t ide                                      ;the-y vvere perfizuoedn          by
                                                     .ts''a'f'< ;''.QA'e ;:.test '-procS&durg. however, cojle.yir:eendenly. I combintin, they provide a sizflienbsisM for judging thie th,       :         coiitroroom'het loaWd
                                 .                                                     y
                                    . 4i-oi"a                            -1..
is tet                        was t                                   nse tot           necessityto hA                    rze the beat load and hesitp .
                           .r:o : ?c;hfti3                                                   rom 4j1141 Iata wis'takeniililithe ctrol room HVAC M-1: tUiOiie                                                     air flow mce twn-'ta                               on -teinlet ilde of tii cot                        .,:
                .            oom.                                                   . 6               werecr tAken at ie supply.au return poifts so fo the.
ffi;-                     tion o lheat 16dk                                              n iditywl luave to 1,e esmd
                 'Di.
                   '                          '            .  . '             '"4''

Control Roomn Hcatp'p adue o Loss of HVAC: VYC-1502 Rev. 0                                              Page 1     Tl flown                       obtine forair 1.ow through the' control room:
    -AirFo16 =S59&7*fmi              10
* Reprn empi~i-re                  3. Fdb ThMown qations:-are tme irsu'lt of sipems                            n encrg balance relatio)ns. 7b yai~e wh~hat dddtoiostair. betWe~e-a'pOitsl-anlJ-2
                                                                                                                      .4
           ~ Assushg the ncom~u~ir was~at 60  wa  A          . -ih~~a
                                                           !RH'                                 es I (frp test daita), ,the enthalpieargieby W..
          *-                                                  J,
                  .. - -                O6298!
w     1~oHJroL air..si W
                         -.                     oizes.                         lb-
                                      .p.        ..                                 "
Si.                           4-
      '(S

ControI Room Hentup *Ci tl6 Loss of HVAC; VYC-1502 Rev. 0                                              Page.LS For b.2 2
4)=0.60 VA=0'0217 psia      (at t      737)
I                          p=14.7-jpsa!
14.7;-0.2413G2 h2&871U                lbdau
    'Durnn         "Js est'It ~iflle ssmdtath                         ao1oionhectolroom is neglgbe
* fj'-.. o-O-,.01f V38 W       .103
'      Ths M        th   nryde                   ote i rmhasu                    jsiiithe control room isgiveby:
                          ~!          r
                                .Crit (287 IN4 47)
                               $P;                         -2916,64BTU/ZifiU or175,007TIqhr          or 4861 BTU/se
            'Ii au or tc.otoiro                        etla         oldb      tlat+         O    ae on tile Aii          .
                                            ,-      ~
                                                                               .5 .'. 4
                          ?~-~4                              .4

w:PF LACoMtrRoom             Heatup due to Lss of HVAC; VYC-1502 Rev. 0                                                   Pagc?-
4,1.2 Heat Load Test `2 To he p onfirin te heat load that Ws calculated in the prLvious section the control room heat lqed is cate frpm data uicuded in a imneoraudurn that reported a control room HVA                i iyt              l5].             Sbment of the refece includcs data taken on the INMVAC Ssy The daaet                              s are inched in Apedix C.:
iAC                                            m ismor c~ pldx t                    e datataken for and used in the p[u. pO                    Som                            fte           t    rool              system is in order.
AiVAC i'        3   'ca          s          c view of thcotrol room VAC systen F[16]. 0 .32n                     Ihe daa irmnthse IWACeiability sd gives enougb j~f~~iotiaii~aroiundtsystem g ' t'-P''me                     -c on itio ns of -thwoid                             an and toodtnie determ    the heat oIbti~lic 3          fro               s8,-}igme                   10.12-2]
:Gmass>- ofir flRing-siro         <                bub=
the-          ioF                   51;G6 &deg;F.
603r~
          'MWv                                      .                                    *thits uifif       on. First, caculate thie.;
L04~O                                                                                   .3.
W             .'          :                        -i3          300ii               8981(f g      -.                                          - l     . ...."b -day         .1 W     . .; -1314275 5 '-
              ' -'    ~.:''-.
71:

Control Room Heap d                 Loss of HVAC, VYC-ISQ2 Rev. 0                                      Page Zf E
                                                                                            -     L
 .1:                                   .                                                      oJ3 Li C..                  ..          :       i r.4 W- ''-                  ::U        i Ciel                                  r          &deg;                              q Ai Hsl.                                  e                                               Co nt ro I
                                                                .                                                         Room PAA
                                   *..                     A rnO~~~.24+W01040 FinrW                  Contro .po          HVA Si,      chpemanati    -  of thc enteijag and leaving .
pr                                                          W
                                                             =                          Re                  a         on
         .               -.                     .'-iCt..                   1.9 K:(093O.556fW~Q.24O~-tij v.'-           ' '-
V(OW-144                                t     -^',                                                           .
      ''       ;.-      '.,       f.         -A4U                 uI         J                '       UMJ.....,;,T.-:t ', F' i ~*X,
              ,{*;            ,*
ttv        .
* 4
                                                      *1.    .-    .;:;;:bk         l   7 C,.

     - -                                                                      .1".- -
roomlm lcawop dac t oss~ of RVAC VYC-15(02 Rev. 0                              Page .za specifically, at dth inlet' t =54.47~
                                   --.0v0081374 1093 4:0.444 t- i                                    .I
      *.                             ~(.1093-03556 51.6 ;.bb813-74-O.240 (S4.4-51.6)                     IA 1093 *0.444 .54.&,,,51.6:
                                   )r~"'b240.&       (1061+~0.444r)
                                        =  O2454.4`7A5310-s'(1061 +0.44454A4)
M.
VM                                                                          10360F 103-0M6'OW-.3).240M,.20 b
                                             ..        109340444t-t      1 ~0.                          *1
   &#xa3;                       .h'.k                ''W1093      +OA444'7.I .03 I
                                        - .203.1      .&265   '10' '(161 .0A444'3.1) qjli 4

WEMCW :1.,1                                    - 9 11111M Ccoftr.o=Rv         Heatup due to Loss of fWAC; VyC-1502 Rev. 0                                        NCg Pn   Zdi and th6 total heat addition to the air stream from the control room is:
If                     qz,.im7(h 6 -h) m        668.85Tb dry airyminute h6 = 26.582BU/Th dry air 2 21;5782'TU18/d             air
                        ;q -.68.85(,6.582-21.578)                   = 3346.9 DBMtnin o, 200,815 BT1hr or 55.78 1T/sec Teimz-:ertainties in this ialue of control room heat load are not entirely known since
                 .u.ncrtaunt values 'are not directly attributable to the instruments used.
4:13        Determin'tion of Limiting Control Room Rest Load Wo v
                    *.1..TW.!    USJ f rcnilcontrom hebt     etla load have v
heen calculated using independent methods, i::.o         t tIn  oniend eonnel.                values obtained are: 175,000 BTU/hr and
                ;2.!8s 13T/lir. T a value was determined with data taken on a warm day inU                               mid-
                                       -         valuUw* obtained on a cool day in late:May. Vaiations in solar beat S :caJ  ld &#xa3;                      o          ancy an iistruiment accuracy can accunt for this diference. Eh o &#xa2; dfthese    ;ues                liaCIt joad fmcoiductior into or out of the control room thiugh the control room wall,                         ~ eat   h.. load. Thus, there is an inherent conservatism in these v auesd                 eihtir wiseg Ohuc ideli will be as electrical .l6ads (conduction and C               p.;ng explicpit iputs in the GOTHIC model input.)
T     r.,. to nswe          a   ding value for control room beat load is used, the average of the
                  - --vQues  two f10 % nid e.tho control room heat load used for dtrmining the control room r           to a loss of HVAC due to an Appendix R fie event. This Ibi e c'ntrolthe                                  information received regarding the potential electrical loAds
                       .:q
                        =17 000 + 200815S) 2                  1.1 = 206,700 BIU hr or 57A2 BlTU/sec n .'*        '               ,'

___ _  _ ~_   _  _   ZMMMU Control Room Heanip due to Loss of HVAC; VYC-1502 Rtv. 0                                                         Pagc3 4.2        Characterization of Abstract Heat Conductor As xnitioned previously, the transient temperature of the control room due to a loss of HVA: dependion twoimportant actors: te heat load inthe control roor and the interial heat asorbig               r                  This section attempts to characterize the internal heat absorbing
                ;cr            as a GOTIC heat conductor using test data obtained during a control room heaup O.-          ~test.
                ,n a hdated room, the:rinsient temperature rise is predicated on two major factors: the heat load and the heat absorbing structures. For the control room, the previous sections have identified the hiat load. 'Thus, to be able to predict tbh nransient temperature profile, the heat
:absoing struturs must be able to be modeldd; However, the coutrl room has a very cornpl4 array of beat absorbing structres (cabinets, panels, etc.) making it difficult to model with                      TIhus, a test was performed to gather data (tempeture rise versus time and
                                         -accacy.
                '.-he Io9Mb)hat can be uicd to charicterize the heat absorbing struchues.
                .,; theIn iedviosections, the control room beat load during the test is characterized. Given thi isEhea loiad and kkowi ambient conditions, it is possible to modify the GOTHIC model develod Previously to-attempt to model ti test condidons. An abstt inteal heat S;flUXt? ca~r.        i b -iriad in the GOTHIC model othe test. By vanying the surface area of
                .^abs.irthhat            coridur-itisthe                            abh       trbmto etanient'test    data and thus have a fairly good is                tivetni                   conddcormodel fbr use in laier control room heatup analyses.
The;, Id, .tes         iS~a seres of temperatures at differeft locations within te control room take
...         :.over      a          d of boif3Srnutes,: Bcae                         the o introl room is a three-dimensional structe with '," c"rr"n              d  ,''.       lize heat sorc lights, rilays,-'power pies, etc.) but the GOTHIC     - -.
                           ~lt                                 the conto room is one dmensionial.(a lumped parameter volume)
                ,it is not                                 'GOTHI          model iW matche etest results exactly. However, to adeq.te, : .barde            ii               int        hieat conducts, matcing the slope of the temperaue
                ' riewl i.nuat bmfiinL'I                                               .s                         w @f'lt offi
::ope th temperat a found that the slpe of thei temperature rise was fairly contt over the different locatioesof the tempeathm eadings. In fact, it was much more constant ta ,.the,          olute                          s betwenlo       ons.
Th'e est da -is include in Appendix D. TMe test                           data is plotted in Figure 4, Figure 5 and
    -Pigure                6 along with-les`howing the least res                            linear data fit for each of the data sets.
                 ?';Te following equaions s wused to obtain the lea,                       t squares fit. Note that since the slope is theon.ly variable of concern, only the slope is calc                      dated.

             ... '-.            S.,..,.
4 p          .'*.
                                                                                                                  *f            'I...
I I'      *. :
                     .1*
            * *I.
0 Ak
                                                                                                                                              .1 81                                                                           .
                                                                                                                                               .I 1.'.. lyE
* a 77            -.-                                                                                                                   a, M-..A                                                   76~
09
                                                                            . .   .        . lo-p           .. -,
                                                                 .              -        .                                                        II
                                              . : i     :...41P .          . .. -
                                                           #- .1    .
                                            .   .Po od.09 73"                            I'AAr 6 72 71                           A
                                                                        .1 I            I    l 0                                                                     20                      30   40
                                                  . I.
                                                                   ...                   Timle ..(minutes) 5S
                                              *            .:  .      .    .    :   . I

S...      4 I.
         'S...
9 4*      4*
a' I
r   'i
                                                                                                                                                                                                             -C:)
: 74.      -.                       5                                                                                                                   &#xa3;q-A'.
72 -                                         2,~:2*'                                                                                                 F         I
  ,4.
S..
* I1 I
       .5...     ..
                                                                     .0a                                        10.>                          20                30                              40 4.
                        - . p me (mn                .'.
/                                                                       .      . ..          .                                  .....
                      ,    . , - .. : I.: ,
i..:.     .
                                                                       ..       ..   -; -. :     ' : '..L'   -           "" ,      -
                                                                                                                                        . ...  .. o- ,  "..-  -
                                                                                                                                                                !  S..-    .5..   .. - . . .. -   I .. -

tV.
4   ..
                                                                                                        ''*..'.94-.
4 I
I :....
Vew                     t              -d         dI"errg      *1 LY1
                                                                                                                                      .(  .
B I
I F           i II I
I 0%
                                  -                          -U i
i 4
t I
I i
i I
i'    . _-                                     .
~.F;          ..       ~             ..                                 .                                                :.         :
                  -.              s.'.72_
                           -;-{;:~~~~~~.-.                            .   -. 4...-;... Tie( n
.t'.             r.   ..           .
o     51           15        0   2530 LA
             @0r;~..i"'.:"'":>             :':''-    ''I.4.
                                                             .. ;o i -i       !       '- .;            *.        .2   *
* 5

Control Room Heatup due to Loss of UVAC; VYC-1502 Rev. 0                                                         Page 3Y y =x+b where           m-
                   .                                         .b               P)~L      ).(1   XiY)(~
Ht )     (N  2 N    'Nx?-  Exd Purting the data for ithe first data set in the equation for m gives:
                                                        ,36,8so73                   25
-*.3.                                                        .i'.75X2':                                752 1'-'
_ _  _ _      --. _ _i   _  _4-           .                   :                 y?.2s13 A5.x 4              .          :-.       ..         12.. 5.8      :.     :*.
                                                                                   -22                311376 i2 ;   3 t              *1               i               ?4                     0                1524 3 ..                10 .                          76.2                 400                752 6                    2s                           76.4                 625                1910 7                   30                            76.9                 9                  2307 S                   3                            .77.2                 1225               2702
            .Su6       .03A                                                         3500               10700    .
C..
i      . 8 10700 Fm .-t140"F603.4min 8 '3500rin     2-(140mW)    2
                                                        =      0.13387Fmin
                                                                                                                  ... . '- ..  , , I .,.. ....

Cootrol Room Hcap due to Loss of HVAC; VYC-1502 Rev. 0                                                                 Page, The balance of the datacan be similarly manipulated and the results are:
Data Sit               Slope 1                      0.1338 &deg;F/mii 2                      0.1317:OF/rnin 3          1            0.08262 0Flrin 4                      O.1057.F/z&deg;in 5                      0.1 14.6'/bin
                            .6                      0.1264 6F/itin 7                 1 0 .1414L&deg;P/nin 0 F/nuin
                                                     .$ .0.1350:
eb9                             .
E.ch sl           reperaure                                     increase in a local region in the control room. As it LStrmind inpart by the at load and the intenal heat conductors in tiat local r
* such,
                             .e~gion. iSince othei tes data a given i averageheat load in the control room during the
                           . test (Section 4.1.1), th.aberage of these slopes will give a iepresentation                            of the combined a verage:iffc of all t                    *hea coidudors in the control room.
1.e Ave       A slope=0.12216 &deg;P/min t te 6vpe of the transient temperature curve for the GOTHIC model; Ibe v<t-. -:;:dificA~ors kd4i"iod to the initial input model are-d-; ors #12b: t heat sin:: Area 25,00D fP
::                             .                    l                     -  all init teperatures at 73.16 'F 6 o-.fHicients - all temperature boundaiy conditions set to 70&deg;F to
                                       , ;r fithe rielfIvly mild conditions in the rooms adjacent to the control room and thi; .utd.oi a-r ii                  atu=e (ly cloudy day based on t est data.)
C.o:ers.HeaCeis -Heatr #1 (eBTtUical oad)                                           /see based on results
                                      -f..
                                         ,mSwdoi~fi"4.n          Al;t,   .Heatr    #2  (occupants)   set to 3.25  BTU/sec      based on 30 siated r~fiq.                        eq la 390; `Ulhr [14][10].
                                     -;Volumetic Fan - i trip set to 106 seconds, off trip set to 0 seconds (no opening of fi' . .r' , '. i nesgt'ies.) ::., ,-
::'                                                          tnpseto   troIO's-  seconds, offtrip set to seconds(no opening of tr~tFS;j;           ;ceillg tiles^):                 . -
Volume                       itionstiitil - Volume I (Coitrol Room) set to 73.16 &deg;F 6, Volume 2 (Drop Cei"ing                       set to 72.1 TF to match Time 0 test data.
AircraV Tim. 0 katr        Wxtmpca  ftvc tz IWAldta.
    ',a'.' ....
                                   , ..bi 0.Wide wdriempMqWfWubi"v 1etred on T U80offtiew kmlcrcwx WM a~n udhoflonlmy a bzr. nt S~
O-;R*-qy                                    CM                               'voq   tile' wall      would aftatdkwaU XhS twlthnei b;asi2   e  t1w sde4Wo                Rai IuoA4 MP at VsIl wien proffa          could IMportulsted hu the Wo
                                        'Coadlwo,     fthtfi    oaJ66ow bedt up Is neJ1bla.
e.-
* 8i7@-'-li
 >A'-S'                                           '    -;. :'.,.d..1'

                                                 .        .      . 1 ...             -       I Control boom Heatup due 'o Loss of HVAC; VYC-1502 Rev. 0                                                                                                                            Pagc 3b Run Citol Par60                           meter - the tes lasted forT35 minutes; the GOTHIC problem will
                      - un fr6;                          nto be-lie to cover theist peniod and see any short team trends.
The GOTJHC inisft fXbles forthis r.odel -areinciudd in4Appendix A. The GOTHIC output results &#xa2;incijide on icwflche u-d the filenen CONTROL.SOI.
Ths.he,         tso Is hii                        are shown :.r7and                                                                 ieS. -The temper e plot sbow a C minitl    slsatp incease in                                   peure f"(o6v& b alower but steady increase. The initil s.~h^ ii           se inD c6~ie                                     is net vdij seen m the t d but'can be explaind by the
: . beaioftihe-'h                                      er coefflcxent ie                                   tb .code, -Irnaly, the ldifferential temperatue
:al . aS.onlt'l the heat dnucisisvery smal.                                                                    .Thisleads to a ve' y small h &#xa2;sfer'coeficint~idch
                         .                                                       is based n 'difrcnidal tlhrture (the correlAtion heat t>r.f.f :..eflieic ~.).3Ts in~itialy, very: lttlehea+/-is removed from the vapor phse in the odeL; tliiitiieatup                                                     i-svery sm*.ro an adiabatic p of the room.
A s th e romhe                         .                eea                                                                               t conductor sudraes fncreass as
<                   nf;ercl_             f                                                                                          th(e conductanceof heeatstrctur
                                     ;             siah                          nom anditigates the control room tempeatre rise. Tbe e  'al~Z            >ii0                                    -f                                         gs:co1ex.i1ciiiafed from GOTiIC~o                                                         -;. ..
oiwaum;ton-.....................................
I               ...-- -    *. - ; - .
* a T'T                                                                                 .       ,                           .u         i           the of                   5.......
                .-                         TAGa                                                                                                        of 25,6.                n 10se)*.60-sedzhi                            0.1231         OFlmini         .
g  w~AlIO5         Cnaedwtthtstsoeof.0J216'Fmi"'e
                                   &#xa3;                 y 82&deg;M 0
to.55 0fe.       At         360          subseunt-dalss    T     eut   76D swthin       1% 6'o'f    the'    desrird       ~~
                                                  -4..                                                 *-w            .-  n   O.                                           o         ;        .          .          -.
5'."I                                                 (5-               b                          ui        de
                -use g -:\   .....      ' ~.::; -.      '   :
                                             .-     ':'5-         ,...................................
7-

Control Room Heatup due to Loss of EVAC; VYC-1502 Rev. 0 Page 37
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I9                                                                        -,Ij                      4     I t'-,'~~~F~ur ''','&#xb6;,''','    -7'''          Control Roo                             .''u      Tes M:a Voum Vapor''''' ..
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I                                                                                                        . ...
..          .              Control Room Heanup due to Loss of HVAC; VYC-1502 Rev. 0                                                        Page39 43          Control Room Heatup Daring App. R Event Given the information available fim the test data, it is novw possible to adequately model the controom during 8 ioss 0ofVA.C' following an Appendbc R event. The generic control room GOHC input model descn led previously is used wvith modifications to the model mad dn the bisis of the test results describedin previous sections.
                          'The modifications and additions to the intial input model are:
1.h rm.1Conductors - #12 sbstract hdat sink Area =2,500 ft
* i          m'her nta  t       - all inii                           78 F
* Heat Triasferr;          ffcients - fbr walls x ed to outside conditions, the temperature
                                      ;;.oundsrycondhtitns set to r                tve aeag SOLAiRempers
                                      . .olcri/Hee          - eater#l (eliloa)stot57.42.BTU/se based on results
         .- ;f1:~~~
            ..                        in         Section 4.1.^3-.;
                                      - umetric Fa          -  Otrip set to 11IOF control room temp                     e in accordance with
       -.      .-                                            piously V        disc.ed
                              .O..n,                             tp     st to          00 cwntrol room tempeatwure in accordance with the:                    re1VOus>ydiscus            *              .       *
       .                 *       -        olum.ii iti i-                   Vatues:1 ;       aiia2sare settb-78 F;
                                       &Pa-U *rl;IIleters
                                                    *                 - the G               problem will rim for a four hour transient g:                        : (d400 3:.:
    .;       .          ThieGOT                Cipt      tbleisfor this indoe are incluWed in Appendix B. Te GOICoutput
               ;,i      r
                     ,ndej       t             i         ']EA                       the
                                                                                      . f:lena'ie..                Sor.
bip-e.           z        rLore         ~igrmel2                 9     The ittiismy          show that the control
   .4.        ;       .ac:onno included:inh6idel.                                   f feci of tli oeiBj oP the ceiling tiles caicearlyj be ...
                       'tan-the              s     ft   conDrl room since h      tbe eat sources are in th control   e      room.
                      ,!t      .r*old  r              y be expeted that the teipeintue seenu              bythfie operaors ad esqsipent wold b..dless than hoow                      thitetis                                       tby ceiingftotl Cocan. lmem-on.
oe
                       -experience owsth                         is a notceble differen e in                tempe lie         at floor level ad the;
                       .temeue at ceilingtIamlan heatd room.
                   .he
                                      ' es'" hk $h''- b' th                                           '          oh     eln.Cmo                     'i .'ijsn
                       '. ,', owstw f',t,,                                                                                                  d      '"
km'M-tM                                      a etdro" 4).-Y

Conftro Room Heatp duc to Loss of HVAC; VYC41502 Rev. 0                                 pape WVo ibv4 Roc  8HrM-Lca      ft HVAC GOV8&#xb6;V~on S~(AO        -Aprd199 f*         I 2          ~ContW Room Hoalup, Loss of HVAC
:.TV1         TV12 Its           ---..----    z............... . . -. . -. ,. . .
                                                                                      .14.
CnrlRoomiFiue9 He~atip, .Lcisi of HVAC Tiica=Tmin

__ .      A . .. *-                                                  Ig Control Room Heatup due to Loss of HVAC; VYC-1502 Rev. 0                                                                                       Page 'e 4.4             Control Room Heatup Mitigs ion Options As sden by the results in Section 43, during the first four hours of the control room loss of HVAb transient i1he                              pet re h the control room stays under 120 'F. However, due to the I                     eatoadn the room and the insulating effect of the thick- concrete wallsv                                                       , the steady state terpraue is eictd to be welr above 120 "F. To mitigate this temperature rise, two possiile options are explored. Any option chosen wil have to do one of two things, either singl or in combinaion. Either the heat 1o6d will have to be reduced or somie alternate ventilationlcoolin will have to be prov ded.
i i
The first option explored was to provide a means of alternate ventilaton. Given some tenipqrary venilaion usizng outside air the control r im temperature can be maintained at or L                  below 120&deg;F. ;:A GOTHIC model was used to determine the amount of air flow required.
                    .lbesecond option was to reduce the heat load and iprvide some means of alternate venfi1tion (at-some reduced level fromntthe first option.)
                   .be          asic f'GIOTM codtrol room model &veloped .previously was used wit some or n odiIcations.i schtic of the newGOTHIC model is 6hown in Figure 10. As seen on rA:ie.shnac, an'dditionai flow boundary coilition id a flow path to that boundary conldition were ded.                                      ls th' mifions-are outlined below for each optioiL Option i- The inputftables for this option are included as Appendix E. The output for this u'iLs! insluded onmicrofiche uder'thte fiename nOPPTONN .SOT .
                  <;, -Oit2oh2-           TIe input taibles for dii option are included as Appendix F. The output for this is insl'dad on microfiche under the idenazr 4t ~OPTION2.SOT".
   .   :                            Fluid Bo                      .Conditions - a flow boundiry condition was added to simulate a eci:s:ant.,'ntila<flon air flow;                                            nthis case, the flow was set to a function ibich had zewoflow upt&'4400 seconds (four hcurs) and 110 cfs (6600 cfi for Option 1) or 56.cfs -30009cf;for Option 2)therea                                                        te ' wd-iswactualyregativne, ito th bo      ,iM           o                    lin                        the control room press           boundary condition to provide
     ^  .                 .          -theWei   .
Flow Pah-;aflow path was modeled connecting tbe new fluid boundary condition
             *t                          the c                         o            the flow path is not niendedto model any particular
                                    .vea                       iiction and for s-mplici was                                v sct io the same parameters as the other control                          mdmy condition connestn flow path';
CoolIrfeateri/ - a forcing fiunction was assigned to the electrical heat lead; for Opion I, tifing                            fimction was unity (no c                                 'inheat load) but for Option 2, fiorciggthis     fntion was set to 0.48 at four hours to represent a reduction in heat load The tempat                      tsn                  results of these models re shown in Figure }I and Figure 12.
PSI..::                                                                                .                     :   :         .
                                                                                                                                                                                        . P t  4  r.,     .?-,-     "wt  '   ;*r    ' :-!..
                                                        ''ti
:-             :'   '   .................................                                                          . :, .. . . Z;

Control ikoom Hcatnp.4tae to los of HVAC: VYC-I502 Rev. 0                    Page .L VY   frol ROOM HmdAtup-osW I AC OPtSo I UmAD28 14:2020 1998 2MI    frsonl 5.V,(QA)-C -AOI 99 10A.                     IV5p.
. -I.r
                                            .~- r~~n~pu          ~      ...                 C
                                                         .O              2 2tM    *                                           .4:a vb2Md v.I----                           -

Control Room~ Heamup duco t Los of HVAC; VYC-I 502 Rey. 0                                 Page~~
* I              olF oaito.
M         Los* ofHVAC.Opgton 2'!.
32.0               IS     1A         Ia    2.
2'          F6n im
                                      *3.       7..2             1A         1a1.. 21.8
* FigireV
                 .            Contol oam eatu Re                        Otion 2
                                                                       'lts        -
I      Na~W~AC~cPZ 2~:         C~oIA~m~I~Heat                         Load

1-IA~
Control itoom Hcit                            duc6i Loss of HVAC; VYC-1502 Rev. 0                                                        Page   '
5.0                     lConcdiMous
 ...                                It has beei shown-hi.forthe fixtt four hours aftera control room loss of HVAC tmsient, 2?                                  i."m~initol
                                     .su.fi.cien+/- ced ing                             -,
WS          -
                                                                                                  -room te er pa not rea
                                                                                                                           *x.200 2given         opertor actontoremove 110F -To p m the co tol room tp                                 f.                d.iig l2Os                     o, sorm wfourctdon is requied to itigate the 0 . ;tion-h           ai' been idht tia which ensure control room
:;t>
d.e dedtr                        120I in a long tiemi Wenti:.
             .                      a                                            ventiltionproviding at least 6600 efin of outside air, or
* D.                  :-;iS             3 rewdui               by at leat 52% accpanie by at least 3000 cfm of outsde air.
b                '           :    o*d                                                              eas....t            ,ar.                   .*. ",X 511~
W-1-*                                                       I -;
Wlft;*                                               I MS ',*@
          ,*->  Ss                .                   '
            *N ,- 'vfi S@,
m.*                                    *-'-.,21  sZ v-    -A-           ..'. .- '1'^''~'.......................'-,....
             ';L'...............
m                       *;                  ~~~~~~~-                                                              t'''
                                                                                                                                                                                                  &#xa2;r 4''.
w--S                                  e                                           b;e                    **->?                 -*e,,                                           ,          *
     ,+/-~
       -'                                      >                                   i-    -  ...-                                  *
::*1'..-:,;;,,-_..k-......=r
                                                                                                                   - -                                                                      C.-r

BVY 05-072 Docket No. 50-271 Exhibit SPSB-C-52-4 Vermont Yankee Nuclear Power Station Proposed Technical Specification Change No. 263 - Supplement No. 30 Extended Power Uprate Response to Request for Additional Information Calculation VYC-2405, Rev.0 Total number of pages in this Exhibit (excludina this cover sheet) is 85. 1

CALCULATION COVE1R PAGE O IP-2            f    IP-3           El 1A}            []PNPS         ED VY Calculation No.             This revision incorporates the following MERLIN    Sh   1 f VY 41Re.0                DNs     or Minor Calc Changes:                     Seet I of 85

==Title:==
Dryw6lI Temperature Calculation for a Station                                          ZQR   EINQR Blackout Event at Extended Power Uprate.
Discipline: Fluid Systems Design Engineering                  Design Basis Calculation? ZYes   []No This calculation supersedes/voids calculation: N/A Modification NolTask No/ER      No:  EPU EI        No software used 0        Software used and filed separately (Include Computer Run Summary Sheet).
If BYES', Code: GOTHIC V7.0p2 O        Software used and filed with this calculation. If "YES', Code:
System NoIName:
Component NoJName:
.(Attach additional pages if necessary)
                      '      '                            Print/Sign REV #        STATUS             PREPARER                REVlElENNR/           OTIEER       APPROVER        DATE (Prel, Pend,                                 DESIGN            REVIEWER/
A, V, S)                                    VERIFIER             DESIGN
                                    -    .                          .        VERIFIER o            Prelim         Liliane                  Alan L.                             James G.
Schor                    Robertshaw                          Roge 7 t/:&A                  7/C                           A ENN-DC-126 REV.A ATTACHMENT 9.2                                                 CALCULATION COVER PAGE ENN-DC-126 REV, 4 ATTACHMENT 9.2                                                CALCULATION COVER PAGE

  -Entergy                Calculation VYC-2405 Rev. 0          Page 2of 85 ae f.
ENN-DC-126 REV. 4 ATTACHMENT 9.6                   CALCULATION RECORD OF REVISIONS RECORD OF REVISIONS Calculation Number: VYC-2405 Rev 0                             Page    2    of   85 Drywell Temperature Calculation for a Station Blackout Event at Extended Power Uprate.
Revision No.           Description of Change                  Reason For Change 0             Original Issue                        N/A

Entergy                 Calculation VYC-2405 Rev. 0                Page 3 of 85 ENN-DC-126 REV. 4 ATTACHMENT 9.4                                CALCULATION 

==SUMMARY==
PAGE CALCULATION 

==SUMMARY==
PAGE Calculation No. VYC-2405                                                     Revision No.0 Drywell Temperature Calculation for a Station Blackout Event at Extended Powver Uprate.
CALCULATION OBJECTIVE:
This calculation will address the VY drywell temperature for Station Blackout (SBO) at Extended Power Uprate (EPU) Conditions. The calculation will look into means to mitigate the drywell temperature for this event, such that there will be no need for Emergency Depressurization.
CONCLUSIONS:
See Section 7.0 ASSUMPTIONS:
See Assumption in Section 4.0. (see also list of open items - assumptions which need verification or implementation - Section 4.1)
DESIGN INPUT DOCUMENTS:
See Design Input Documents identified in References Section 8.0 AFFECTED DOCUMENTS:
See Assumption Section, Section 4.2 METHODOLOGY:
See Section 3.0

    - Entergy                         Calculation VYC-2405 Rev. 0                                           Page 4 of 85 TABLE OF CONTENTS Cover Sheet (ENN-DC-126 Attachment 9.2)................................................................                                                 1 Record of Revisions (ENN-DC-126 Attachment 9.6).................................................................                                       2 Calculation Summary Page (ENN-DC-126 Attachment 9.4) .......................................................... 3 Table of Contents ..........................................                                                                                            4 List of Effective Pages (ENN-DC-126 Attachment 9.15) ........................................                                                          5
* 1.0 Background.6 2.0 Purpose.6
* 3.0 Method of Analysis.7 4.0 Inputs and Assumptions.7 8.0 Backgr oun d............................................................................................................. ;...........................6 5.0 Input and Design Criteria .........                                                                                                                  15 6.0 Calculation / Analyses .........                                                                                                                        ;26 7.0 Results and Conclusions .........                                                                                                                       ;;67 Computer Run Summary Sheet (ENN-DC-126 Attachment 9.10)..71 Calculation In p          act Page (ENN-DC-126 Attachment 9.7)7..............................................72 Review Calculation Design Verification and Review (ENN-DC-134) .............................................                                            77 Files on CD.............................................                                                                                          85 Attachments Attachment A.Al                                                                                                                          - A39 TOTAL NUMBER OF PAGE (Including Attachments) ......................................                                                                 124

6 f .. . .) i w s w w w
 -Eniergy               Calculation VYC-2405 Rev. 0              Page 5 of 85     -
ATTACHMENT 9.15                                                   LIST OF EFFECTIVE PAGES LIST OF EFFECTIVE PAGES Calculation Number: VYC-2405                   Revision Number: 0     Page 5 of 85 Torus Temperature Calculation for.a Station Blackout Event at Extended Power Uprate.
PAGE        REV.           PAGE           REV.        PAGE     REV.
All        0

  -- Entergy                 Calculation VYC-2405 Rev. 0                 Page 6 of 85     b'

===1.0 Background===
The Station Blackout (SBO) torus temperature calculation (Reference 1) accommodated a higher coping time of 2 hours versus 10 minutes previously assumed. In addition to an increased coping time, Reference 1 also eliminated the potential need for Containment Overpressure (COP) for the SBO event. In the process, it was determined that in order to implement coping strategies for the two hours, two additional parameters need to be analyzed:
     -   Drywell temperature and the coping strategy to accommodate an expected higher drywell temperature, and
     -   Procedural direction for the operators (if needed) to limit the drywell temperature while ensuring capability of HPCI/RCIC to maintain vessel level 2.0 Purpose This calculation will address the VY drywell temperature for. Station Blackout (SBO) at Extended Power Uprate (EPU) Conditions. Th'e calculation will look into means to reduce the drywell temperature for this event, such that there will be no need for Reactor Pressure Vessel.
Emergency Depressurization (RPVED Reference 30).
This analysis will address control of the drywell temperature by controlled depressurization (cooldown) and will show that RCIC/HPCI injection is maintained until power is restored and the low pressure pumps (RHR and CS) are available.
As indicated in Section 1.0 of Reference 1, for the SBO event, the Alternate AC (AAC) power source is restored at 2 hours into the event. After the restoration of power, torus cooling and drywell spray will become available.
2.1 Acceptance Criteria To evaluate the results the following criteria are applied:
: 1. The maximum allowable drywell bulk average temperature should remain below the EQ temperature (340TF for the first 30 minutes and 325TF for the next 270 minutes)
(Reference 19).
: 2. The maximum allowable drywell surface temperature is 281 F (Reference 20).
: 3. The maximum allowable drywell air pressure is 56 psig, (Reference 27).
4 Maintain the torus pressure below PSP curve (Reference 30) during the 2 hour coping duration and the 10 minutes of low pressure pumps restoration period.
: 5. The analyses should provide assurance that there is no need to spray the drywell in the unsafe region of the DWSIL curve (Reference 30).

s +   - t  D D w w ffi Entergy                   Calculation VYC-2405 Rev. 0             Page 7 of 85 3.0 Method of Analysis The model developed in Reference 1 is modified to accommodate changes related to the purpose described In Section 2.
The GOTHIC code (Reference 7), Version 7.Op2 has been selected for use in this analysis. This code was used in the original suppression pool temperature calculation (Reference 6) and in the analysis for SBO at EPU conditions, Reference 1. This specific version of the code has been installed and complies with the ENVY SQA procedures ENN-IT-104 (it replaced VY procedure AP-6030) as documented in calculation VYC-2208 (Reference 8).
The following changes to the input SBO-NoLeak-80 to produce SBO-drywell2 are being added:
      -    drywell heat load
     -     drywell heat slabs
     -     leakage from drywell to wetwell
     -     modifications to the vacuum breaker modeling The GOTHIC input file for the case SBO-dryvell2 is presented in attachment A.
4.0 Inputs and Assumptions The inputs and assumptions for the SBO event were developed in Reference 1. For completeness, they are added to this calculation. The more important modifications to the model have been made, for this analysis, by the addition of the Drywell Heat Loads and Drywell Heat Slabs (see Section 5.0 for details).
The SBO scenario postulates a complete loss of onsite and offsite AC power. The vessel is assumed to be isolated at the start of the event.
The scenario is modeled as follows:
: 1) Scram occurs at time zero.
: 2) The MSIVs are isolated at time zero (this is a conservative assumption for the drywell temperature calculation since the energy transferred to the condenser while the MSIVs are opened will remain in the vessel).
: 3) The Reactor Vessel level is maintained with HPCI or RCIC in a band between 127-177 inches above Top of Active Fuel (TAF). Level is maintained with HPCI at a nominal flow of 4250 gpm. In reality HPCI flow will be adjusted to keep level in the band and to prevent excessive start/stop cycles. The HPCI (or RCIC) modeling in the GOTHIC input as a continuous flow (lower flow) or as intermittent flow has no effect on the drywell temperature analysis results.
The choice of RCIC or HPCI or the flow capacity has no effect on the analysis since HPCI injects intermittently to maintain inventory or can be throttled as required to maintain level.
If RCIC were used, it would inject more often.

Entergy                  Calculation VYC-2405 Rev. 0                Page 8 of 85    NV
: 4) HPCI takes suction from CST at 135 0 F. The CST inventory available for injection is 75000 gal.
The GOTHIC input value in Ibm = 75000 gals
* 1 ft3 n.48 gals /0.01627 Ibn/ft3 (Reference 31 for the density a 135 OF) = 616271.7 ibm
: 5) Power is restored at 2 hours. Torus cooling is initiated at 2 hours and 10 minutes. Two RHR Service Water pumps are available at 2 hours and 10 minutes, delivering 4700 gpm. The second RHRSW pump is discontinued at 16 hours in the transient to maintain the Corner Room temperatures below the EQ limit (Reference 1, Attachment B). The drywell temperature analysis is performed for only 25000 seconds for the base case and for 14400 (4 hours) seconds for the sensitivity cases since, after 2 hours and 10 minutes (7800 seconds),
the low pressure pumps are available to spray the drywell, if needed, hence there is no need to analyze the drywell temperature for a longer duration.
: 6) An orderly reactor cooldown is initiated at one hour in order to maintain the drywell.
temperature below the EQ limit (Reference 19) and the drywell shell metal below 281 TF (Reference 20). Two cooldown rates will be analyzed: 800 F/hr and 45TF/hr.
: 7) The -RPV level is controlled by HPCI until the CST is depleted or HPCI shutoff pressure is reached. When the pressure permissive is reached, one Core Spray pump starts (after 2 hours and 10 minutes) to inject into the vessel . After the level is recovered in the normal range, the Core Spray system is used to maintain the level with the vessel pressure being controlled by an SRV cycling between 50 and 100 psig. The suppression pool is cooled continuously by the RHR system. The reactor vessel is maintained in this configuration. The RHR pump in torus cooling is also available for drywvel spray after 2 hour and 10 minutes.
.8) The HPCI turbine takes steam from the vessel to provide its motive power. It returns the exhaust steam to the torus. The steam to the turbine is not modeled since the model assumes SRV opening and closing to maintain pressure. Any steam not removed by the HPCI turbine will be removed through the SRV. to maintain a certain pressure. The total flow through the SRV is increased, but the details of SRV flow are not important for this applications and the two (SRV flow and HPCI turbine) can be combined for model simplicity.
: 9) The liquid leak is modeled as a fixed flow of 8.4585 lb/sec (61 gpm, Reference 3) [(61 gpm
     /60 s/min /7.4805 gal/ft 3 /0.0161 ft3 Ilb = 8.4585 lb/sec)] and it stays on for the entire transient. (Analyses will be performed with and without leak for one depressurization (cooldown) rate: 80 TF/hour). In reality, the leak is variable depending on pressure. Assuming a density of 62 lb/ft3 and fixed flow is conservative for the drywell temperature analysis.
Analysis. of drywell temperature for a 45 TF/hour cooldown with no RPV leakage was not performed because for the case .with 80 TF/hr cooldown, for the period of interest the drywell temperature stays below 300 0F for both cases (with and without RPV Leakage). For the 45
     *F/hour cooldown, the temperature in the drywell for the no-leak case is expected to remain below 300 TF as in the 45 TF/hour cooldown case with RPV leakage for the analysis duration.

 -  Entergy                  Calculation VYC-2405 Rev. 0                 Page 9 of 85
: 10) The analysis will assume a fouling of 0.0018 in the tubes and 0.0005 in the shell. This corresponds to an overall RHR Heat Exchanger (RHRHX) fouling of:
R. = Rfi (d) + Rfo where:
Rfi and Rfo = tube and shell fouling factors, respectively (hr-ft2 -&deg;F/Btu) do = outside tube diameter (in) di= inside tube diameter (in) do = 0.625 in (Reference 22) di= 0.527 in (Reference 22)
Rfo = 0.0005 hr-ft2 -FJBtu (Reference 22) 2 Rfi = 0.0018 hr-ft?-F/Btu    (from 0.0020)
Overall RHRHX fouling 0.625 Rf =0.00l8*(o 62)+0.0005=0.0026 0.527 This number compares well with the maximum fouling calculated in Reference 23 of 0.002307 and 0.002445 for the RHR HX E14-lA and RHR IIX E-14-IB, respectively.
: 11) A variable SW temperature is used, consistent with Reference 1.
      *Since this change of depressurization (cooldown) function of service water temperature requires procedure changes, it is added in Section 4.1 as an unverified Assumption.
      - For SW> 75 0F, depressurize the vessel at 80 &deg;F/hr or higher.
      - For lower SW temperature (SW < 75TF no restrictions on depressurization) rates.
: 12) - Various assumptions made concerning the added Heat Conductors:
      - The heat load decreases linearly when the temperature difference between RPV and drywell becomes smaller.
      - For all conductors, only heat conduction is conservatively assumed in the air and concrete layers.
      - The outer surface boundary condition is conservatively assumed to be adiabatic (i.e., the heat transfer coefficient is set to zero)

 ---Entergy                   Calculation VYC-2405 Rev. 0               Page 10 of 85 . -
Table 1 - Vessel and Core Initial Conditions and Parameters, Primary Variables Parameter                  Nominal         Analysis        Basis
_                          Value           Value Initial Reactor Power      1912 Mwth       1950 MWth       100% power +2 % uncertainty (per NEI-87-001, SBO can be performed at 100%
power, however this analysis used 102%
power, consistent with CLTP and the Reference   I analysis).
Core Decay Heat            ANS 5.1         ANS 5.1 +2 ca   ANS 5. 1 1979 standard+2 a uncertainty
                                                           -(Reference 24)
MSIV closure time          3.0-5.0 sec     0.0 sec (MSIVs  Minimum value allowed retains the not modeled)    maximum energy in the vessel.
RPV Pressure                1015-1025 psia 1045.2 psia     Higher value, conservative, maximizes (Reference 28)                  the vessel energy.
Initial Vessel Level       162 inches      172 inches      Analysis value conservatively accounts for 3 inches increase above normal (uncertainty and operational fluctuations) and 7 inches for dimensional uncertainties. These assumptions are LOCA assumptions and are judged conservative for SBO.
Core Flow Rate             48.0E6 lb/hr    51.36e6 lb/hr   Includes ICF of 7%.
Initial Feedwater Flowrate 7.876e6 lb/hr   8.076e6 lb/hr   Reference 25 (TE 2003-20)
Initial feedwater          393.5-393.6 OF  393.9 OF        See discussion in Reference 25.
temperature                                                Feedwater is tripped at time .0, due to SBO. The feedwater is used only for the steady state initialization.
SRV Cycling                1080-1047.6 psi 1080-1047.6 psi The setpoints for the SRVs are nominal.
(between RPV    (between RPV    No additional as found allowable of 3% is and Drywell)    and Drywell)    added since it will have no effect on the drywell temperature since the SRVs open to remove the decay heat and, until the depressurization starts, indifferent of setpoints, the SRVs will cycle to remove the decay heat.
The operators will take manual control of the SRV and will cycle between 800 and 1000 psig (EOP Reference 37) to reduce the numbers of times the valves cycle. There is no effect on the calculation since the valves in any operational mode will open to remove decay heat.
Vessel Leak                61 gpm          61 gpm          A constant 61 gpm leakage is assumed; (Reference 3). The analysis will be performed with & without leakage, since the drywell temperature will have a different profile for the cases with no leakage.

v . w t t. o s A. . S \ \ ^
i- Entergy                   Calculation VYC-2405 Rev. 0                   Page 11 of 85           N:
Table 2- ECCS Initial Conditions and Parameters Parameter        Nominal         Analysis Value Basis                     Comments HPCI flow rat    4250 gpm        4250 gpm            Tech Spec Flow       Since the flow is intermittent there is (Reference 2)        no need to use the min flow of 3570 gpm (uncertainty added) (References 2 and 5). In reality the HPCI flow will be adjusted to maintain level to l                      prevent excessive pump stop/start.
HPCI pressure    1135-165 psia                        1135-165 psia       Reference 5 and 27. HPCI is shut off range                                                                     if vessel pressure drops below 165 psia.
CST Temperature  120 OF           135 OF                                  OPEN Item CST available    75000 gallons   75000 gallons       Available CST        Per Reference 3, the Tech Spec value inventory        (VY Tech                            inventory for HPCI   can be used.
Spec -                              injection            An administrative limit for the CST Reference 2)                                             level of 25% is required.
Core Spray Flow  Curve of flow   Same as             The core spray       The Core Spray System will be used vs. vessel-      nominal.           flow rate used in    for level control only after the CST is torus AP.                           the SBO analysis     depleted and/or the low pressure is of Reference I will  reached.
be used. The flow rate is determined   OPL4 -Reference 5 as a function of the vessel-torus AP.
(consistent with the LOCA analysis)
RHR Flow         7000 gpm        6400gpm             6400gpm used in      Consistent with Reference I (t=7800 seconds)                                     analyses             limiting case and Reference 5.
RHR Hx Fouling   0.0005 shell,   0.0005 shell,                            Assumption input #10, supported by 0.0018 tube     .0.0018 tube                             Reference 23.
RHR Hx Tube      NIA             5%                  Allowable            Design value providing margin above Plugging                                             plugging margin      the current plugging value of 3.6%
RHRSW Flow       4700            4700 gpm (2         4700 gpm             (Reference 4)
RHRSW pumps) 4700 gpm =
650.98 lb/sec (at 85 OF)
RHRSW Inlet      32-85 OF        Variable, see       If SW is > 75 'F,    Maximum Allowable Service Water Temperature                      assumptions,        depressurize the     Temperature (Reference 2) only for based on            RPV with rates 80    depressurization rates > 80 TF/hr 0
depressurization      F/hr or higher. For lower depressurization rates the (cooldown) rate                          SW has to be below 75 'F. This requirement is derived from the torus temperature calculation (Reference 1). The rate of depressurization was shown in this calculation to have minimal impact on the strategies to control the drywell temperature for
                                              ._                             SBO.
'I

'FEntergy                     Calculation VYC-2405 Rev. 0                      Page 12 of 85           -
Table 3- Primary Containment Initial Conditions and Parameters Parameter          Nominal Value       Analysis          Basis                   Comments Value Drywell            110-170 OF          170 OF            Reference 5             The highest drywell Temperature                                                                      temperature is used.
Drywell Pressure   16.4 psia           16.4 psia         VY Tech Spec (Reference 2)
Wetwell            88 OF               90 OF             Maximum Tech Spec       A 2.F uncertainty is applied Temperature                                              Value (Reference 2)     via procedure to account for instrument uncertainty (Reference 26)
Wetwell Pressure   14.7 psia           14.7 psia         Normal Torus operating pressure (vented to atmosphere via Standby Gas Treatment System)
Drywell Humidity   20 -100 %           100% (base        Nominal Values:         Use maximum drywell case)             VY UFSAR                humidity consistent with (Reference 27)          Reference I for the base Sensitivity                               case. Sensitivities performed at                              performed at 20% drywell 20% humidity                              humidity.
Wetwell Humidity   100%                100%              Nominal Values:         Minimal to no impact on the VY UFSAR                SBO drywell temperature.
(Reference 27)
Wetwell Water      68000 ft            68000 ft3         Minimum Tech.
Volume                                                   Spec. Value (Reference 2)
Drywell free       128,370 -131,470    131,470 ft        Reference 5             The maximum value in volume             ft3                 (includes                                 OPL-4A is used for vents)            Consistent with         SBLOCA, IBLOCA and Reference 6, the        Small Steam Breaks.
The values        volume of the           Vents: 16703 ft2 (VYC-proposed are      drywell side of the     2306 -Reference 32) consistent with   torus-drywell vacuum OPL4A             breakers of 372.3 ft3   Total Drywell Volume =
will be added to the     131470 - Vents Volume +
proposed value.         Drywell side of Vacuum Breakers = 131470 -16703 +
372.3 = 115139.3 ft3 Wetwell free       For the minimum     Nominal           Reference 5             The value at Dp>0 of volume             water level of      Values used. I       _     _      _     105,932.0 ft3 is used for a

'Entergy                        Calculation VYC-2405 Rev. 0                   Page 13 of 85 Parameter           Nominal Value        Analysis        Basis                  Comments Value 68000 ft3 , the                      Consistent with        total volume of 105,932 +
wetwell free         The values      Reference 6, the       68,000 = 173932 ft3 volume is            proposed are    volumes of the 107,104.8 ft3 for   consistent with drywell side of the    Used in calculation: 173932 Dp =0.0 and          OPL-4A          torus-drywell vacuum   + 99.4 = 174031.4 ft2 105,932.0 for                       breakers of 99.4 ft3 Dp>0.0 where Dp                      will be added the is the pressure                      proposed Value.
difference between drywell and torus.
Vacuum Breakers-    0.5 psi              0.5 psi         0.5 psi                Reference 2 pressure difference between wetwell and drywell for vacuum breakers to be fully open Dr-ywell-to         Max allowable        Base case                              Reference 5 for the max Wetwell Bypass      area =0.12 ft2       =0.12 ft2                              leakage, Reference 33 for Leakage                                                                         Tech Spec Allowable
                    .Tech Spec Allowable Sensitivity
                    =0.0033 ft2          =0.0033 ft2

AEntergy                     Calculation VYC-2405 Rev. 0                 Page 14 of 85 4.1- Assumptions that need Verification or Iniplementation
: 1) Two (2) hour restoration of outside power (coping time).
: 2) TVn (10) minutes to start RHR flow through the RHRHX, and the use of 2 RIIRSW pumps and CS.
: 3) Acceptability of using 75000 gal from CST (change of level setpoint).
: 4) Maximum CST temperature of 135TF.
: 5) The depressurization rate function of Service Water temperature needs to be verified and proceduralized as follows:
For SW> 750 F; depressurize the vessel at 800 F/hr or higher.
          - For lower SW temperature (SW S 75 0 F) no restrictions on cooldown rates.
4.2 Affected Documents
: 1) DBD - for Residual Heat Removal - change the maximum tube side fouling resistance from 0.002 hr 2 -f'F/Btu to 0.0018 hr-ft 2 -&deg;FfBtu as well as the total fouling.
: 2) Change the description of the SBO event in the DBD for Safety Analysis.
: 3) Change all DBDs and documents that address the SBO coping time (identify and modify).
: 4) Change DBD Containment Pressure Suppression System to incorporate results of this calculation.
: 5) Review following documents for need of modification: VY UFSAR, and PUSAR.
: 6) Modify SBO procedure (OT-3122-Reference 36) to incorporate cooldown at one hour and provide guidance to the operators such that RPVED is precluded based on the results of this calculation.
Note: Section 4.1.& 4.2 items are being tracked via LO-VTYLO-2005-00135.
I.

Entergy.               Calculation VYC-2405 Rev. 0             Page 15 of 85     -- :
5.0 Input and Design Criteria The GOTIJC input from Reference I is modified to implement the features described in this section. The modified input is called SBO-dryvell2.
The main features added to the SBO model are the drywell heat load and drywell heat structures.
A schematic of the system modeled is presented in Figure I
                                                    *Orywefl L' ,/van WetwaC.
bxus Weat servlce Water Figure 1 - VY Containment and the Associated Systems Note: only RCIC pump is shown in this simplified model. Actually, HPCI is assumed to inject.

Ad Entergy                  Calculation VYC-2405 Rev. 0              Page 16 of 85    '
5.1 Drywell Heat Load Calculation The Drywell Heat Load Summary at Current Licensed Thermal Power (CLTP) is summarized in ReferenceJ1O. The total amount of heat given to the drywell at CLTP is 1,691,300 Btulhr. The drywell heat load was recalculated in Reference 18 for the Extended Power Uprate (EPU) as 1,700,675 Btu/hr. Since the Extended Power Uprate is performed at constant pressure, only the feedwater pipe and valves will be at higher temperatures (Reference 18), hence a higher Q for this component; (124,000 Btu/hr -Reference 10 versus 133,375 Btu/hr at EPU -Reference 18) is calculated.
The total power to the drywell for EPU is presented in Table 4.
Table 4 Drywell Heat Load (Reference 18)
Item      Component                                           Cooling Load No. I                                                      (Btu/hr) 1         Reactor Vessel                                      459,000 2         Recirc. Pumps, Valves and Pipe                      278,000 3         Feedwater Pipe & Valves                             133,375 (EPU Modified-Reference 18) 4         Steam Pipe & Valves                                 212,000 5         Condensate & Instrument Lead Lines                  82,000 6         Control Rod Drive Pipe                              50,400 7         Clean-up Pipe & Valves                              17,800 8         Shutdown Supply Pipe                                8,100 9         Steam Safety/Relief Valves                          206,600 10        Biological Shield (Gamma Heating)                   16,400 11        Safeguards System Piping                            82,000 12        Steam Leak                                          155,000 Total                                                      . 1,700,675 Btu/hr

Entergy                 Calculation VYC-2405 Rev. 0                Page 17 of 85        -
5.1.1 Drywell Load Modeling The drywell heat is modeled as two heater #45H and 4#6H. The heater 5 represents the heat source which varies function of the liquid temperature in the vessel, while heater 6 represents the heat source which varies as a function of the vapor temperature in the vessel. See explanation of these two heaters in Section 5.1.2.
The heat loads which are exposed to the steam atmosphere (for Heater #611) are:
Table 5 Heat Loads Exposed to Steam Item No.'     Component                                        Cooling Load I_                                                (Btu/hr) 1            30% of Reactor Vessel Heat Load                  459,000*0.3 = 137700 4             Steam Pipe & Valves                              212,000 9             Steam Safety/Relief Valves                       206,600 12           Steam Leak                                       155,000 Total                                                          711,300 The normal level is at about 0.3 of the total vessel height. From Reference 34 the distance from the 152 inches above TAF to the top of the vessel is 21.432 ft in the GOTHIC vessel model and to the vessel bottom is 41.193 ft.
The middle range of 152 inches is calculated as (177 inches + 127 inches)/2 = 152 inches.
The model assumes a normal level of 172 inches (Table 1) which is 20 inches above the 152 inches, hence from 172 inches above TAF to the top of the vessel there are -21.432 - 1.667 =
19.765 ft The liquid height = 41.193 + 1.667 ,42.86 ft Total GOTHIC vessel height = 62.625ft (from Reference 34 =330.542 - 267.917 =62.625 ft)
Steam region = 19.765 /62.625 = 0.31 (used 0.3)
Heater 6 load = 711300 /3600 = 197.58 Btu/sec Total Heat load = 1,700,675/3600 = 472.41 Btu/sec Thus, Heater 5 load = 472.41 - 197.58 = 274.83 Btu/sec

I Enlergy                   Calculation VYC-2405 Rev. 0                Page 18 of 85     NY; 5.1.2 Transient Heat Load Behavior It is assumed that the heat load decreases linearly when the temperature difference between RPV and drywell becomes smaller. In order to calculate the transient heat load, the following transient heat load pocedure is used
: 1. When the difference between the vessel temperature and the drywell temperature is less than or equal to zero, the power of the heat source is zero.
: 2. When the temperature difference is greater or equal with to Tset, the heater power will increase above the nominal value.
: 3. When a temperature difference exists between Tset, and zero, the power is linearly interpolated between the nominal value and 0.0.
T3w is defined in the way that the calculated power of the heat source is equal to the nominal value at the beginning of the transient. Two GOTHIC control variable CV 41 and CV 42 arc defined as the temperature difference between the vessel internal water temperature and the temperature inside the drywell (CV41) and between the vessel steam temperature and the temperature inside the drywell (CV42), respectively. The model shows higher steam temperature than saturation because of the heat slab exposed to steam which represents the heat structures in the Reactor Pressure Vessel (RPV) exposed to a steam environment. Sensitivity studies which placed this heat structure in liquid eliminated the steam superheat, as expected. This. is a conservatism of the model. In reality all structures will be exposed to Tsat = Tliq= Tvap The control variables are used as the independent variable of the functions, which gives the transient heat loads to the drywell, as described above.
5.2 Drywell Thermal Conductor Model Development The following thermal conductors are being added to the model.
There'are several types of heat sinks and thermal conductors inside the drywell. The components included as heat sinks are the metal mass of 4 RRUs, vent pipes and the drywell liner.
Miscellaneous steel exists in the drywell, but has not been previously quantified in detail.
Minimum heat sink components are considered conservative; therefore, miscellaneous steel is not included as heat sinks.
Drvwell liner divided in (Reference 21):
l) Lower Drywell spherical portion,
: 2) Upper Drywell cylindrical portion, and
: 3) Drywell head.
The drywell wall consists of the concrete, the inner surface steel plate and the air gap. Zero heat flux boundary condition on the outside surface of the drywell wall is used.

Enterqy                     Calculation VYC-2405 Rev. 0                   Page 19 of 85 The data on OPL-4A (Reference 5) is used to model the steel liner. The surface area calculation for the liner was performed in Reference 9.
N                         Table 6 - Drywell Steel Liner Elevation          Steel Thickness (in)      Surface Area (ft2 ) GOTHIC Item #-from Ref. 9                                                                          thermal conductor No.
2, page 47 of Ref. 9       El. 237.74'-El     1.0 page 47 of Ref 9      1856.24             5 257.75'          l 3,page 47 of Ref. 9         El. 247.24'-EI    0.8125 page 51 of Ref 9   2041.28             6 257.75' 4.1, page 47,48 of Ref. 9   El. 257.75'-El. 0.6875 page 51 of Ref 9   1250.47             9 259.92' 4.2, page 48 of Ref. 9     El. 259.92'-E1     0.6875 page 51 of Ref 9   3802.73             7 283.69' 5, page 48 of Ref. 9       El. 283.69'-EI. 2.5 page 51 of Ref 9      780.68              8 289.61' 6, page 48 of Ref. 9       El. 289.61'-El     0.635 page 51 of Ref 9    1898.24             10 308.00' 7, page 48,49 of Ref. 9    El. 308.00'- El    1.25 page 51 of Ref 9     1114.72             11 318.50'                                                          _
8, page 49 of Ref. 9       El 318.50'- El     1.25 page 51 of Ref 9     783.4               12 327.75'                                                                   l 9, page 49 of Ref. 9       El. 327.75'-top    1.3125 page 51 of Ref 9   1718.3              13 of drywell                                I Total                                                                         15246.06 ft' The items 2 through 7 have 0.0025 inches of paint per Reference 13 and Reference 9, Appendix VI (for properties) and a 2 inches thick air gap (Reference 14) and a conservative low thickness of concrete of 24 inches is used from Reference 21.
Item 8 (side of drywell head -small cylinder) is modeled, with a 2.5 ft air gap outside the steel wall and conservatively low thickness of 1.5 ft of concrete (scaled from Reference 21). The thermal conductor has an adiabatic heat transfer boundary condition. Only heat conduction is assumed in the air and concrete layers. This is conservative.
Item 9 (top of drywell head) is modeled with a 6.7ft air gap outside the steel wall and a conservative low thickness of 24 inches of concrete (part of the concrete plugs) - (scaled from Reference 21).
RRUs (References II and 12)
ARRU   =  1272.8 ft2 . thickness =0.125 inches (used in OPL4A-Reference 12).

Entergy                   Calculation VYC-2405' Rev. 0              Page 20 of 85        -.
Vent Pipes Vent pipes surface area A vent pipe,= 2885.7 ft2 , thickness =0.125 inches (used in OPL-4A-Reference 12)
NOTE: The total Surface Area of the steel components adds up to the value obtained from OPL-4a (i.e., Table 6 Total = 15246.06 ft2 , RRUs = 1272.8 ft2 , Vent Pipes = 2885.7 ft2 , thus Total = 19405 ft2 )
Total Surface area of Concrete Exposed to Drywell Air Space The surface area of the pedestal is the only concrete component quantified in OPL-4A. .The drywell floor is ignored because it may be covered with liquid and not directly exposed to the drywell airspace. Only the outer surface area of the pedestal was considered in OPL- 4A as well as Reference 9 because the inner surface *has limited communication with. the drywell atmosphere. The biological shield wall (BSW) is a concrete structure surrounding the reactor pressure vessel and located above the reactor pedestal. Because of the proximity to the reactor pressure vessel the BSW is at a temperature greater than the drywell (DW) ambient and thus a heat source (already incorporated into- the drywell ..heater) and a heat sink only when its temperature drops 'below the DW temperature. Because of the uncertainty of the BSW temperature and its limited value as a heat sink, the BSW is not considered here.
The OPL-4A value for the area is used and = 2068 ft2 (A value of 2108 ft2 was used in the model, addressed in Case 5).
Thickness of Concrete Exposed to Drywell Air Space' From Reference 5 = 4ft.
Properties of Materials Table 7 - Thermo physical properties of Passive Heat Sink Materials (Reference 5)
Material              Density (Ibm/ft3 )  Specific Heat       Thermal                 References (Btuflbm-0 F)       Conductivity (Btulhr-ft 0 F)
Carbon Steel          489.0               0.11                32 F 31.8               15 68 OF 31.2 212 OF 30.0 392 OF 27.8 572 &deg;F     26.0 Concrete              145                 0.156               0.92                    16 Paint                 288                 0.2                 0.125                   9

Enteigy                    Calculation VYC-2405 Rev. 0               Page 21 of 85-4 Air Thermal Properties From Reference 15.
T (1F) K(Btu/hr-ft-0 F) Cp (Btu/lbm-0 F) p(lbm/fl 3 )
100       0.0157               0.24            0.07092 150       0.0167               0.241           0.06511 200       0.0181               0.241           0.06017 250       0.0192               0.242           0.05593 300       0.0203               0.243           0.05225 400       0.0225               0.245           0.04617 Heat Transfer Boundary Conditions On the inner surface of all the thermal conductors, the heat transfer coefficient is calculated by the GOTHIC code. The following options are used:
        -Direct'Heat transfer Option.
        -Summation of the condensation and convection heat transfer.
        -Max of Uchida and Guido-Koestel condensation heat transfer option (sensitivity with Uchida for the limiting cases).
        - Radiation heat transfer option is OFF for all heat structures with exception of the drywell dome.(sensitivity with option OFF for the limiting case).
        - The surface orientation is "FACE DOWN for the drywell dome", thermal conductor
           #13,and "VERT SURF' for heat conductors 5 through 12.
        -All thermal conductors use 'VAP" option.
The outer surface boundary condition is conservatively assumed to be adiabatic. The heat transfer coefficient is set to zero.
5.3 GOTHIC Drywell SBO Model Development The following changes to the input SBO-NoLeak-80 (Reference 1) to produce SBO-dry-weIl2 are being added:
    -   drywell heat load
    -   drywell heat slabs
    -   leakage from drywell to wetwell
    -   modifications to the vacuum breaker modeling The GOTHIC input is presented in Attachment A.
The GOTHIC model used for all cases is presented in Figure 2.

A Entergy             Calculation VYC-2405 Rev. 0         Page 22 of 85 SBO - 80P-Noleak-drywelll Mar/15/2005 14:09:56 NGOTHIC Version 7.0p2(QA) - April 2002
       *File: /home/scbor/vyc-2120ccn/SENSITIVITY/SBO/drywell-SBO/SBO-drywell2 I
Figure 2 - GOTHIC SBO Model

9 9 - ale o S Ad Entergy                     Calculation VYC-2405 Rev. 0                Page 23 of 85.
Flow Paths
.Flow path 21 is added to model the vacuum breakers leakage path from the drywell air space to.
the wetwell air space. Per Reference 5, the maximum leakage area is 0.12 ft2 . The elevations and the height of this junction were elected to be the same as the vacuum breaker junction since the leakage is around the vacuum breakers.
The K rever    =K forward = 1.5 (expansion & contraction) (Reference 29)
Thermal Conductors Twelve new thermal conductors were added. The description of the thermal conductors was given in Section 5.2.
The temperatures of 11 of the thermal conductors were set at 160 TF, the pedestal thermal conductor is set at 152 'F. On page 73 of Reference 9 the average temperature for the middle and the top drywell node is calculated as 151.94 'F. Hence, the thermal conductors are set, conservatively at i 60 'F. The pedestal is in the lower drywell and middle drywell hence 152 'F is used (average for middle and top drywell is conservative). In Case 5, the temperatures of the heat slabs which represent the drywell wall were set at 170 'F (very conservative assumption).
Functions Two new functions are added Function 17 (FF17) represents the power to the drywell from the structures exposed to steam.
The function multiplies Q initial and represents (Tliquid - Tdzyell).
The FF17 is:
AT (CV41)
 -500             0 0                0 380              1 380000           1000 The independent variable is the temperature difference between        (Tuquid - Tdrwell), CV4 1.
Function 18 is identical to the Function 17, but the independent variable is CV42 (T vapor       TdyweCll)
Tref = Initial Vessel temperature - Initial Drywell temperature = 550VF -170GF =380 OF where 550 'F is the initial vessel temperature, and 170 'F is the initial drywell temperature. (550 is determined from the GOTHIC model at time zero and 170 'F is the maximum drywell temperature, OPL4A-Reference 5).

Entergy                   Calculationr VYC-2405 Rev. 0               Page 24 of 85 Valves A new valve (V5) was added to represent the vacuum breaker. It opens on trip 33 (0.5 psi difference between wetwell and drywell (Reference 2) and it closes on trip 34 (0.3 psi-arbitrary, since the vacuum breaker valves reseat when the pressure difference becomes less than 0.5 psi. A quick close valve is used for this component since the valve will close as soon as the 0.5 psi difference between wetwell and drywell disappears.
The vacuum breaker valve is modeled as Valve Type 3, with an area of 17.6737 ft2 (Reference 9). Note: the area of the valve from Reference 9 is slightly larger than the area of the flow path in which it is located. The valve area will have minimal impact on this analysis because the flow is limited by the area of the flow path. The area of the valve was changed to the area of the flow path in the final case analyzed, Case 5.
Materials Four new materials are added. The properties for the new materials are described in Section 5.2.
Trips and Controls Trip 18 is modified to ADS when the vessel pressure difference between RPV and drywell is lower than 100 psi instead of 50 psi in the original model. This trip is not used, however, the SRV valves will open at a AP of 100 psi, not 50 psi.
Trip 21 is modified to start depressurization (cooldown) at one hour (3600 seconds) in order to limit the drywell temperatures.
Trips 33 and 34 are added to open the vacuum breakers valves at 0.5 psi pressure difference between wetwell and drywell (trip 33) and close it on a AP of 0.3 psi.
Coolers & Heaters Two new heaters are added, 5H and 6H to model the vessel heat to the drywell. These heaters are described in Section 5.1.
For heater 5H the heat rate of 274.83 is multiplied by the FF 17, while for heater 6H the heat rate of 1997.58 is multiplied by FF 18.

Entergy                  Calculation VYC-2405 Rev. 0                Page 25 of 85      -
Heat Transfer Coefficients Types Two heat transfer coefficients arc added.
The following options are used:
       -Direct Heat transfer Option.
       -Summation of the condensation and convection heat transfer.
       -Max of Uchida and Guido-Koestel condensation heat transfer option (sensitivity with Uchida for the limiting cases).
       - Radiation heat transfer option is OFF for heat transfer coefficient type 6.
       - Radiation heat transfer option is ON for heat transfer coefficient type 7.
The use of the radiation option has no effect on the results at these low temperatures
        - The surface orientation is "VERT SURF' for heat transfer coefficient type 6.
        - The surface orientation is "FACE DOWN for the drywell dome", heat transfer coefficient type 7.
The use of the surface orientation is appropriate since this is the thermal conductor physical arrangement.
    -  -The heat transfer coefficient types 6 and 7 use 'VAP" option since this is the drywell medium.
       -Convection bulk T model: Tg-Tf. The bulk temperature is the calculated vapor temperature. Tf is the maximum between the calculated wall temperature and the calculated saturation temperature.
       -  Condensation heat transfer Bulk T Model : Tb -T, used. Tb is the minimum between the calculated vapor temperature and the calculated saturation temperature.
Control Variables Two control variables are added, 41 and 42 they represent the AT between Tliq in RPV and Tv drywell and between Tvap in RPV and Tv drywell, respectively. See Section 5.1.2 for additional information on the operation of these Control Variables.

                                                                                     .       _:- 11 I Enter           gy       Calculation VYC-2405 Rev. 0              Page 26 of 85 6.0 Calculation / Analyses Five (5) cases are analyzed:
Case I is called SBO-dryivell2. It is the base deck, developed from Reference 1 and described in Section 5.3. Case 1 assumes no RPV leakage, depressurization (cooldown) with a rate of 80 TF/hour a 100% humidity and base deck inputs as described in Section 5.0.
Case 2 is called Case SBO-drywell2-80-sensy2-NoLeak. Case 2 is identical to Case 1 with the change in humidity, changes in the leakage area and minor changes in the heat transfer type 6 and 7. Case 2 assumes 20% humidity and minor changes in the heat transfer type 6 and 7. These changes are described in Section 6.2.1.
Case 3 is identical to case SBO-drywell2-80-sensy2-NoLeak but with leak. It is called SBO-drywvell2-Leak-80-sensy. These changes are described in Section 6.3.1.
Case 4 is called SBO-drywell2-Leak45-sensy and is identical to Case 3 but with a slower depressurization (cooldown) rate. It assumes a depressurization of 45 TF/hour with leak in order to show that with an early depressurization and slower cooldown rate the results are not changed and the drywell temperature is not impacted negatively by a slower cooldown. Consistent with Reference 1, the RHRSW temperature is changed to 75 'F.
Case 5 addresses changes found during documentation and as part of review. These changes are described in Section 6.4.1.
Case 5 is called SBO-drywell2-comments. The following changes are made in Case 5 to address changes found during documentation and review:
            . change the temperatures of the steel structures from 160 'F to 170 'F (very conservative assumption),
            . change the K reverse for the Junction 3 from to 3.93 from 3.964,
* set the V3 Valve with the same area as the junction, and
* change the area for the pedestal from 2108 ft2 to 2068 ft2, consistent with OPL-4A.
Case 5 changes are described in Section 6.5.1.
* Entergy                    Calculation VYC-2405 Rev. 0              Page 27 of 85 6.1.1 SBO-drywell2 Model Development SBO-dryvell2 represents the base model for this calculation. The modification to the input are presented in Section 5.0 and the GOTHIC input deck is presented in Attachment A.
6.1.2 Case SBO drywell2 Results Figure 3 through Figure 11 present the main parameters for the base case SBO-drywel]2. Figure 3 presents the drywell temperature. The drywell temperature increases to about 285 0F after one hour. The heatup is arrested due to depressurization. At about 4 hours into the transient the temperature in the drywell starts increasing due to lower heat removal into the passive heat sinks (walls). The maximum drywell temperature is 290 'F. The air gap acts as an insulation and the steel liner is almost at 245 'F. However, after 2 hours and 10 minutes the low pressure pumps are available so the operators can spray the drywell with the RHR pump, if needed. The results indicate that the temperatures in the drywell stay below the EQ limit and the'drywell liner is well below the 281 OF for the SBO coping duration.
 -Figure 4 presents the containment pressure. Due to the higher leak area the drywell and the wetwell are at the same pressure. At about 2 hours the pressure in the drywell is too low to spray the drywell, (unsafe area of DWSIL(EOP-3 -Primary Containment Control -Reference 30))
however the pressure increases' to about 6 psig at about 12000 seconds at which point the operators would able to spray the drywell with the RHR pump, if needed.
Figure 5 presents the RPV pressure. At one hour into the event it is assumed that the operators start depressurization (cooldown). The pressure drops to the HPCI shutoff pressure of 165 psia at about 12000 seconds. At that point only about 450000 lb were injected from CST. (Figure 12). At
-this point the RPV is depressurized and the CS is available to inject.
Figure 7 presents the RPV level. The core stays covered. There is a dip in the normal level at about 12000 seconds when HPCI stops injecting and CS pump has not yet injected. This is due to the fact that the CS pump was set to inject at 14000 seconds; however CS is ready to inject at 7800 seconds..
Figure 8, Figure 9, and Figure 10 presents the drywell liner temperature. The drywell liner stays below 260 F for the 7 hours analyzed. After 2 hours and 10 minutes the low pressure pumps are available for suppression pool cooling, drywell spray and maintaining vessel inventory.
Figure 11 presents the suppression pool temperature. Since the vessel is depressurized early, the suppression pool temperature is below the maximum of 182.2 'F calculated in Reference 1, hence no containment overpressure is required.

Entergy                                                                                Pa.g Calculation VYC-2405 Rev. 0                Page 28 of 85 -K SBO    -    SOP-Noleak-drywelll Mar/03/2005 13:28:01 GOTHIC Version 7.0p2(QAI - April 2002 File: /home/schor/vyc-2120ccn/SENSITIVITY/SBQ/drywell-SBO/SBO-drywell2 L                   Drpwell Temperature Ior]la         Dlye l 3 . 14.4    8    2.
n                           eas
                                 ..                    Time (Sac) goe       1-4,24#     M..J6112S 1tt46,49 Figure 3 Drywell Temperature Case SBO-drywell2 SBO - 80P-Noleak-drywelli Mar/03/2005 13:33:59 GOTHIC Version 7.0p2(QA) - April 2002 File: /home/schor/vyc-212occn/SENSITIVITY/SBO/drywell-SBO/SBO-drywell2 3-centulneza             Wrestflr!
            -i            .-                Illli
            -                Sc           1                            T p    ..                           1             T        l S (see)
               <nne ww~f"}       srISsItSIS 11,45,45 Figure 4 Containment Pressure - case SBO-drywell2

- Entergy                        Calculation VYC-2405 Rev. 0             -Page 29 of 85 -
SBO - 80PF-oleak-drywelll Mar/03/2005 13:29:48 GOTHIC Version 7.0p2t(A) - April 2002 File: /home/schor/vyc-2I20cn/SENSITIVITY/SBO/drywell-SBO/SBO-drywell2 24                  VT1 pressures
                             .to C-
:c                                .~
                          -c
: 1.                                   Tt    (c1o) .
sg  t   p~tJ^    1.,/S3fl/g3  IK.4s.4I Figure 5 - RPV Pressure - Case SBO -drywell2 SBO -. 80F-Noleak-drywelll Mar/03/2005 13:30:33 GOTHIC Version 7.Op2CQA) - April 2002
          .Pile: /home/schor/vyc-2120ccn/SENSITIVITY/SBO/drywell-SBO/SBO-drywell2 4    ,               Reactor Vessel Temperature Tv4 n.4 ST4 Th4 Z         Liqui Oatu                          r       To II iiSj       T i.n Tilke (;ec)
G*t     7 0*4211 ".I/6912M           4149M Figure 6- RPV Temperature - Case SBO-drywell2

':Entergy                        Calculation VYC-2405 Rev. 0                            Page 30 of 85 ; '
SBO - 8CP-Noleak-drywelll Mar/03/2005 13:01:07 GOTnIC Version 7.0p2(QA) - April 2002 File: /home/schor/vyc-212Occn/SENSITIVITY/SEO/drywell-SBO/SBO-drywell2 21               P" Liqid Lewel 114 34   _--              -       -
                    -s            %             _       _      I o    C          l        C.....1....
N    4           12l         .      20 2 8 32 , *3z T5a* (see) 4#=2t 1.021ti)s     N"/0/2185 21,46.44 Figure 7- RPV Level - Case SBO-drywell2 SBO - 8OP-Noleak-drywelll Mar/03/2005 13:29:11 GOTBIC Version 7.0p2(QA) - April 2002 File: /bene/schor/vyc-2l20ccn/SENSITIVTY/SBO/drywell-SBO/SBO-drywell2 48               s5uf att Tetzature TJS 73I T78 o _ _       _   _
I.
4WM    t 172tt*O Pv/&#xa2;1411S 11,4c,43 Figure 8 - Surface Temperature for Heat Slabs 5,6 & 7 - Case SBO-drwell2

                                                                                        . .  - i .. I -
Entergy                           Calculation VYC-2405 Rev. 0            Page 31 of 85 NM SBO - 80P-Noleak-drywelll Mar/03/2005 13:28:24 GOTHIC Version 7.0p2(QA) - April 2002 Pile: /home/scbor/vyc-2120ccn/SENSITIVITY/SBO/drywell-SBO/SBO-ddrywell2 SI                    Surface X~eurezt          ire G738733             i31           120 UU I..ttt...1t.1
                                 '        3.571   7.14)    07     1~--4.2
                             .                             Tiac (ceo)
: m. l 1.h24,a          gg..}e)2eS5 u1,4c.,,
Figure 9 Surface Temperature for Heat Slabs 8,9,10,11 - Case SBO-drwell2 SBO - 80F-Noleak-drywelll Mar/14/2005 14:45:07 GOTHIC Version 7.Op2(QA) - April 2002 File: /homelschor/vyc-2120ccn/SENSITIVITY/SBO/drywell-SBO/SBO-drywell2 surface tesperatmra T313 7312 z
tI 0.
TSae (Sao)
M    7.022(0)         H2st C/141205 11t2SS4S Figure 10 - Surface Temperature for Heat Slabs 12 and 13 - Case SBO-drwell2

-Entergy                         Calculation VYC-2405 Rev. 0                                 Page 32 of 85 S80 - 80P-Nolean-drywelll Mar/03/2005 13:31:38 GOTHIC Version 7.0p2(QA) - April 2002 Pile: /bome/scbor/vyc-2120ccn/SENSITIVITY/SBO/drywell-SBO/SBO-drywell2 2                 VttwtU Tnperature ST2 1i2 o= _          i b
Z
                .v-M i
                                                      .ia*     (sec)
              *t?    ?w.*P2tts)  rPrWS1266  1J.46,48 Figure 11 Suppression Pool Temperature case SBO-drywell2 SBO - 80W-Noleak-drywelll Mar/14/2005 14:02:38 GOTHIC Version 7.0p2(QA) - April 2002 Pile: /home/g6chor/vyc-2120ccn/SENSITIVITY/SBO/drywell-SBO/SBO-drywIll2 40              zutegrated )Wcl Flow c33 r    o _
                          .                  ..            ,                . I                           f tXI                                 I.2t}Wl/2s1.XX S.F     7.2       1t0.t     24;4   :18 21.6 25. 2 S    (Se).
Figure 12- Integrated HPCI Flow - SBO-drywell2

Ad Entergy                  Calculation VYC-2405 Rev. 0             Page 33 of 85       r A
6.2 Case SBO-drvwell2-80-sensV2-NoLeak 6.2.1 Model modification Table 8 presents the modifications to the base deck SBO-drywell2 to produce SBO-drywell2 sensy2-NoLeak.
The following modifications were made:
The Heat Transfer Coefficient Types 6 and 7 were modified to use Uchida correlation for.
condensation heat transfer instead of MAX (maximum of Uchida or Guido-Koestel). For this case since there is no.RPV leakage, the choice of condensation correlation should have a minimal impact on results.
For the Heat Transfer Coefficient Type 7 the radiation option was turned off. Again, at these small temperatures, the radiation has a minimal impact on results.
The humidity in Volume 1 (Drywell) was modified from 100% humidity to 20% humidity to encompass all the humidity range in the drywell (Reference 5).
The reverse loss coefficient for the vacuum breakers was changed from IeI8 to 3.964 (equal to the forward loss coefficient).
A coefficient of 3.93 should have been used. This is corrected in Case 5.
The vacuum breaker reverse coefficient is the weighted sum of the flow paths 7, 8 & 9 of Reference 9. (Same as the forward loss coefficient)
K reverse = 1.168 (15.63/16.23)2 + 2.528 (15.63/16.23)2 + 0.5 (1.53/1.53)2 = 3.93 (A K of 3.96 was used, less than a 1% difference)
The area of junction 21 is changed from 0.12 ft2 , maximum leakage to 0.0033 ft2 (allowable Tech Spec leakage) -Reference 33.
Table 8; Input Modifications -SBO-drywell2-80-sensy2-NoLeakvs. SBO-drywell2

Entergy                    Calcufation VYC-2405 Rev. 0                          Page 34 of 85 Modifications in /hame/schor/vyc-2120ccn/SENS]TIVITY/SBO/drywell-SBO/SBO-dryhvell2-80-sen Mar/10/2005 10:42:03 GOTHIC Version 7.0p2(QA) - April 2002 File: /home/schor/vyc-2120ccn/SENSITIVITY/SBO/drywell-SBO/SBO-drywell2-80-sensy2-NoLeak
     >                             Graphs (continued)
Graph                                      Curve Number Gp          Title        Mon     1        2          3        4         5 42    Heat to the sup          CQ4H     CQ2H 43    Leak Flow                FL4       FL19      FL20 44    Integrated Leak          cv40 45    Title                    cv39 46                             FV18      FL18      FOID 47                             cv38 48    Surface  Tempera         TB5       TB6       TB7 49    Surface  Tempera         TA8      TA9        TA10 50    Surface  Tempera         TA11      TA12      TA13 51    Surface  Tempera         TB8      TB9        TB10     TB11 52    Surface  tempera         TB13      TB12 53                           TP8t630 TP9t600 TPlOt60 TP11t60 54                           TP13t5S TP12t50 55    Drywell Tempera          lVl       TU -
Heat Transfer Coefficient Types - Table I Heat.                 Cnd       .Sp        Nat         For Type Transfer      Nomtnal     Cnv Cnd Cnv          Cnv         Cnv      Rad
         # Option         Value FF Opt Opt HTC            Opt         Opt      Opt 1 Correlat              0                   VERT SURF  PIPE  FLOW  OFF 2 Correlat              0                   VERT SURF  PIPE  FLOW   OFF 3 Correlat                                  FACE DOWN  PIPE  FLOW  OFF 4 Correlat                                . FACE UP   PIPE   FLOW  OFF 5 Sp Heat          0.
6 Direct                  ADD UCHI          VERT SURF PIPE FLOW OFF 7 Direct                  ADD UCHI          FACE DOWN PIPE FLOW OFF Run Control Parameters (Seconds)
Time     DT      DT      DT       End   Print Graph        M'x. Dump Phs Chng Int     Min     Max    Ratio Time       Int       Int     CPU      Int Time Scale I    le-06       1.      1. 100.         5.      0.1   le+06       0. DEFAULT 2    le-06       1.      1. .1200.      50.       1. 1e+06       0. DEFAULT 3    le-06       1.      1. 1300. 500.       10. le+06       0. DEFAULT 4     le-06       1.      1..25000. 600.       10. le+06     -0.      DEFAULT

7~Entergy                   Calculation VYC-2405'ReV. 0                     Page 35 of 85   '
Modifications in /hane/schor/vyc-2120ccn/SENSITIVITY/SBO/drywell1-SBO/SBO-drywel112-80 -sen Mar/1012005 10:42:03 GOTHIC Version 7.0p2(QA) - April 2002 File: /hcrme/schor/vyc-2120ccn/SENSITIVITY/SBO/drywell -SBO/SBO-drywel112-80-sensy2-NoLeak N                 Volune Initial Conditions Liquid Relative Li-quid Ice         Ice Vol   Pressure Vapo        T7~. Humidity Volume Volume Surf .A.
             # (psia)        R                   (M     Fractio Fract. (ft2) def         14.7      80. 80.          60.       0. 0. 0.
1          16.4     170. 170.          20.       0. 0. 0.
2          14.7      90. 90.        100. 0.39497      0. 0.
3          16.4      90.-    90.        100. 0.00595     0. 0.
4       1045.2 549.97 533.12            100. 0.60794      0. 0.
                                         .Graphs Gr~ph               TitleCurve                    Number     4      5 1    DrywI 1 Teapera           IVi       Thi 2     Wewel1 Tempera           1V2        112 3    Contalrinent Pre         PRI        PR2 4    Reactor Vessel           TV4        11-4     ST4     TD4 5.-  PRiR Heat Exchan 6    Reactor Vessel 7    Torus Water Vol          AL2 8    Heat Exchanger                      t2lH 9    Wetwell. Vessel           1L2       TL4      TLI Conductor Tenpe          NAl        TA2      TA3     TA4 11   Integral Vessel 12   Vapor & Conduct:
13   Liquid & Conduc          TL4        TA2 14   Vapor & Conduct.         TV2        TA3 15   Liquid & Conduc          TL2        TA4 16   Vapor Heat Tran          HAI.
17   Liquid Heat Tra          hA2 18   Vapor Heat Tran          HA3 19   Liquid Heat Tra          hA4 20   Feedwvater & Bre         FL9        FL4 21   RPV Liq~uid Leve         LL4 22   SRV and ADS Flo          FVIO       FV11 23   Feedwater Entha          cv29 24   RPV Pressures            PR4 25   Feedwater Conrtr         cv27       cv28 26   Inte~rated Feed          cv4 27                            FL5 28   Vessel Droplet.          AD4 29   ECCS Injection           FL7        FL12     FL14 30   RPV Pressure             PR4'       VC2V     FL7 31   ADS Valve Posit          VC3V 32   SRV Position             vc2V 33   Cool dow~n FLo.w         FV16 34   Vessel Drop Dia          D14 35   Reactor Vessel           PR4 36   Suppression poo           112 37   Reactor Vessel           PR4 38    Supression Poo           112 39      HPIFlow Rate          FL18 40   Integrated HPCI          cv39 41   Core Spray Flow          FL8

AdEntergy                 Calculation VYC-2405 Rev. 0                         Page 36 of 85' Modifications -in/home/schor/vyc-2120ccn/SENSITIVITY/SBO/drywell-SBO/SBO-drywel12-80-sen Mar/10/2005 10:42:03 GOTHIC Version 7.0p2(QA) - April 2002 File: /home/schor/vyc-2120ccn/SENSITIVITY/SBO/drywell-SBOISBO-drywel12-80-sensy2-NoLeak Flow Paths - Table 2 FlIow   Flow     iyd.      Inertia Friction RelatlAve Dep Man Strat Path    Area     Diam. Length Length Rouc ;h- Bend Trn Flow
             # (ft2)-   (ft.)       (ft)       (ft.)       nes55    (deg) opt Opt 1    283.529      6.75      89.13           0.              . O.  -  NONE 2 286.114       1.948        4.16'          0.               -1.  -  NONE 3      15.63. 1.5625      44.925       28.72                  0.  -  NONE 4 0.001005 0.03568            0.1                             0.  -  NONE 5       3.14         2.      0.01                             0.  -  NONE 6       3.14         2.      0.01                             0.  -  NONE 7       3.14         2.      0.01                             0.  -  NONE 8       3.14         2.      0.01                             0.  -  NONE 9       3.14         2.      0.01                             0.  -  NONE 10    0.09945 0.35584         0.01                             0.  -  NONE 11     0.3978 0.3554          0.01                             0.  -  NONE' 12        3.14        2.      0.01                             0.  -  NONE 13       3.14         2.      0.01                             0.  -  NONE 14       3.14         2.      0.01                             0.  -  NONE 15       3.14
* 2.         0.01                             0.  -  NONE 16       3.14         2.      0.01                             0.  -  NONE 17       3.14         2.      0.01                             0.  -  NONE
       -18       3.14         2.      0.01                             0.  -  NONE 19     0.5454 0.8333           0.1         0.1                     -  NONE 20     0.5454   0.5454         0.1         0.1                     -  NONE 21     0.0033         1.
* 1.           0.                    -  NONE Flow Paths - Table 3 Flow    Fwd. Rev.             Critical Exit. Drop Path    Loss     Loss     Caow. Flow        Loss Breakup Coeff. Coeff. Opt. Model Coeff. Model I     4.2243   4.2243     ON     TABLES          1. OFF 2          1. 0.78    ON     TABLES          1. OFF 3      3.964    3.964     ON       OFF           0. OFF 4          0.            OFF     TABLES          1. OFF 5                        OFF       OFF           0. OFF 6                        OFF       OFF           0. OFF 7                        OFF       OFF           0. OFF 8                        OFF       OFF          .0. 0FF 9                        OFF       OFF           0. OFF 10                        OFF     TABLES          0. OFF                             )
11                        OFF     TABLES          0. OFF 12                        OFF       OFF           0. OFF 13                        OFF       OFF           0. OFF 14                        OFF       OFF           0. OFF 15                        OFF       OFF           0. OFF 16               le+18 OFF          OFF           0. OFF 17                        OFF       OFF           0. OFF 18                        OFF       OFF           0. OFF 19                        OFF       OFF           0. OFF 20                        OFF       OFF.          0. OFF 21         1.5      1.5 OFF         OFF           0. OFF

:Entery                Calculation VYC-2405 Rev. 0              Page 37 of 85 6.2.2 Case SBO drawell2-80-sensy2-NoLeak Resuflts Figure 13 through Figure 20 present the main parameters for the case SBO-drywvell2-80-sensy2-NoLeak. Figure 13 presents the drywell temperature. The maximum drywell temperature is about 289.4 TF and is reached after one hour and 30 minutes.. The heatup is arrested due to depressurization. At about 4 hours into the transient the temperature in the drywell starts increasing due to lower heat removal into the passive heat sinks (walls) . The air gap acts as an insulation and the steel liner is almost at 255 'F. However, after 2 hours and 10 minutes the low pressure pumps are available so the operators can spray the drywell with the RHR pump, if needed. The results indicate that the temperatures in the drywell stay below the EQ limit and the drywell liner is well below the 281 OF for the SBO coping duration.
Figure 14 presents the containment pressure. Due to a lower leak area the drywell and the wetwell are at not at the same pressure, the vacuum breaker opens to relieve the pressure difference at about 14000 seconds. At about 2 hours the pressure in the drywell is too low to spray the drywell, (unsafe area of DWSIL (EOP-3 -Primary Containment Control .-Reference 30)) however the pressure increase to about 6 psig at about 10800 seconds at which point the operators would be able to spray the drywell with the RHR pump, if needed.
Figure 15 presents the RPV pressure. At one hour into the event it is assumed that the operators start depressurization. The pressure drops to the HPCI shutoff pressure of 165 psia at about 12000 seconds. At this point only about 450000 lb were injected from CST (Figure 17). The RPV is depressurized, and the CS pump is available to inject.
Figure 16 pump presents the RPV level. The core stays covered. There is a dip in the normal level at about 12000 seconds when HPCI stops injecting and CS does not inject yet. This is due to the fact that the CS pump was set to inject at 14000 seconds; however CS is ready to inject at 7800 seconds, provided the pressure permissive is reached.
Figure 18, Figure 19, and Figure 20 presents the drywell liner temperature. The drywell liner stays below 260 'F for the 7 hours analyied. After 2 hours and 10 minutes the low pressure pumps are available for suppression pool cooling, drywell spray and maintaining vessel inventory.
The suppression pool temperature for this case is very similar to the case SBO-drywell2 since the input changes results in minor changes to the drywell temperature and pressure but not in the suppression pool temperature since the heat transferred to the drywell is not subtracted from the vessel energy.

- Entergy                       Calculation VYC-2405 Rev: 0                Page 38 of 85; -
SB0 - 80F-Noleak-drywell2-sensitivities-set2 Mar/10/2005 10:36:15.
GOTHIC Version 7.0p2(QA) - April 2002 File: /home/schor/vyc-2120ccn/SZNSITIvITY/SBo/drywell-sBo/sBo-drywel12-80 5Drwell Temptraftre Tvi TlI
                      *0'
                    -                             Dime (see) 4ms i 0Li Figure 13 - Drywell Temperature -Case SBO-drwell2-80-sensy2-NoLeak EBO - 80F-Noleak-drywell2-sensitivities-set2                     :
Mar/10/2005 10:56:11 GOH1IC Version 7.0p2(QA) - April 2002 Pile: '/ome/schor/vyc-212 0ccn/SENSITlVITY/SBO/drywell-SDO/SBO-drywelI2-80 3         ContA  maut resmre
                                                       -Pi
                                                       -l
                                                           .Dnryll relsurt f.t             . I..f2 Time (Ste)
Figure 14 -Containment Pressure - Case SBO-drwell2-80-sensy2-NoLeak

=Entergy                 CalculationVYC-2405 Rev.O                    Page 39 of 85 - I NVyy SBO - 80P-Noleak-drywell2-sensitivities-setZ Mar/10/2005 11:11:35 GOTHIC Version 7.Op2(QA) - April 2002 File: /home/schor/vyc-212Occn/SENSITIVITY/SBO/drywell-SDO/SBO-drywell2-80 24            "V Frsusmrts
:-oI -'nIN'1 1L             -1 I       FVat (Ste)'
F-Figure 15- RPV Pre ssure - Case SB0-drwell2-80-sensy2-NoLeak SBO - BOF-Noleak-drywell2-sensitivities-set2 Mar/10/2005 11:15:46 GOTHIC Version 7.0p2(QA) - April 2002 Pile: /home/schor/vyc-2120ccn/SENfSTIVrmY/SDO/drywell-SBO/SBO-drywell2-80 I          sRPV Liquid Levl LL4
                 *7 ft   "-  /t/J       .   -   -    -   -    - .  -
Tih   (Sao)
Figure 16 - RPV Level - Case SBO-drwell2-80-sensy2-NoLeak

                                                                                                             . . . I . - ; Ws
                                                                                                                            -1. -- z
.. l7ntergy                                                                                          '.y
                                                                                                      .1 .
Page 40 6f185        h -
Calculation VYC-2405 Rev. 0 SBO - SOF-Noleak-drywell2-sensitivities-set2 Mar/10/2005 11:10:47 WCMHIC Version 7.0p2(QA) .- April 2002 File: /home/sdcor/vyc-212 eccn/STNSITIVITY/SBO/drywell-SBO/SBO-drywell2-80 0           Xtegratd )WC1 71w.
j L.
Us             1                       1'.1
                                       .6 7.          .8O   4.
* 2A.nt T3mc (sec) 4*Th  ?.$ZCt9A o.zII*IU*IS *I,@.452 Figure 17 Integrated HPCI Flow - Case SBO-drwell2-80-sensy2-NoLeak SBO - 80P-Noleak-drywell2-sensitivities-set2 Var/1D/2005 10:33:12 GOrBIC Version:7.0p2(QA) - April 2002 File: /home/schor/vyc-2120ccn/S              SITSVITY/SBO/drywell-SBO/SBO-drywe112-80 Surface Temperature TDS TD6 737 TJc (Ste) 6Tir 7.02CsA) 1a1.0pSI8S *141,-52 Figure 18- Surface Temperature for Heat Slabs 5,6 & 7 - Case SBO-drwell2                                                    sensy2-NoLeak

I . -  Entergy-                       Calculation VYC-2405 Rev. 0                        Page 41 of'85   ''
SBO - B0F-Moleak-drywell2-sensitivities-set2 Mar/10/2005 10:34:30 GOTHIC Version 7.Op2tQA) - April 2002 File: /hone/echor/vyc-2120ccn/SEHNSITrVITY/SB0/drywell-SBO/SBO-drywell2-B0 Surface Teartrature DZ8 529 TR30         11 a.
R.
I-SIAR (see)
                       -e74ts        -1t~e291266%      t 5 Figure 19 - Surface Temperature for Heat Slabs 8, 9, 10 & 11 - Case SBO-drwell2-80-sensy2-NoLeak' SBO - 80F-Moleak-drywell2-sensitivities-set2 Mar/10/2005 10:32:24 GOaTIC Version 7.0p2(QA) -PApril 2002 Pile: Jhome/schor/vyc-2120ccn/SKNSITxVIr/SBO/drywel1             -SBO/SBO-drywe112-80 52             surac       ta 7trature D313 D312 o _
o                               NMat 31)112 a.s At Tlse (stc)
                    $*t7000"llt                       oa Figure 20 Surface Temperature for Heat Slabs 12 & 13 - Case SBO-drwell2                                                               sensy2-NoLeak

Entergy         '   - Calculation VYC-2405 Rev. 0                 Page 42 of 85        :-
6.3 Case SBO-drywell2-Leak-80-sensv 6.3.1 Model modification Table 9 presents the modifications to the deck SBO-drywell2-80-sensy2-NoLeak to produce SBO-drywell2-Leak-80-sensy The following modifications were made:
On BC 13, the ON trip is set to zero (0). This allows for a constant leak of 8.4585 lb/sec to leave the vessel.
The end time was changes to 14400 seconds (4 hours) since the purpose of this calculation was to show that the drywell temperature stays below the EQ drywell temperature and the drywell shell stays below 281 TF for the duration of 2 hour and 10 minutes.
Table 9 SBO-drywell2-Leak-80-sensy vs SBO-dryvell2-80-sensy2-NoLeak

                                                                                    ,~~~~~~~   ~_t .. ' _ V'' _ _
.. . .. Entergy                   Calculation VYC-2405 Rev. 0                                  Page 43 of '85 PNa 4
Modifications in /home/schor/vyc-2120ccn/SENSITIVITY/SBO/drywell-SBO/SBO-drywel12-Leak-8 Mar/14/2005 15:38:40 GOTHIC Version 7.0p2(QA) - April 2002 File: /home/schor/vyc-2120ccn/SENSITIVITY/SBO/drywell-SBO/SBO-drywell2-Leak-80-sensy Fluid Boundary Conditions - Table 1 Press.         Tem.             Flow          ON      OFF BC#    Description      (psia) FF       (F) FF (lbm/s) FF Trip Trip IF   RHR/LPCI Suctio        20.            160       v-0.002 8           1      13 2C   RHR/LPCI Discha        20.            160 3F   LPCS Suction           20.            160       v-0.002 7         30       31 4C   LPCS Discharge         20.            160 5F   Feedwater            1000.              el 5         1000 9         1       5 6F   RHRlTorus Suctl        20.            160       v-0.002 6         21        1 7C   RHR/Torus Disch        20.            160                           0 8F   HPCI/RCIC Sucti        20,   .        160        -326.1 9           1       0 9C   HPCI/RCIC Disch        20.            160 1OP  Cooldown Inlet          1. 10             1 11                    21        0 11F  Cooldown Outlet         1. 10             1 11          1 12 12F  CST Tank              14.7            135             587 13      28       27 13F  Vessel Leak          1050.          554.        -8.4585 13          0 14C  Vessel leak to       1050. 0          554 Run Control Parameters (Seconds)
Time    DT      LT      DT       End     Print Graph          Max      Dump Phs Chng Int    Min     Max    Ratio Time         Int        Int       CPU      Int Time Scale I    le-06       1.      1. 100.            5.      0.1    le+06          0.      DEFAULT 2    le-06       1.      1. 1200.         50.        1. le+06          0.      DEFAULT 3    le-06      .1.      1. 1300.       500.        10. le+06          0.      DEFAULT 4    le-06       1. 1. 14400.       600.        10. le+06          0.      DEFAULT Graphs Graph                                       Curve Number I         Title        Mon      1           2          3          4            5 1    Drwell Tempera            lVi       TL1 2    Wetwell Tempera           TV2       TL2 3    Containment Pre           PR1        PR2 4    Reactor Vessel            TV4        TL4        ST4         TD4 5    RHR Heat Exchan           xqlH 6    Reactor Vessel            AL4 7    Torus Water Vol           AL2 8    Heat Exchanger            tllH      t21H 9    Wetwell. Vessel           TL2       114         ILl 10   Conductor Tempe           TA1        TA2        TA3         TAM 11   Integral Vessel                        V4 12   Vapor & Conduct                      . A1 13   Liquid & Conduc           TL4       TA2 14   Vapor &Conduct            TV2       TA3 15   Liquid & Conduc           TL2       TA4 16   Vapor Heat Tran 17   Liquid Heat Tra           hA2 18   Vapor Heat Tran           HA3 19   L~quid Heat Tra           hFM 20   Feedsater & Bre           FL9       FL4

lEntergy                   Calculation VYC-2405 Rev. 0                Page 44 of 85 "
Modifications in /home/schor/vyc-2l2Occn/SENSITJVITY/SBO/drywell-SBO/SBO-drywel12-Leak-8 Mar/14/2005 15:38:40 GOTHIC Version 7.0p2(QA) - April 2002 File: /hcme/schor/vyc-2120ccn/SENSITIVITY/SBO/drywell-SBO/SBO-drywel 112-Leak-80-sensy Graphs (continued)
Graph                                   Curve Number Title       Mon      1       2      3       4     5 21   RPV Li uud Leve           LL4 22   SRV and ADS Flo           FV10   FV11 23   Feedwate erEntha          cv29 24   RPV Pres sures            PR4 25   Feedwate erContr          cv27   cv28.
26   Integrat ed Feed          cv4 27 RHR F10                     FL5 28   Vessel DDroplet           AD4 29   ECCS In3 3efflon          FL7    fL12    FL14 30   RPV Pres sure             PR4    VC2V    FL7 31   ADS Valv vePosit      - VC3V 32   SRV Posi tion             VC2V 33   Cooldowni. FLow           FV16 34   Vessel DDrop Dia          D14 35   Reactor Vessel            PR4 36   Suppress sion poo         TL2   -
37   Reactor Vessel            PR4 38   Su press sion Poo         LL2 39   HPC] Flc w Rate           FL1B 40   Integrat led HPCI         cv39 41   Core Spr ra Flow          FL8 42   Heat to tne sup          .CQ4H   CQ2H 43   Leak Flc w              .FL4     FL19    FL20 44   Integrat ,ed Leak         cv4O 45   Title                     Cv39 46                             FV18   FL18    FD18 47                             cv38 48   Surface  Tempera          TB5    TB6     1837 49   Surface  Tempera          TAB    TB6     TAI1 50   Surface  Tempera          TAll   TA12    TA13 51   Surface  Temra            TB8    TB9     TB10 I      TB11 52   Surface  tempera          TB13   TB12 53                          TP~t600 TP9t600 TPIOt6O TPI1t6O 54                          TP13t58 TP12t50 xXooos xwOL              xxx xooooXc voooo= sooooax xouooxx XXXoox

tEntergy                     Calculation VYC-2405 Rev. 0             Page 45 of 85 6.3.2 Results Case SBO-drvwell2-Leak-80-sensy Figure 21 through Figure 29 present the main parameters for the case SBO-drywell2-Leak sensy. Figuire 21 presents the drywell temperature. The maximum drywell temperature is about 290 TF and is reached at about 12240 seconds (3.4 hours). The drywell heatup rate is arrested due to depressurization, however the leak brings enough energy from the vessel to continue the heatup. At 7800 seconds the drywell temperature is 285.8 OF, well below the EQ limit of 325 OF.
Figure 22 presents the containment pressure. The available water to spray the drywell (Reference
: 30) is the Diesel fire pump per Appendix M of OE 3107 (Reference 35) and it takes about one hour for aligning the fire pump for drywell spray. The drywell pressure is high enough to allow for drywell spray, if needed. The drywell temperature does not exceed the EQ drywell temperature limit and the drywell shell temperature stays below the limit of 281 TF hence the analysis shows that drywell spray is not needed for the coping duration. At about 3 hours and 30 minutes the wetwell pressure reaches equilibrium with drywell and slightly exceeds the drywell pressure. The vacuum breakers do not open during the time of interest.
At about 4 hours the wetwell pressure is about 26 psig, close to the PSP limit of 27 psig.
However at this time the RHR pump is available for containment spray.
Figure 23 presents the RPV pressure. At one hour into the event it is assumed that the operators start depressurization. The pressure drops to the HPCI shutoff pressure of 165 psia at about 12000 seconds. At that point only about 540000 lb were injected from CST (Figure 25). At this time the RPV is depressurized, and the CS pump is available to inject.
Figure 24 pump presents the RPV level. The core stays covered. There is a dip in the normal level at about 12000 seconds when HPCI stops injecting and CS does not inject yet. This is due to the fact that the CS pump was set to inject at 14000 seconds; however, CS is ready to inject at 7800 seconds, provided the pressure permissive is reached.
Figure 27, Figure 28, and Figure 29 presents the drywell liner temperature. The drywell liner stays below 280 TF for the 4 hours analyzed. After 2 hours and 10 minutes the low pressure pumps are available for suppression pool cooling, drywell spray and maintaining vessel inventory.
The suppression pool temperature is not a parameter of importance for this calculation. In Reference I it was shown that the suppression pool temperature is lower for the cases RPV with leakage and lower for earlier depressurization hence the maximum suppression pool temperature will be lower that 182.2 TF, calculated in Reference 1.

-- Entergy                     Calculation VYC-2405 Rev. 0                  Page 46 of 85 SDO - 8OF-Noleak-drywell2-Leak-80-sonsitivities Mar/14/2005 11:02:51 GOTHIC Version 7.0p2(QA) - April 2002 Pile: /home/schor/vyc-2l20ccn/SHNSITIVITY/SBO/drzywell-SBO/SBO-drywel12-Le 1r*rell Teaperature Il        blvi         td    mpral     xre J--
            .0=    1. 2( a)              MU                    .~M TUC (see)
Figure 21 -Drywell Temperatures - SBO-drywell2-Leak-80-sensy SBO - BOP-Noleak-drywell2-Leak-80-sensitivities Mar/14/2005 11:03:43 GOTIC version 7.Op2(QA) - April 2002 File: /bome/schor/1yc-2120ccf/SENSITIVITY/SBO/drywell-SBO/SBO-drywel12-I.e 3Cetaiment pressu" It7.2  (0.8
                              .                     T~t~e (See drme      s2}"!!~@5         s8$
Figure 22- Containment Pressure - SBO-drywell2-Leak-80-sensy
* v - t ;# .
.. Entergy                     Calculation VYC-2405 Rev. 0                 Page 47 of 85 Pg4o8 SBO - 8OF-.Ioleak-drywcll2-Leak-80-Bensitivities Mar/14/2005 12:05:31 GOTHIC Version 7.0p2(QA) - April 2002 File: /bone/rc!cor/vyc-2120ccn/SENSITIVITY/SBO/drywell-SDO/SBO-drywell 2-Le 24           pv ?ressurts 5R4 Zo                               I rime (see)
VW$ 1.67201)s "-103/2015  16,1.
tI 1 Figure 23- RPV Pressure - SBO-drywell2-Leak-80-sensy SBO - 8OP-Noleak-drywell2-Leak-80-sensitivities Mar/14/2005 11:04:51 GOTHIC Version 7.0p2(QA) - April 2002 File: /home/schor/vyc-2120ccn/SmNSITIVIrY/SBO/drywell-SBO/SBO-drywell2-Le 21            ra   Liquid Level LL4 o-   _
T     '   )
tia   (sic)
Figure 24 RPV Level - SBO-drywell2-Leak-80-sensy

-Entergy Calculation VYC-2405 Rev. 0                        Page 48 of 85
                                                                                                     \'Jy.
SBO - BOP-Noleak-drywell2-Leak-80Csensitivities Mar/14/2005 11:09:57 GOTHIC Version 7.0p2(QA) - April 2002 File: /home/schor/vyc-2l20ccn/SENSITIVITY/SBO/drywell-SBO/SBO-drywell2-Le 40                 Xategrated 210        7ic.
S                                                      j-LW1 Tim   (9SC) t42/2 somg             -111s2l285  IS~A3.2 Figure 25 - Integrated HPCI Flow - SBO-drywell2-Leak-80-sensy SBO - 80P-Noleak-drywell2-Leak-80-sensitivities Mar/14/2005 11:10:57
               'GOTHIC Version 7.0p2(QA) - April 2002 Pile: /home/schor/vyc-2120ccn/SENSITIVITY/SBO/drywell-S5O/SBO-dryweel22-Le 13                Leak ?low
                                       .4 1S       L20 1A o~      ~ ~.i24;
                                                        ~ TUm    ~ S.;c)3.
(
(see) 11.2 12.6 r 40       7 42.112X 1      }2@SAtt Figure 26-Leak Flow - SBO-drywell2-Leak-80-sensy

- Entergy                      Calculation VYC-2405 Rev: 0               Page 49 of 85' SBO - SOF-Noleak-dzywell2-Leak-80-sensitivities Mar/14/2005 11:07:22 GOTHIC Version 7.0p2(QA) - April 2002 File: /home/schor/vyc-2120ccn/SENSITIVITY/SBO/drywell-SBO/SDO-drywell2-Le 48               uface Temptratre                  -
XIS 136        137 o
                                         .,.6       at
_1222 Time (sta) s~~~~~-      .,-,,_,--.--,. -
Figure 27 - Surface Temperature, Heat Structures 5,6,7 - SBO-drywell2-Leak                                                                  sensy SBO - 80P-Noleak-drywell2-Leak-80-sensitivities Mar/14/2005 11:06:34 GOTHIC Version 7.0p2(QA) - April 2002 Pile: /home/schor/vyc-212Occn/SKNSISIVITY/SBO/drywell-SBO/SBO-drywell2-Le 51             Smrfaee reperat~x Si 2        a sime (see) 4*3   7h31*k1 *^)I)S1212            11.21.1 Figure 28 - Surface Temperature, Heat Structures 8, 9, 10, 11 - SBO-drywell2-Leak-80-sensy

LII)qery    -.
Calculation VYC-2405 Rev. 0                    Page 50 of 85       W-SBO - 80P-Noleak-drywell2-Leak-80-sensitivities Mar/14/2005 .11:09:22 GOThIC Version 7.0p2(QA) - April 2002 Pile: /hboe/schor/vyc-2120ccn/SENSITIVITY/SBO/drywell -SBO/SBO- drywell2 -Le 7...
52         surface te.eratnre
                                .all Tin
     .              I.
Figure 29 - Surface Temperature, Heat Structures 12,13 - SBO-drywell2-Leak                                                  sensy

-Entergy                   Calculation VYC-2405 Rev. 0             Page 51 6f 85 6.4 Case SBO-drywell2-Leak-45-sensv 6.4.1 Model modification Table 10 presents the modifications to the deck SBO-dryivell2-Leak-80-sensy to produce SBO-drysvell2-Leak-45-sensy.
Two modifications are made, the depressurization (cooldown) table, is changed from 80 0F/hour to 45 0F/hr (same cooldown curve as in Reference I- Function 10).
The RHRSW temperature is changed from 85 TF to 75 &Fconsistent with Assumption 13 and Reference 1.
Table 10 SBO-drywell2-Leak-45-sensy vs SBO-drywell2-Leak-80-sensy

- Entergy                  CalculationVYC-2405 Rev. 0                  Page 52 of'85      N     :.
Modi ficati ons in /home/schor/vyc-2120ccn/SENSITIVITYiSBO/drye1ell-SBOISBO-drywel12-Leak-4 Mar/14/2005 17:49:07 GOTHIC Version 7.0p2(QA) - April 2002 File: /hone/schor/vyc-2120ccn/SENSITIVITY/SBO/drywell-SBO/SBO-drybell2-Leak-45-sensy Functions FF#    Description     Ind. Var. Dep. Var. Points.
0 Constant                                -        0 I RHR H Tube         Reynolds N Nusselt Nu        34 2 RHR HX Shell       Reynolds N Nusselt Nu        34 3 Decay Heat         Time (sec) Decay Heat        50 4 Punp Heat          Time (sec) Heat Rate          6
         '5 Feed Enthalpy          cv4       Dep. Var. 34 6 RHR/Torus          Time (sec) Flow (gpm)         3 7 LPCS Flow Curve        cv26      Flow (gpm)     13 8 LPCI Flow Curve        cv26      Flow (gpm)     12 9 Feed Flow              cv28       Dep. ar.       3 10 Cooldown Pressu    Time (sec) Pressure (        39 11 Cooldown Temper        cv33       Dep. Var.      3 12 'Cooldown Flow         cv32       Dep. Var.      3 13 Constant             Ind. Var. Dep. Var.         6 14 'ECCS Pump Heat Time (sec) Heat Rate             6
: 15. Check Valve         Ind. Var. Dep. Var.         6 16 SW.                Time (sec) Service Wa         6 17 Drywell Power C cv41              Dep. Var.      4 18 Drywell Power C cv42              Dep. Var.      4 Heat Exchangers - Table 2 Heat Scndy Scnd Scndy Scnd             Ext;    Ext. Ext. Ext.
Ex. Flow Flow        Terrp Temp      Flow Flow      Heat Heat
         # (lbm/s) FF        (F)     FF (llm/s)       FF (Btu/s) FF
       *1H        1. 16       75.
Graphs Graph                                      Curve Number 4         Title       Mon      .1         2       3      4   5 1   Drywell Tempera           TV1       TLI 2   Wetwell Tenpera           TV2       TL2 3   Containment Pre           PRI       PR2 4   Reactor Vessel            TV4       TL4      ST4    TD4 5   RHR Heat Exchan           xqlH 6   Reactor Vessel            AL4 7   Torus Water Vol           AL2 8   Heat Exchanger            tllH      t21H 9   Wetwell, Vessel           1T2       TL4      TL1 10  Conductor Tempe           TA1       TA2      TA3    TM 11   Integral Vessel          QL4       QV4 12  Vapor & Conduct           TV4       TA1 13  Liquid & Conduc           TL4       TA2 14  Vapor & Conduct           TV2       TA3 15  Liquid & Conduc           TL2       TM 16  Vapor Heat Tran           HAI 17  Liquid Heat Tra           hA2 18  Vapor Heat Tran           HA3

- Entergy                Calculation VYC-2405 Rev. 0                   Page 53 of 85          :
Modifications in /hcme/schor/vyc-2120ccn/SENSITIVITY/SBO/dry)well-SBO/SBO-drywel12-Leak-4 Mar/14/2005 17:49:07 GOTHIC Version 7.0p2(QA) - April 2002 File: /home/schor/vyc-2120ccn/SENSITIVITY/SBO/dryw l -SBO/SBO-drywel12-Leak-45-sensy
      *'.                       Graphs (continued)
Graph                                  Curve Number Title        Mon    1        2       3      4       5 19  Liqutd Heat Tra         hA4 20  Feedwater & Bre         FL9      FL4 21  RPV Liquid Leve         LL4 22  SRV and ADS Flo         FV10     FV11 23  Feedwater Entha         cv29 24  RPV Pressures           PR4 25  Feedwater Cortr         cv27    cv28 26  Integrated Feed         cv4 27  RHR Flow                FL5 28  Vessel Droplet          AD4 29  ECCS Injection          FL7      FL12    FL14 30  RPV Pressure            PR4      VC2V    FL7 31  ADS Valve Posit         VC3V 32  SRV-Position            VC2V 33  Cooldown FLow           FV16 34  Vessel Drop Dia         014 35  Reactor Vessel          PR4 36  Suppression poo         TL2 37  Reactor Vessel          PR4 38  Suppression Poo         LL2 39  HPCI Flow Rate          FL18 40  Integrated HPCI         cv39 41  Core Spray Flow         FL8 42  Heat to the sup         CQ4H     CQ2H.
43  Leak Flow               FL4      FL19    FL20 44  Integrated Leak         cv4O 45  Title                   cv39 46                          FV18     FL18    FD18 47                          cv38 48  Surface Tempera         TB5     TB6      TB7 49  Surface TeTpera         TAB     TA9 .TA10 50  Surface.Tempera         TA11    TA12     TA13 51  Surface Tezpera         TB8     TB9      TB10   TB11 52  Surface tenpera         TB13    TB12       . -
53                        TP8t600 7P9t600 TP10t60 TPllt60 54                        TP13t5O 7P12t50 55

.Entergy                 Calculation VYC-2405 Rev. 0                 Page 54 of 85 Modifications in /home/schor/vyc-2120ccn/SENSITIVITY/SB0/drywell-SBO/SBO-drywell2-Leak-4 Mar/14/2005 17:49:07 GOTHIC Version 7.0p2(QA) - Apr1l 2002 File: /hure/schor/vyc-2120ccn/SENSITIVITY/SBO/drywell-SBO/SBO-drywell2-Leak-45-sensy Function 10 Coolda-i Pressure Ind. Var.: Time (sec)
Dep. Var.: Pressure (psia)
Ind. Var. Dep. Var. Ind. Var. Dep. Var.
: 0.      1078.5         600. 1014.5 1200.       953.5        1800. 895.3 2400.       839.8        3000.      787.
3600.       736.8        4200. 689.1 4800.       643.7        5400. 600.7 6000.       559.9        6600. 521.3 7200.       484.8        7800. 450.2 8400.       417 6        9000. 386.8 9600.       357.8       10200. 330.5 10800.       304.9       11400. 280.8 12000.       258.2       12600. 237.1 13200.       217.3       13800. 198.9 14400.       181.7       15000. 165.6 15600:       150.8       16200. 136.9 16800.       124.2       17400. 112.3 18000.       101.4       18600.      91.3 19200.         82.1      19800.      73.6 20400.         65.8      21000.      58.7 21600.         52.3      22200.      46.4 1000000.        46.4

  -Enlergy                  Calculation VYC-2405 Rev. 0                Page 55 of 85    -
6.4.2 Results Case SBO-drvwell2-Leak-45-sensy Figure 30 through Figure 38 present the main parameters for the case SBO-drypvell2-Leak sensy. Figure 30 presents the drywell temperature. The maximum drywell temperature is about 293 0F and is reached at the end of the run (4 hours). The run was not extended beyond the 4 hours even though the drywell temperature continues to increase because at 4 hours the RHR pump is available for drywell spray, if needed. The mission time of 2 hour and .10 minutes is.
achieved. The drywell heatup rate is arrested due to depressurization; however the leak brings enough energy from the vessel to continue the heatup. At 7800 seconds the drywell temperature is about 290 TF, well below the EQ limit of 325 F .
Figure 31 presents the containment pressure. The drywell pressure is high enough to allow for drywell spray after one hour into the transient, if needed. The available water to spray is the Diesel fire pump (Reference 30) per Appendix M of OE 3107 (Reference 35) and it takes about one hour for aligning the fire pump for drywell spray. The drywell temperature does not exceed the EQ drywell temperature limit and the drywell shell temperature.stays below the limit of 281
 'F for the mission time of 2 hour and 10 minutes hence spray from Diesel fire pump is not needed.
The vacuum breakers do not open during the time of interest.
At 4 hours the torus pressure is about 25 psig and increasing, close to the PSP limit of 27 psig.
However at this time the RHR pump is available for containment spray.
Figure 32 presents the RPV pressure. At one hour into the event it is assumed that the operators start depressurization. The vessel pressure during the 4 hours of the run time does not reach the shutoff pressure for the HPCI pumps, so at 4 hours the HPCI pumps still inject to maintain inventory. At 4 hour into the event only about 540000 lb were injected from CST (Figure 34). At this time the RPV is not depressurized, and the HPCI pump continues to inject.
Figure 33 presents the RPV level. The core stays covered and HPCI maintains inventory for the duration of the analyses. There is no need to continue the calculation beyond 4 hours because the-coping time of 2 hours was demonstrated.
Figure. 35 shows that the leak is maintain constant for the duration of the transient.
Figure 36, Figure 37, and Figure 38 presents the drywell liner temperature. The drywell liner stays below 280 TF for the 4 hours analyzed. After 2 hours .and 10 minutes the low pressure pumps are available for suppression pool cooling, drywell spray and maintaining vessel inventory.
The suppression pool temperature is not a parameter of importance for this calculation. In Reference 1 it was shown that the suppression pool temperature is lower for the cases with leak and lower for earlier depressurization hence the maximum suppression pool temperature will be lower that 182.2 "F, calculated in Reference 1.

Entergy                     Calculation VYC-2405 Rev. 0                Page 56 of 85     N.
SHO - drywell2-Leak-45-sensitivities Mar/09/2005 18:54:07 GOTHIC Version 7.0p2(QA) - April 2002
     ;      Pile: /home/schor/vyc-2120ccn/SENSITIVITY/SBO/drywel    -SBO/SBO-drywell2-Le I                IryweU Tesptratlare DzywtV1 21       rature T1    TLI 5.
                 .S IZ Time (sea)
               ,, matl      ,    "W69120*3  IS15 ,64 Figure 30 -Drywell Temperature -Case SBO-drywell2-Leak-45-sensy SO        - dryvfell2-Leak-45-sensitivities Mar/14/2005 17:59:16 GOTHIC Version 7.0p2(QA) - April 2002 File: /hboe/schor/vyc-2120ccn/SENSITiVITY/SBO/drywell-SBO/SBO-dlywell2-Le 3               contaiment ?rtssoxt o                                       !
In     t    2   1 9.A TAt (SeC) arst             pq..eS2BO     15,123,4 Figure 31 -Containment Pressure - Case SBO-drywell2-Leak-45-sensy

WEntergy                       Calculation VYC-2405 Rev. O                Page 57 of 85  .
SBO - drywell2-Lcak-45-sensitivities Mar/09/2005 18:59:19 GOTHIC Version 7.0p2(QA) - April 2002 File: /home/schor/vyc-2120ccn/SENSITIVITY/SBO/drywell-SBO/SBO-drywell2-Le
      .'      I24                PV Pressuvres                            -I gat        S Tint  ($to) arms T~fttfa      ox)
I-41104    IS,12,44 Figure 32- RPV Pressure -Case SBO-drywell2-Leak-45-sensy SBO - drywell2-Leakc-45-sensitivities Mar/09/2005 19:05:53 GOTHIC Version 7.0p2(QA) .- April 2002 Pile: /home/schor/vyc-2120ccn/SENSITIVITY/SBO/drywell-SB0/SBO-drywell2-Le 21 .              .??   Liquid Level LL4 Iir~
                                -                        TItse (see) fwrms71.S.)te1     H.,-IjV260% 15,1224 Figure 33 - RPV Level - Case SBO-drywell2-Leak-45-sensy

'Entergy                         Calculation VYC-2405 Rev. 0                               Page 58 of 85  -Y--
SBO - dxywell2-Leak-45-sensitivities Mar/C9/2005 19:14:51 GOTHIC version 7.Op2(QA) - April 2002 File: /home/schor/vyc-2120ccn/SZNSIT7IVITY/SBO/drywell-SDO/SBO-drywe               12-Le 9N       40               ThtjArated Ha'l t10W
                             .                  3.6            7.2        1.1            i Time  (see) 8*1     ?.!12(11) .,-y091295   13.1t.04 Figure 34 - Integrated HPCI Flow - Case SBO-drywell2-Leak-45-sensy SBO - drywell2-Leak-45-sensitivities Mar/09/2005 19:16:54 GOTHIC Version 7.Op2(QA) - April 2002 File: /home/schor/vyc-212Occn/SENSITIVlTY/SBO/drywell-SBO/SBO-drywel12-Le 43               SLea    Flow 11L4 7LI       rL20 a..
0n t          ..
IA
                         *          .       3.      .       .4         .s     11.2 .22.8 Tine (sec)
Zest  ?.ei2(fi) Mtr&sect;/O2f2C5 X5.12."4 Figure 35 - Leak Flow -Case SBO-drywell2-Leak-45-sensy

 - Entergy                      Calculation VYC-2405 Rev. 0         -      Page 59 of 85    -
SBO - drywcll2-Leak-45-sensitivities Mar/09/2005 19:06:59 GOTHIC Version 7.0p2(QA) - April 2002 Pile: /home/schor/vyc-2l20ccn/SENSITIVITY/SBO/drywell-SBO/SBO-dry-well2-Le Surface Traperatore TDS Ta6 S3?
I.- _
o
* a SA  (set) 4UeJ 7.A,2(9A    P..,1U112Se5 1.12.,4 Figure 36 - Surface Temperature, Heat Structures 5,6,7 - SBO-drywell2-Leak                                                             sensy SBO - drywell2-Leak-45-sensitivities Mar/09/2005 19:08:37 GOTHIC Version 7.0p2(QA) - April 2002 File: /hoete/schor/vyc-212occn/SENSITIVITY/SBO/drywell-SB0/SBO-drywell2-Le 51              5srface Tmperatere 138 133 Tale Tall 3                               7 2
                                                        ?S    (Sao)
                  $WM* 7.6t2mA) er/491260       15.1264 Figure 37 Surface Temperature, Heat Structures 8,9,10,11 - SBO-drywell2-Leak                                                             sensy

Entergy                     Calculation VYC-2405 Rev. 0              Page 60 of 85 SBO - drywell2-Leak-45-sensitivities Mar/09/2005 19:11:15 GOTHIC Version 7.0p2(QA) - April 2002 File: /hcze/schor/vyc-2120ccn/SENSITIVITY/SBO/drywell-SEO/SBO-drywell2-Le surface teaperature Th13 TM1
                       .           _S_              6
* C                ,           bo TIAC (see)
Cmii1.)iUA)   4.z/,It*5 M22,1I4 Figure 38 - Surface Temperature, Heat Structures 12,13 - SBO-drywell2-Leak                                                        sensy

:Entergy                    Calculation VYC-2405 Rev. 0               Page 61 of'85' 6.5 Case SBO-drvwell-comments 6.5.1 Model Modifications This case addresses the reviewer comments and also some discrepancies found during the documentation. The following changes are being made:
-change the initial temperature for Heat Structures 14 from 160 0F to 170 'F.
-change the Knverse in junction 3 to 3.93 from 3.964.
-change the flow area of the valve V3 to 15.63 ft2 , same as the flow path flow area
-change the surface area of the concrete pedestal to 2068 ft2 The changes are made to case 2 but it could be done to any of the other cases.
Table 11 presents the modifications made to file SBO-dryIvell2-80-sensy2-NoLeak to create SBO-drywel12-comments.
Table 11 Comparison between SBO-dryvell-corments vs SBO-drywell2-80-sensy2-NoLeak.

Entergy         -            Calculation VYC-2405 Rev. 0 Page 62 of 85 Modifications in /home/schor/vyc-2120ccn/SENSITIVITY/SBO/dry3el1-SBO/SBO-drywell-coment Mar/15/2005 14:07:44 GOTHIC Version 7.0p2(QA) - April 2002 File: /howe/schor/vyc-212Occn/SENSITIVITY/SBO/drywell-SBOlSBO-drywell-comments Flow Paths - Table 3 Flow FWd.           Rev.            Critical   Exit     Drop Path Loss           Loss     Ca.      Flow     Loss   Breakup r Coeff.       Coeff. Opt. Model   Coeff. Model
        .1    4.2 1243   4.2243     ON. TABLES       1. OFF 2          1.      0.78    ON    TABLES       1. OFF 3     3..964     '3.93     ON      OFF        0. OFF 4          0.             OFF    TABLES       1. OFF 5                         OFF      OFF        0. OFF
: 6.                        OFF      OFF        0. OFF 7                         OFF      OFF        0. OFF 8                         OFF      OFF        0. OFF
    . 9                            OFF     -OFF        0. OFF 10                          OFF    TABLES       0. OFF 11'                         OFF    TABLES       0. OFF 12-                         OFF      OFF        0. OFF 13                          OFF      OFF        0. OFF 14                          OFF      OFF        0. OFF 15                          OFF      OFF        0. OFF 16                  le+18 OFF        OFF        0. OFF 17                          OFF      OFF         0. OFF 18                          OFF      OFF        0. OFF
      .19                          OFF      OFF         0. OFF 20                          OFF      OFF        0. OFF 21          1.5        1.5 OFF       OFF         0. OFF Thermal Conductors - Table 1 Cond                         Vol HT     Vol   HT Cond S. A. Init.
         #    Description          A Co        B   Co Type (ft2)     T.(F) Or 1 Steam E'xposure         4     1     4     1   1  2965.72 609.23 1 2 Liquid Exposure         4     2     4     2   1  11521.8 647.4 1 3 Torus, Vapor            2     3     2     5   2  13553.7    90. I 4 Torus, Liquid           2     4     2     5   2  13553.7    90. 1 5 Lower Drywell           1     5     1     6   3  1856.24 170. I 6 Lower Drywell           1     5     1     6   4  2041 28 170- I 7 *Middle.Drywell         1     5      .1   6   5 3802.73 170'. I 8 Middle Drywell          1     5     1     6   6   780.68 170. .I 9 'Middle Drywell         1     5     1     6   5 1250.47    170. 1 10 Middle drywell           1     5     1     6   7 1898.24    170. 1 11 Middle DrpRl1            1     5     1     6   8 1114.72    170.. 1 12 Top Drywell              1     5     1     6 13    783.45   170. 1 13 Too Drywell              1     5     1    '7 14    1718.3   170. 1 14 RRUs                     1     5     1     6 11    1272.8   170. 1 15 Vent Pipes               3     5     3     6 11    2885.7   160. I 16 Concrete Shield          1     5     1     6 12     2068. 152. 1 Modifications in    i /hane/schor/vyc-2l2Occn/SENSITIVITY/SBO/drywell-SBO/SBO-drywell-ccmnent Mar/15/2005 14:07:44 GOTHIC Version 7.0p2(QA) - April 2002 File: /home/schor/vyc-2120ccn/SENSITIVITY/SBO/drywll-SBO/SBO-drywell-coments Valve/Door Types Valve                    Stem Loss        Flow Type       Valve      Travel Coeff. Area F*       Option       Curve Curve      (ft2)
I     QICK OPEN         0       0 2    CHECK CALVE        0      15       3.141 3    QUICK CLOSE         O      0       15.63

:;        Ia
 -"Entergy         .               Calculation VYC-2405 Rev. 0              Page 63 of 85 65.2 Case SBO-dr'well-commeznts Results The results of this case are presented in Figure 39 through Figure 40. Figure 39 presents the drywell termperature. Due to the fact that the initial temperature for the drywell thermal conductors increase by 10 0F, the drywell temperature is increased from 289.4 &Fto 295.2 T.
The containment pressure (Figure 40) is identical to the Case 2, hence the changes in the vacuum breaker inputs have no effect on results, as described in section 6.2.1.
Figure 41, Figure 42, Figure 44 presents the drywell liner temperature. The drywell liner stays below 260 0F for the 7 hours analyzed. After 2 hours and 10 minutes the low pressure pumps are available for suppression pool cooling, drywell spray and maintaining vessel inventory.
SB0 - 80F-Noleak-drywell2-sensitivities -coazents Mar/15/2005 14:03:50 GOTHIC Version 7.Dp2(QA) - April 2002 File: /home/schor/vyc-2120ccn/SENSITIVITY/SBC/drywellSBO/SBO-drywell -cOmD S5               DryvelL Texperature TV17L1l
                            . I_
                             *1 go  C.
wT CI-             1 fig I .,_ EIvus..A       Jwf.. ..
M.2 Xle3 aL (sac)
Figure 39- Drywell Temperature - Case SBO-drywell-comments

Eniergy                         CalculatiorinWC-2405 Rev. 0                  Page 64 of'85 .-
SBO - S0P-Noleak-drywrell2-sensitivities-comr       ents Mar/15/2005 14:05:35 GOTHIC Version 7.0p2(QA)               - April 2002 Pile: Tloine/Ichor/vyc-2120cen/SENSlTIVITY/SBo/drywell-SDO/SBO-drywell-com
         %      3                  Containment Preseure PRIPRfl p.2 Time (sec)        Xle3 Figure 40- Containment Pressure - Case SBO-drywell-comments SBO - 80P-Noleak-drywel2-Bensitivities-cornents Mar/15/2005 14:04:06 GOTHIC Version 7-Cp2(QA) - April 2002 File: /boxne/schor/vyc-212Dccn/SENSITIVITY/SBO/dryvell-SBO/SBO-dxywell-com Surface Temperature f5 TE16TH oB76B
                  .-          o-C F11 a
t.2 Time (3ec)         Xle3
                     ..& T.ft    a f..M24     ..,,4_.5 . ... .
Figure 41 - Surface Temperature, Heat Structures 5,6,7 - SBO-drywell-comments

E-Lntergy.                        Calculation VYC-2405 Rev. 0                                Page 65 of 85 NY-SBO - 60P-Noleak-drywe1l2-sensitivities-cOcmeants Mar/15/2005 14:04:52 GOTHIC Version 7.Op2(QA) - April 2002 File: /bone/schor/vyc-212 Occn/SENSITIVITY/SBO/drywell-SBO/SBO-drywell-cotn 51                   Surface Texperature H8e  T9*T10 11 N3 :
14                f
                              -U             3.6     7.2-     10.T     (14e4            25.2 XIe3
                                                      -       Tize (aec)
               .-   il .4      {1       .-   MI526fit I~t.4A.      ..      .
Figure 42 - Surface Temperature, Heat Structures.8,9,10,11 - SBO-drywell-comments SBO - BOP-Koleak-drywell2-sensitivities-comments Mar/15/2005 14:05:12.
GOTHIC Version 7.0p2f(A) - April 2002 File: /bomne/schor/vyc-2120ccn/SENSITIVITY/SBO/drywell-SBO/SBO-drywell-cocm 52                   Surface temperature CD                                                      -
N         t                                      BeatSlab 12 CD 14s4e                                     (ac) Eeat     13
('4
('           3.           .       814         .. 18... Z.2 Time (3ec)               X1e3
              .---     1-1          ---  -------  --------
Figure 43 - Surface Temperature, Heat Structures 8,9,10,11 - SBO-drywell-comments

i-Entergy.
Calculation VYC-2405 Rev. 0                  Page 66of 85 SBO - SOP-Noleak-drywell2-sonsitivities-comments Mar/15/2005 14:05:12 GOTHIC Version 7.0p2(QA) - April 2002 File: /bome/schor/vyc-2120ccn/SEMSITIVITY/SBO/drywell-SBO/SBO-dryvell-com 52         Surface temperature
                        .213Tn12 a--
C3 CO o       AdHeat                Slab 12 8 N. C3 o                      eatL   1
: 3.       '     114.4       18   '1. 5.2 Time (sec)              Zle3 Figure 44- Surface Temperature Heat Structures 12, 13 - Case SBO-drywell-comments

4 $ *   .-  zw . .- .    . ..                                        .. .. ..  . - .. ..
 *Enfegy                              Calculation VYC-2405 Rev; 0                 Page 67 of85 7.0 Results and Conclusions Assuminge Station Blackout with RPV depressurization (cooldown) at I hour after the event the following results and conclusions arc found:
: 1) The drywell temperature for all cases analyzed stays below the EQ drywell temperature profile for the entire SBO coping period of 2 hours and the additional 10 minutes to power the low pressure pumps (i.e., the drywell temperature for all cases analyzed stays below 300 TF for more than 4 hours of transient).
: 2) The drywell liner temperature stays below the design temperature of 281 TF for more than 4 hours after the SBO event.
: 3) The drywell pressure stays below the design pressure of 56 psig.
: 4) For 2 hour and 10 minutes the wetwell pressure stays below the PSP curve for all cases analyzed.
: 5) The analysis shows that there is no need to spray the drywell when in the unsafe region of DWSIL.
: 6) There is enough inventory in the CST to insure that the CST is not depleted before the time of low pressure pumps availability such that the core stays covered. A CST inventory of 75000 gallons was assumed.
: 7) The maxim suppression pool temperature for all cases stays below 182.2 'F.
: 8) The analysis predicts a conservatively high drywell temperature. Several factors contribute to this conservatism:
* The heat transfer from the vessel to the drywell is based on a constant heat transfer coefficient at normal operating differential temperatures. However, this heat transfer coefficient will vary with the temperature difference to the 1/4 power based on the dependence on the Grashoff number.
* The heat transfer to the drywell from the drywell heaters is not subtracted from the vessel.
* The reactor building side of the thermal conductors are considered adiabatic.
* A constant leakage is assumed; the leakage will decrease as the vessel is depressurized.
: 9) No restriction on the rate of cooldown is applied to protect the drywell temperature beyond the restriction of depressurization (cooldown) function of RHRSW temperature (Reference 1).
Note: "Unverified Assumptions" and "Affected Documents" items are being tracked via LO-VTYLO-2005-00135 (also see Section 4.1 and 4.2).

Entergy                   Calculation VYC-2405 Rev. 0              Page 68 of 85 8.0 References The References are divided into Section 8.1 and General References (Section 8.2). Section 8.1 includes all references.
8.1 Design Input References and General 

==References:==

1 VYC-2398 Rev 0 Torus Temperature Calculation for a Station Blackout Event at Extended Power Uprate, dated March 2005.
2   VY Tech Spec.
3   NUMARC 87-00, dated 11/20/87, including NRC accepted errata and Q&A from NUMARC seminars and Topical Report F.
4   VYC-2397, Maximum RHRSW Flow to RHRHX for SBO & Appendix R, dated 1/24/2005.
5 DRF 0000-0011-5646, OPL-4A (Containment Analysis Input Values) for Vermont Yankee Nuclear Power Station EPUIMELLLA+, dated 2/6/03.
6   Calculation VYC-1628 Rev 0, Torus Temperature and Pressure Response to Large Break LOCA and MSLB Accident Scenarios, dated 4/27/98.
7   George, T. L., et. al, GOTHIC Containment Analysis Package, Version 7.0, July 2001.
8   VYC-2208 "GOTHIC 7.0 Code Installation Validation and Verification at VY", dated July 18, 2002.
9 VYC-1457, VY Containment Heatup Analysis - Appendix R Application, dated 8/19/96.
10 GE Design Specification, No. 22A1 184, "Drywell Atmosphere Cooling System", Table I, Drywell Cooling Load Summary, #8 on Sheet 9.
11 VYC-1254 Rev 3, Containment and RPV Volume Calculations, dated 5/21/98.
12 VYC-1850 Rev 1, OPL-4A Input Preparation, dated 6/22199.
13 GE Design Specification # 22A1182 RevI        "Protective Coatings-Special" 14 Drawing G191526 Rev 2.
15 J. P. Holman, "Heat Transfer", McGraw Hill Book Company, 1981.
16 Standard Review Plan Branch Technical Position CSB 6-1.
17 VYC-1628D Rev 0 CCN02, Torus Temperature and Pressure Response for to Appendix R and Station Blackout Scenarios, dated 06/23/2003.

5Enlergy                 Calculation VYC-2405 Rev: O             Page 69 of'85 18 VYC-2279, Evaluation of EPU Impact on Ambient Space Temperatures During Normal Operation, dated 8/26/2003.
19 EQ Manual Vol 1.
20 22A1265, Rev 1, Reactor Containment Design Specification, September 1969.
21 VY Drawing G191489 Rev 2.
22 VYC-1290 Rev 0, Vermont Yankee Post-LOCA Torus Temperature and RHR Heat Exchanger Evaluation, approved Augustl, 1994.
23 VYC-2045 Rev 0 Residual Heat Removal Heat Exchangers Fouling Factors and Projected Heat Rates for Cycle 21, dated 12/11/99.
24 GE-VYNPS-AEP-146, Letter Michael Dick (GE) to Craig Nichols (ENOI), VYNPS EPU Task T0400: Decay Heat for Containment Analysis dated March 10, 2003.
25 TE-2003-020, Feedwater Parameters for Power Uprate, April 14, 2003.
26 VY Memo VYS 2000/39, P A Rainey/T. P. Bowman to J. R. Lynch, "Torus Temperature/SW Design Temperature Recommendations", April 16, 2000.
27 VY UFSAR Rev. 19.
28 OP 0105 Rev 11, Reactor Operation.
29 Crane Technical Paper No. 410, Flow of Fluids through Valves, Fittings, and Pipe, 1976 Crane Co.
30 EOP-3 Primary Containment Control; Rev3, dated 10/19/02.
31 ASME Steam Tables -Third Edition, 1977.
32 VYC-2306, Torus Temperature for Appendix R events at EPU Conditions, dated 08/29/2003.
33 VYC-I 850A Rev 0, "OPL 4A Input Preparation", dated 7/28/99.
34 VYC-1 628B Rev 0, "Torus Temperature and Pressure Response to Small Break LOCA Scenarios, Model Development", dated 11/3/98.
35 OE-3107 Rev 17, EOP/SAG Appendices dated 04/29/2004.
36 OT 3122 Rev 19, Loss of Normal Power, dated 04/18/2000.
37 EOP-1 Rev2 RPV Control, dated 10/19/02

Entergy.                CalculationrVYC-2405 Rev. 0           Page 70 of 85 8.1 General References Georgejy. L., et. al, GOTHIC Containment Analysis Package, Version 7.0, July 2001 GOTHIC 7.0 Code Installation Validation and Verification at VY, dated July 18, 2002.
ASME Steam Tables -Third Edition, 1977 J. P. Holman, "Heat Transfer", McGraw Hill Book Company, 1981 Crane Technical Paper No. 410, Flow of Fluids through Valves, Fittings, and Pipe, 1976 Crane Co.

Entergy                      Calculation VYC-2405 Rev. 0                      Page 71 of 85     *w>
ATTACHMENT 9.10                                                                  COMPUTER RUN 

==SUMMARY==
SHEET COMPUTER RUN 

==SUMMARY==
SHEET Page I of I Calculation No. VYC-2405                               Revision 0                        Date 16 March 2005 Sheet I of 1

==Subject:==
Drywell Temperature Calculation for a Station Blackout Event at Extended Power Uprate Code GOTHIC V7.0p2                            Catalog No.02543                           Version 7.0 SQA Classification Level A Run Title (variable, described in Section 6.0 and Att. A)
Run No. No                                Run Date                              By Output Use:                 5    Variable Values As Noted                3 Plot Attached 0    Disk.                                   D File No.
Description Of Output:
* Figures in text.
* Input file on Disk.
* Multiple cases were run, all are described in Section 6.0.
* One case (base case) is attached in attachment A).
* The Figures in text, for each case, have the date of the run & the run name.
Comments: None (Attached additional pages if necessary)
Review:               0  Information Entered Above is Accurate 0  Input Entry Accurate 0  Code Properly Executed (Based on User Manual) 0  Output Accurately Extracted or Location Specified Reviewer Comments          /lo.t'       e Preparer (Print/Sign)       Date                       Reviewer (Print/Sign)             Date Liliane Schor                                           Alan L. Robertshaw 3l15105                                                                            3//0f

' Entergy       -
Calculation VYC-2405 Rev. 0                                      Page 72 of 85 Calculation Impact Review Pages (ENN-DC-126 Attachment 9.7)
From System Engineering NUCLEAR                        QvAMRA                  ENN-DC-126         REV.4
       -= EnteW
       .:        tA  .E r    MANAGaEENT..
NUAL        N.                                         PAGE 35 OF 57 ArrAcENr 9.7                                                             CAuLiAnom luPAcT REviEw PAGE CALCULATION IUPACT REVIEW PAGE Date: 14 Febsiary 2005                                                                     aR    .      3NQR (Note: X Indicates required distribution)
To: __Mechanical Engineering                   X Licensing                       X  Operations I&C Engineertng                    -    Elect Maintenance                 Chemistry
_Electrical Engineering              -      I&C Maintenance                   HP/Radieolgical
* _ CIM Engineering                      _ Mech Maintenance                -    Conputer Applications X Systern Engineering                -      Component Engineering        _    Rad Engineering Reactor Engineering              -      Program Engineering               ISI Engineering X DBD Owner RHR. P. Perez            -      Nuclear Engineering          -    [ST Engineering BXBD Owner SA. P. Perez          _      E                            _    PSA.:
(Name)                              (Other)                  --
Fromr: Llllnre Schor 802-451-3013 (Originator Print Name and Phone extension)
Calculation No.: VYC-2405                                            Revision No. Q Tite:     DrvweD Temperature Calculation tor a Station Blackout Event at Extended Power Upzrate.

==Reference:==
WA Date'Response Required       16 Februarv 2005 MESSAGE: Work organizations are requested to review the subject calculation (parts attached) to-Identify Impacted calculations, procedures, Technical Speckiatlons, FSAR sections, other design documents (e.g.. EQ files.DBD, Appendr          R. ISIAST, PSA, MOVsIAOVs, etc.) and other documents.
which must be updated because of the calculation results. Also provide the name of tie Indh/ldual responsible for the action and the tracking number. The tracking Rem should Incude a requirement to ensure that any ER Implementation associated with the Item is completed prior to revising the Impacted document. Sign and return the form to the originator.
IMPACT     REVIEW    RESULTS:
Affected Documents              Responsible              Tracking                       Remarks Individual             Number r-0oP S-3~        ,  J Responding SupervisorfManager (or designee):            &'.i       Ljecsw -//oS           4 J       .
Name/Signature                                   Dale

t w fl
                                                                       .-                           I ... I' .
- .. rEntergy.                           Calculation VYC-2405 Rev. 0                                          Page 73 of 85     ;-
Calculation Impact Review Paues (ENN-DC-126 Attachment 9.7). Continued From System Engineering NUCLEAR                       (QUMMRIELZATD            ENN.DC126          REV. 4
                  .         .nery           MANUAL                                    USE0ADOrAL       PAGE 350F57 ATTACHMENT 9.7                                                              CALCULATI       IMPACT REvcw PAGE CALCULATION IMPACT REVIEW PAGE Date: 14 February 2005                       '                                                 R .           NOR (Note: X Indicates required distnbution)
To: _       Mechanical Engineering              _XjLlcensing .. -               _      Operations
                     -     I&C Engineering      .             __Elect Maintenance        .      -. Chemistry
_ Electrical Engineerlng                 - I&C Maintenance               __HP/Radiological
                    -      Civil Engineering                      Mech Maintenanoe             -      Computer Applications
_XSyslem Engineering                       Component Engineering        -      Rad Engineering
__Reactor Engineering             .        _ Program Engineering               - ISI Engineering
* X DBD Owner RHR. P. Perez                - Nuclear Engineering              - IST Engineering X DBD Owner SA. P.- Perez              __EO                .            _ PSA (Name)                               (Oter)
From: Liliane Schor 82-451-3013
                               *   (Originator Print Name and Phone cxtension)
Calcutatiop No.: VYC-2405                                                Revislon No. 0 Tite.      Drywel Temoxatur         Calcuation for a Statfon Blackout Event at Externded Power Uriate.

==Reference:==
WA Date Response Sequired:         16 February 2005 MESSAGE. Work orgarnzations arm requested to review the subject calculation (parts attached) to ideny irmcted calculations. procedures, Technical Speciriations. FSAR sections other design dctirents (e., EO Oes, DBD, Appendbx R. IISVIST. PSA, MOWstAOVs, etc.) and other documents, which must be updated because of the calcubaSion resuLs. Also provide the rame of the individual responsible for the action and the tracldrig nurmber. Jhe tracking Item should include a requirement to ensure that any ER lmplementation'msaociated with the kern Is completed prior to revising the
               *Ipacted docurnenL Sign and return the form to the originator.
IMPACT REVIEW RESULTS:
Affected Documents               Responsible            Tracking                        Remarks Individual            Number OP 4032                             S. Jonasch                                    See attadhed comnents Responding SupervisortManager (or designee): Stephen Jonasch                                               3160 NametSignature                                     Date

Eniergy                               Calculation VYC-2405 Rev.O                                     Page 74 of 85     .-
Calculation Impact Review Pages (ENN-DC-126 Attachment 9.7). Continued From RHR and SA DBD Owner ATTACHMENT 9.7                                                                 CALCULATION IMPACT REVWW PAGE CALCULATION IMPACT REVIEW PAGE Date: 14 Februarv 2005                         ;DR                                             O            NOR (Note: X indicates required distribution)
      - To:  _     Mechanical Engineering               -,X...Licensing
* X.Operations
* I&C Engineering                             Elect Maintenance             -  Chemistry Electrical Engineering                 __I&C Maintenance                 __HP/Radiological Civil Engineering                 . __Meeh Maintenance                __Computer Applications X   System Ergineering                         Component Engineering            Rad Engineering
             =-Reactor Engineering                        -    Program Engineering            _ ISI Engineering X DBD Owner RHR. P. Perez                 __Nuclear Engineering              _ IST Engineering X DBD Owner SA. P. Perez                  _    EQ           *_                  PSA (Name)                .             (Other)
From: Ullane Schor 802-45143013 (Originator Print Name and Phone extehsion)
Calculation No: VYC-2405
* Revision No. 0 Title,    Drywell Temperature Calulation or Station 3lackout Event at Extended Poweir Uorate.

==Reference:==
NIA
* Date Response Required:             16 Februarv 2005 MESSAGE: Work organizations pre requested to review the pubject calculation (parts attached) to identify impacted calculations, procedures, Technical Speeifications, FSAR sections. other design
       .documents:(e.g., EQ files, DBD, Appendixr               ISI/iST, PSA, MOVs/AOVs, etc.) and other documents, whch rmust be updated because of the calcutation resuls. Also provide the name of the individual                .
        *responsblefor the action and the tracking number. The traciog Item should Include a requirement to ensure that any ER implenerdtati6n associated with the Item Is completed prior to revising the
       .inpacted document Sign and returi the formito the originator.
IMPACT REVIEW RESULTS:
Aftfected Documents                 Responsible          Tracking                      Remarks
                .                                   Idivivdual           Number
      . SA-DBD                                     P. Pertz          LO-VTYLO-         May need update for Drywell -
                                                                     .2005-00135       SBO assumptions I Methodology
             ._    -___._.of_'_             _                 _                            VYC-2405.
Responding Supervisor/Manager (or designee):                 _______              __ __        ___       ___
* Namrnwg"nhhre,-     -                         Crate

Entergty          -               Calculation VYC-2405 Rev. 0                                        Page 75 of 85N. -
Calculation Impact Review Pages (ENN-DC-126 Attachment 9.7), Continued From Operations - EPU Engineering ATrACHMENT S.7                                                             CALCutArlnn    IuPjtT R~vFwPI~&e CALCULATION IMPACT REVIEW PAGE Date: 14 February 2005                                                                  MOR              0 NOR (Note: X indicates required distribution)
To: __Mechanical Engineering                   X tLcensing              .         X Operations
               -     I&C Engineering                          Eect Maintenance            _-Chemistry
_ Electrical Engineering                       &acMaintenance                _ HPlRadiological
               -     Civil Engineering                      Mech Maintenance                _ Computer Applications XSystem Engineering                    _Component Engineering                Rad Engineering
_ Reactor Engineering                      Program Engineering               ISI Engineering X DBD Owner 1RHR P. Perrez              = Nuclear Engineering -              IST Engineering
* X DBD Owner SA P. Perez               _    EQ                             _  PSA (Name)                             (Other)
Frorn: tLlilane Schor 802-451-34 )13 (Orig.nator Print Name and Phone extension)
Calculation No.: WC-2405                                               Revision No. a

==Title:==
Drywell Temperature Calculation for a Station Blackout Event at Extended Power Uprate.

==Reference:==
NIA Date Response Required:         16 FebruarY 2005 MESSAGE: Work orgarizations are requested to review the sutbect calculation (parts attached) to identify Impacted calculations. procedures. Tedrual Specifications. .FSAR sections, other design documents (e.g.. EQ ies. DBD. Appendix R. ISVIST. PSA. MOVsIAOVs. etc.) and other documents.
which must be updated because of the calculation results. Also provide the name of the individual responsible for the action and the'tracking number. The tracking item should indude a requirement to ensure that any ER implementation associated with te item Lscompleted prior to revising the Impacted documenL Sign and return the form to the originator.
IMPACT REVIEW         RESULTS:
Affected Documents                Responsible            Tracking                       Remarks
                            .              .        tndividual            Number EOP-3 Study Guide OT-3122 Lesson    Plan for EOP-3 ON-3147              .                                             ..
O N 3 143                            __  _   _  _  _ _7__                              _ _   _  _  _  __  _ _ _
Responding SupervlsorlManager (or desinee): Brvan-Croke                                          3116/2005 Name/Signature                                     Date

PNe7 Entergy                  *" Calculation VYC-2405 Rev. 0                                           Page 76 o6f 85'                 -.
Calculation Impact Review PaNes (ENN-DC-126 Attachment 9.7), Continued From Licensing NUCLEAR                      QUVAU" RuATD              EN      1-DC-Z6 R V.4 MEntefy               KANAGEME:NTI ElANUALe yoMMAAEOENL                      rUSE           PAGE 35 OF 57 ATTACHMENT 9.7                                                           CALCULA ON ImPACT REviEw PAGE CALCULATION IMPACT REVIEW PAGE Date; 14 February 2005                                                                SQR            [QNQR (Note: X Indicales required distribution)
To: _ Mechanical Engineering                  X LUcensing                       X Operations I&C Engineerkg                   _    Elect Maintenance             __Chemistry
__Electrical Engineering                   I&C Maintenance                   HP/Radiologlcal
                = ClvD Engineering                   __Mech Maintenance                 = Computer Aplications
__ System Engineering                __Component Engineering                Raod Engineering
_    Reactor Engineaftg.                -  Program Engireering            _ ISI Engineering X DBD Owner RHR P. Perez            _    Nuclear Engineering               IST Engineering X DBD Owner SA P. Perez             _- EQ                               =_PSA (Name)                              (Other)
Frorn: L113ans Schor 802-451.3013 (Originator Print Name and Phone extension)
Calculation No.: VYC-2405                                           Revision No. Q

==Title:==
_Drywell Temperature Calculation for a Station Blackout Event at Extended Power Uprate.

==Reference:==
WA Date Response Required:     16 February 2005 MESSAGE: Work rganizations are requested to review the subject calculation (parts attached) to Idenm Impacted calculions, procedures, Technical Spocifications, FSAR sections. other design documents (e.g.. EQ ess. DBO, Appendix R, 15SVIST, PSA, MOVs/AOVs. etc.) and other documents.
which must be updated bemuse of the calcutation'results. AJso providc the name of the Individual responsible for the action and the tracidnig nurmber. The tracking item should Include a requiement to ensure that any ER hlmiipmentalion associated uvth the Item Is completed prior to revising the Impacted document. Sign and return tho torm to tte originacor.
IMPACT REVIFW RFSLLTS:
Affected Docurnents             Responsible            Tracking                       Remarks Individual .          Number         _
              -       07-'7-           .                            _-e    2 '                                           __Z I        -                      I.                                                                                     ..
i2 6-OEZ-Date Responding SupervsortManager (or designee): -                   Qo   i  A    C Namelisqgature P St;{-,"' fydor 5<
9          '               ,c
                                                                       .      - ,,,-.           0p'       ___       -
* I Entergy - -                 Calculation VYC-2405 Rev. 0                              Page 77 of 85          -
Calculation Design Verification and Review (ENN-DC-134)
          'ENN-DC-134 REVISION I                ATrACHMENT 9.1                DESOGNVERInAcoN COVERPAGE DESIGN VERIFICATION COVER PAGE El   IP-2                OIP-3                   OlAF                   EPNPS          [VYY Document No. WC-2405                                          Revision O   Page 1 of 8 Tatle: Drywell Temperature Calculation for a Station Blackout Event al Extended Power Uprate M Quality Related        LiNon Quality Related DV Method:       [3Design Review -         QAlternate Calculation        OQualification Testing VERIFICATION                                        VERIFICATION COMPLETE AND COMMENTS REQUIREDDISC                     NE                    RESOLVED (DVprint, sign, and dac)
Ol                Electrical El               Mechanical El    . Instrument and Control O              (:Civil/Structural Design Eneineirinj!        Alan L Robertshaw.
                    .              FluidSystems                c        ,               7         2/11/o     <:
PrlntrSlgn After Comments Have Been Resolved Originator:
Liliane Schor Date:
3 1710 I      I Design Verification for VYC-2405 Page I of 8

; Entergy                               Calculation VYC-2405 Rev. 0                                             Page 78 of 85 -
       'ENN-DC-1 34 REVISION I                          ArrAcHMENT 9.7                    CALCuLA.nON DESIGN VERIricATroN CHECKUST IDENTIFICATION:                                                                                       DISCIPLINE:
Document

==Title:==
Drywell Temperature Calculation for a Station                                         EICiviV/Structural Blackout Event at Extended Power Uprate                                     [l Electrical Doc. No.:                 WC-2405                                   Rev. 0 QA Cat.                    E T& C Verifier:
Alan L. Robertshaw WrSin                                              I)=
I]El Mechanical Nuclear Mmag1eraubriatxi~ frarn                                                                                       s uipx         >"orptring          -                                                                           Other:
pa'rifiC16"mL r          n pirmicisi                   sip                                   Design Engineering.
          ] N/A                                                                                               Fluid Svstems METHOD OF VERTFICATION:
Design Review            0                       Alternate Calculations          El         QualificationTesi         El 1      Design Inputs -Were the Inputs corredy selected and                                                    Rerertcrc Incorporated Into the design?                                                         PagcNu. S&ctin s 0rVYC-2405 Design h-puts Include design baser, plart operational conditiors, perforrmance       OR requirements, regulatory requirements end comrmitments, codes, standards, feld data. etc. Al Information used as deain khuts shoutd have been                  Paragraph No.
reviewed    and approved by the responsible design organization, as 8pplicable.
Al hnpuAs   need to be retrlevable or excerpts of documrents used should be           Completion of the Referenre BDoxes is attached.                                                                            optional for ESquestions.
See site specific design hnput procedtres for guidance In Identeifyng Iputs.
Yes0                            No l                          NiA       l Verifier Comments:
Section 5.0, Input and Design Criteria, of VYC-2405, has been satisfactorily reviewed. Any identified Design Input needing verification is listed in Section 4.1 of VYC-2405 and tracked via LO-VTYL.O-2005-00135 (see Itern #2 ofthis Calculation Design Verification Checklist). All other Design Tnput 'has been verified in VYC-2405.
Resolution: None needed.
Attachment 9.7                                                            Calculation Design Verification Checklist Design Verification for VYC-2405 Page 2 of 8

Ad ~Entergy                      Calculation VYC-2405 Rev. 0                             Page 79 of 85 '
a      Assumptions - Have the assumptions been verified?                           'Reference Page No. Cacki,,Lon Suxnmry Pare aMdSect 4 1 orVYC-2405 YesL              No0                  N/A  LIOR Par..Sph No.
Vcriflcr Comments:
Section 4. 1, Assumptions that need Verification or Jmplementation, and LO-VTYLO-2005-00135 CA02, document identified "Unverified Assumptions" from VYC-2405 Drywell Temperature Calculation for a Station Blackout Event at Extended Power Uprate. The following lists the various "Unverified Assumptions' from WC-2405:
: 1)  2 hour restoration ofoutside power (coping time).
: 2)  10 minutes to start RHR through the RHRHX, 2 RHRSW pumps and CS.
: 3)  Acceptability of using 7500C0 gal from CST (change of level setpoint).
: 4)  Maximum CST temperature of .1357F.
: 5)  The depressurization raie function of Service Water temperature needs to be verified and proceduralized as follows:
                       - For SW> 757F; depressurize the vessel at 800F/hr or higher.
                       - For lower SW temperature (SW < 757F) no restrictions on depressurization rates.
Upon verification of these assumptions; the calculation should be revised to convert the assumptions to Design Input and the alciulationr Status should be changed.
Resolution:
LO-VTYLO-2005-00135 CA02 has bocn issued to track thcsc Unvcrificd Assumptions.
Upon verification of these assumptions, the calculation should be revised to convert the assumptions to Design Input and the calculation Status should be changed.
3      Quality Assurance - Isthe Ouarity level correct7Rccmcc
                                                                          .Piae No. ('over Saed or VYC-2405 Yes Z             No l                 N/A [3               OR Pamara;Ph No. _   _   _  _ _  _ _
Verifier Comments: WC-2405 is correctly designated "Quality Related.'
Resolution: None needed.
Attachment 9.7        .                              Calculation Design Verification Checklist Design Verification for WC-2405 Page 3 of 8

Entergy                       Calculation VYC-2405 Rev. 0                                  Page 80 of 85       -
4      Codes, Standards and Regulatory Requirements          -Rerrot Are the applicable codes, standards and regulatory               age N.n.
requirements, Including Issue and addenda properlyOR identified and are their requirements for design met?           Panpu  ph No.
Yes [0            No           -          N/A   Ell Verifier Comments: Appropriate use of requirements (inputs, assumptions, methodology) set forth in the VY Technical Specifications and UFSAR have been followed in VYC-2405 as needed.
Resolution: None needed.
: 5. Operating Experience - Have applicable construction and                             Reference operating experience been considered?                                   Page No.
OR Yes E                    No     l                N/A El                 ParamphNo.
Verifier Comments:
Consideration (discussions and reviews) has been given to various timelines for operator actions, equipment start times, cooldown rates, etc- Somne Unverified Assumptions exists which will rely on, in part, to VY operating experience(e.g. two hour restoration of outside power (coping time of 2 hours and 10 minutes), ten minutes to start RH1R flow through the RHRHTX, the use of 2 RHRSW pumps and CS, and the acceptability of using 75000 gal from CST (change of level setpoint). See Item #2, Assumptions, of this Calculation Design Verification Checklist.
Resolution: Consideration has been given to VY operating experience. Some Unverified Assumptions exist (for which commitments have been issued). See Resolution of Item #2, Assumptions, of this Calculation Design Verification Checklist.
Interfaces - Have the design Interface requirements been                           Rtfirenoe satisfied and documented?                                               Pare No              -_-----
                                                                    .0            OR Yes 1                    No   El                 N/A                  larranphNo.
P Verifier Comments:
The relationship between the Fluid Systems (Idesign group") and other organizations within VY have been satisfactorily met using the Calculation Impact Review Page. The Calculation Impact Review Page was completed by persons in the Operations Department, System Engineering, Licensing, and various DBD owners.
Resolution:
None Needed.
Attachment 9.7
* Calculation Design Verification Checklist Design Verification for VYC-2405 Page 4 of 8

AEke~g Calculation VYC-2405 Rev. 0                                   Page 81 of 85                  :
: 7. Methods -Was an appropriate analytical method used?                                     Rfrraur aVaaeN. Section 3 of VYC-2405 Yes0                    No     l                   N/Aa                  OR Panriph No.
Verifier Comments:
The GOTHIC code (Reference 7 & 8 of VYC-2405), Version 7.0p2 was selected for use in VYC-2405.
Resolution: None Needed.
: 8. Design Outputs - Is the output reasonable compared to the                              Refert._ce inputs?                                                                 PTe4NO,                           -...
Io Yes                    'No     l                  N/A                  IPartphNo.
Verifier Comments; The'output is reasonable compared to the inputs. Previous, similar Drywell SBO analyses are familiar with the preparer and reviewer ofihis SBO calculation and thus the output given in VYC-2405 was reasonable for the various changes and modifications made to the previous input.
Resolution: None Needed.
9.-   Acceptance Criteria -Are the acceptance criteria                                       R ererence incorporated inthe calculation sufficient to allow verification        Pa.gc No. Scoion2.1orWYC-240; that design requirements have been satisfactortly                       OR accomplished?
Pau giaph N o. _ _ _ _ _ _    _ _
Yes                    No    O                    NIA      L Verifier Comments:
Section 2.1 was added to include the Acceptance Criteria.
Resolution:
None Needed.
Attachment 9.7                                              Calculation Design Verification Checklist Design Verification for VYC-2405 Page 5 of 8

.Entergy                             Calculation VYC-2405 Rev. 0                                    '          Page 82          of185    --
1,. Records and Documentation - Are requirements for                                                       Reicirence
         . record preparation, review, approval, retention, etc.,                                 Pate No. Secion 4.2 of WC.2405 adequately specified?                                                                  OR Are al documents prepared In a clear leWtile manner suiitate ror rnicroflrning andfor   Pwugraph No. _   _   _  _   _ _ _
other docurnentation storage mnethod? Have el Impacted documents been Identifed for update?
Yes E                      No [li                       NIA     El Verifier Comments: VYC-2405 was prepared in a clear legible manner suitable for microfilming and/or other documentation storage method. All impacted documents (Design Output) have been identified for update in Section 4.2 of VYC-2405 and in LO-VTYLO-2005-00135 CA01.
Resolution: LO-VTYLO-2005-00135 CA01 has been issued to track these affected documents.
If any other documents are identified during the Calculatioi Impact Review process, other commitments will be generated.
: 11. Software Quality Assurance- For a calculation that utilized                                            Retemence software applications (e.g., GOTHIC, SYMCORD), was it                                  P'age We. Sectiop 3 of VYC-2405 property verified and validated in accordance vvith ENN IT-                            OR 104 or previous ste SQA Program?                                                       P=V*      Na            _____
Yes Ax                     No El                        NIA [E1 Verifier Comments: The GOTHIC code (Reference 7 & 8 of VYC-2405), Version 7 .Op2 was selected for use in VYC-2405. This code was used in similar SBO analysis (References I of VYC-2405). This specific version of the code has been installed and complies with the ENVY SQA procedures ENN-IT-104 (replaced VY procedure AP-6030) as documented in calculation VYC-2208 (Reference 8 of VYC-2405).
Resolution: None Needed.
OTHER COMMENTS See Attached list or General comnments from review of VYC-2405 RESOLUTIONS Attached General comments have been made to reviewer's satisfaction.
All comments for "NO" answers have been resolved satisfactorily.
Attachment 9.7                                                       Calculation Design Verification Checklist Design Verification for VyC-2405 Page 6 of 8

-E
 - ntergy                      Calculation VYC-2405 Rev. 0                          Page 83 of 85 General Comments ftrom Alan L. Robertshaw from Review of VYC-2405 (Attached to Calculation Design Verification Checklist)
* In Section 2.0, Add addition Section (2.1) entitled "Acceptance Criteria" and add appropriate acceptance criteria.
Done
* In Section 3.0, discussion of GOTIUC code and V&V calculation need appropriate references.
Done
* In Section 4.0:
: 1. For Assumption #4, please correct the imits for density (value is correct).
: 2. For Assumption #9, add that 61 gpm is from Reference 3.
: 3. Add VYC-2306 to Reference (from Diywell Free Volume of Table 3).
Added
* In Section 5.0:
: 1. In Section 5.3.1 show in more detail how total Heater ,5 loads are calculated.
Done
: 2. In Section 5.4, recommend addition of simple" drawing of Drywell Done
: 3. In Section 5.4 show that the total Surface Area adds up to the Value obtained from OPL-4a (i.e., Table A Total =15246.06 fl,, RRUs = 1272.S fl2, Vent Pipes = 2885.7 ft2, thus Total - 19405  a2 )
AdIded
: 4. Add to description of Item 8 'Side of Drywell Head."
: 5. Add to description of Item 9 'Top of Drywell Head.'
Done, both items
: 6. Is RRU area Reference I1 or 12 or other? Please correct if needed.
Both, added
: 7. Flow Path 21 Forward and Reverse loss coefficients should reference CRANE (or similar), data not found in steam table. Need to add CRANE to references Done
: 8. Please add additional information on new valve added (vacuum breaker, Valve R5.
e.g., size, type, etc.).
Done Design Verification for VYC-2405 Page 7 of 8

'~Entergy                      CalculatioriVYC-2405 Rev. 0                            Page 84 of 85
: 9. Also for Valve discussion, note that the valve area'is different from the line area used in Flow Path 21 Added
: 10. Initial Temperatures of the Thermal Conductors are not given (and reference).
Added I 1.Please flirther explain why a loss coefficient of 3.964 was used instead of the calculated value of 3.93 for the vacuum breaker flow path. Any impact on analysis?
I added this discussion in the document, I also said that the difference is less than /% and will makeno differcnce on the analysis. Anyway, was correctedin Case5.
In Section 6.0:
: 1. List Five case individually to add clarity.
Done
: 2. See minor grammatical / spelling concerns.noted in marked-up calc drafi Done
: 3. Section 6.2.1 please add info on FlowPath 21 (Leakage) area change.
1)one
* In Section 8.0, Reference Section:
: 1. Need to add Reference for ASME Steam Tables and one for CRANE Technical Paper (if needed)
Added
: 2. Are References 11 and 13 used?.
Yes, Refeience 71 and 72 are RRU references and 13 isfor thepaint(addedsome more discussion about the paint)
* In the GOTHIC Input Deck, Thermal Conductor #16, Concrete Shield, a surface area value of 2108 f17 was used. Avalue of 2068 fi2 was given in the text of VYC-2405 (Section 5.4). Both of these numbers are found on page 35 of WC-l 850, Rev. 1 for OPL4A preparation. The largervalue (2108f12) is fortotal surface area, the smaller value (2068 112) subtracts the "Slots." ThevalueusedinVYC-1850, Rev. I CCN-0l (OPL4A. Resolved for Analysis) uses the 2068 ft2 value. Please discuss this in the analysis and determine which value to use, any sensitivities, etc.
CaseS addresses this
* In Attachment A, it appears that Figure on page A14 isa duplicate figure and should be removed.
Removed Design Verification for WC-2405 Page 8 of 8

                                                                  .      -   .  .                         ,             .. __ il, -,   ' _     .
   'Entergy                            Calculation VYC-2405 Rev. 0                                Page 85 of 85 i-N:
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Wednesday. Marrh 16.2005 503:11 PM

BVY 05-072 Docket No. 50-271 Exhibit SPSB-C-52-5 Vermont Yankee Nuclear Power Station Proposed Technical Specification Change No. 263 - Supplement No. 30 Extended Power Uprate Response to Request for Additional Information Calculation VYC-2279, Rev.0 l Total number of pages in this Exhibit l I (excludina this cover sheet) is 14. l

This document contains Vermont Yankee proprietary information. This information may not be transmitted in whole or in part, to any other organization without permission of Vermont Yankee.
I                                                      VY CALCULATION TITLE PAGE VYC-2279                           0                             N/A                                          NIA VY Calculation Number             Revision Number                Vendor Calculation Number                    Revision Number

==Title:==
Evaluation       of EPU Impact on Ambient Space Temperatures During Normal Operation QAStatus:             0 SC ENNS            El OQA                Operating CycleNumber* NIA
* The Operating Cycle Number should only be entered here if the results of the calculation only apply during a specific operating cycle otherwise enter "N/A'.
Calculation Supports A Design Change/Specification? ED Yes El No                     VYDC 2003-008 VYDC/MMrM/Spec No.
Implementation Required?         0  Yes   El No                            Calculation Done as a Study Only? a Yes   0   No Safety Evaluation Numnber. N/A Superseded Calculation Number, Title and Revision: N/A For Revisions: List CCNs, lUs, or SAs incorporated/superseded by this revision: N/A Computer Code(s): None used Are thre open items in this calculadon/revision?      0 Yes 0 No Review and Approval: (Print and Si Preparer. Gene O'Brien Intcrdiscipline (Heat Balance, Att. B only) Preparer(s): N. Zervos Interdiscipline (Heat Balance, Att. B only) Review ): R Srini Independent RceviewersfAansour ojibian-Final Turnover to DCC (Section 2):
: 1)         All open items, if any, have been closed.
: 2)         Implementation Confirmation (Section 23.4)                                               Total No. Pages in Package 64 (including all attachments) o    Calculation accurately reflects existing plant configuration, (confirmation method indicated below)
O    Walkdown        El As-Built input review         0  Discussion with (Print Name)
OR E    N/A, calculation does not reflect existing plant configuration
: 3)         Resolution of documents identified in the Design Output Documents Section ofVYAPF 0017.07 has been initiated as required (Section 23.6,23.7)
_/                                          _I Printed Name                                Signature                Date Page I of      28         Pagess
* For calculations performod using AP 0017 this is the number of pages in the body of the calculation.                  VYAPF 0017.01 For vendor calculations, this is the number of pages of AP 0017 forms added.                                          AP 0017 Rev. 8 (Title page, review forms, data sheets, 50.59, etc.)                                                             Page 1 of I LPC #2

Calculation Number: VYC-2279                   Revision Number: 0                            CCN: 0                    Page 6 of 28 TABLE OF CONTENTS SECTION                                                                                                                            PAGE NO; VY CALCULATION TITLE PAGE .....................................                                                                                     I VY CALCULATION DATABASE INPUT FORM .....................................                                                                           2 TABLE OF CONTENTS................................................................................................................................6 I Calculation ......................                                                                                                               7 1.1 Objective ..................      ....                                                                                                      7 1.2 Summary of Results ......................                                                                                                   7 1.3. Method of Solution ......................                                                                                                  8 1.4 Inputs and Assumptions ......................                                                                                              10 1.4.1 Inputs .....................                                                                                                          10 1.4.2 Assumptions .....................                                                                                                     14 1.5 Calculation .....................                                                                                                          14 1.6 Results .....................                                                                                                              26 1.7 Conclusion .....................                                                                                                           27 Attachments .....................                                                                                                               28 Attachment A .....................                                                                                                                AI Attachment B.......................                                                                                                              BI Attachment C .....................                                                                                                               C1 Attachment D .....................                                                                                                               DI Attachment E .....................                                                                                                               El

CALCULATION NUMBER NrYC-2279 Rev. 0                 CCN 0                             Pmage 7 of 28 I Calculation 1.1     Objective The purpose of this calculation is to evaluate ambient temperature increases in several plant spaces following the increase in reactor power level to 102% of 120%, hereafter referred to as Extended Power Uprate (EPU). The EPU will increase core thermal power from the current licensed level of 1593 MWt to 1912 MWt. For bounding purposes, the 122% (1950.9 MWt) heat balances are used for EPU HVAC evaluations.
This calculation evaluates the EPU impact on ambient air temperature in the following buildings or areas during normal plant operation:
Reactor Building m Drywell Steam u nne i"    the-rReactor Building Areas Turbine Building
* Reactor Feed Pump Room
* Condensate Pump Room
* HP Heater Area (including steam line shelf containing main steam lines)
* LP Heater Area Note: Comments to this calculation provided by letters PUPVY-03-208 dated 7/16/03 and PUPVY 212 dated 7/18/03 have been reviewed and incorporated (see Attachment D).
Increases in area heat gain and ambient air temperatures, as a result of EPU, are predominately caused by increases in operating temperature of piping systems, and equipment, and air-cooled motors operating under increased loads. The preuprate piping system temperatures are selected or extrapolated from a PEPSE Heat Balance that is tuned to match preuprate (current) plant data. The EPU piping system temperatures are selected or extrapolated from a PEPSE Heat Balance that provides the most conservative results.
Affected areas are evaluated to determine the temperature gain due to increases in heat loss from piping and mechanical equipment.
1.2      Summary of Results The results of this calculation show the effects of the EPU in terms of increased ambient temperature and heat load are due to increased feedwater temperature, as well as increased horsepower from the condensate and feedwater pumps. The ambient temperature increases are specified in Section 1.6 (Results).

Calculation Number: VYC-2279                Reision Number: 0                    CCN: 0    Paae 8 of 28 1.3. menthod of Solution This calculation evaluates the temperature increase in a specified area using current fluid and ambient air temperatures and EPU fluid temperatures to predict the EPU ambient air temperature and corresponding temperature rise. PEPSE H-leat Balances at 100% CTP and EPU at 122% CTP (References 7 and 8) are used to obtain preuprate and EPU piping temperatures. The incre'ase in heat loss from piping is determined by comparing the ratio of "temperature differential between EPU pipe and area air temperatures" to "differential temperature between pre-uprate pipe and pre-uprate area ambient air".
The basis for using this scaling approach to determine increased heat loss from piping and equipment can be obtained by referencing the ASHRAE Fundamental Handbook (Reference 12) Section 20.
Formula (9) - AS1JRAE 20.9 is used for flat surfaces qj =(tjs-t os)/R Formula (1O) -ASHRAE 20.9 is used for cylindrical flat surfaces q,=(ti.-t O.)/[r. In (r/r,)]/k1 + [r5 In (r..rz)]/k2 Formula (I 1) -ASIJRAE 20.9 for determining heat flow per area of pipe surface
: q. =qs (r At ;)
Where q,= rate of heat transfer per unit area of outer surface of insulation q0 = rate of heat flow per unit area of pipe surface, Btu/(hr)(ft 2 )
R = surface to surface thermal resistance k = thermal conductivity of insulation at calculated mean temperature.
t= Temperature of inner surface t    = Temperature of outer surface r1 = inner radius of insulation ri, r2 = outer radius of intermediate insulation r1 = outer radius of insulation In = natural or Napemian logarithm For the purposes of this calculation it can be assumed that there is one layer of insulation, therefore Formula 10 can be simplified as follows:
q, =(tL,-t j) / [r. In (r1 /rj)]/kj The increase or delta in heat transfer per unit area of insulation can be stated as follows:
Aq =EPU [(tL,-t .)/ [r5 In (rl/ri)]/kl] / pre-EPU[(t1 5 -t j / [r, In (r1 /rj)]/k1 ]
There is no change in either: rl, r,, or r1 ,.

Calculation Number: VYC-2279             Revision Number: 0                 CCN: 0          Paae 9 of 28 Based on the predicted temperature increases in the various process streams it can be assumed that there is no appreciable change between the preuprate and EPU values for k.
Therefore Aq = EPU [(ti,-t S / pre-EPU[(tL,-t O,)
Present station operating ambient air temperatures are used in the evaluation. If operating data is not available, plant design area temperatures are used. An iterative process using an Excel spreadsheet is utilized. First, an EPU ambient air temperature is estimated. Next, the EPU area heat gain multiplier is obtained using the ratio of the EPU pipe / ambient air temperature difference to the preuprate (current) pipe / ambient air temperature difference as shown below:
(EPU pipe temperature - EPU ambient air temperature)              - EPU Area Heat Gain Multiplier (preuprate pipe temperature - preuprate ambient air temperature)
The EPU factor is obtained by subtracting I from the EPU heat gain multiplier.
The EPU factor is then multiplied by the preuprate temperature difference between the air in and out of a particular air handling unit to calculate the estimated EPU temperature rise. The temperature rise is then compared to the difference of estimated EPU and preuprate ambient air temperatures. If required, a new EPU ambient air temperature is estimated and the process repeated until the temperature rise is equal to the difference of estimated EPU and preuprate ambient air temperatures.
Feed water and Condensate pumps, flows from preuprate and EPU PEPSE Heat Balances along with app! i ate pump curves will be used to determine horsepower changes in the respective pump motors to determ the heat gain increase to the room. The percentage increase in heat gain to the room will be utilized wit     e temperature rise of ventilating/cooling air currently being supplied to the room or area being evaluated.
The heat load from the        ensate and feedwater pumps is evaluated by calculating brake horsepower (BHP) at preuprate and EPUws. BHP is calculated using the following equation from page B-9 of Reference 14:
Bhp (hp) = Q(gpm)*H(fi)* p(lbJft 3 ) / [2          0
* pump efficiency]
The flow, Q, is calculated from the mass flow rate cified on the heat balance using the fluid density, p, calculated at the average of pump inlet and outlet tem >atures. The pump head, H, and efficiency are obtained from the pump curve.
The heat generated is due to pump motor inefficiency and is ca           ted using Chapter 26, Equation 21 from Reference 12:
q (BTU/hr) = BHP (hp)*2545 (BTU/hr hp) *[100-% efficiency]/ % effici The heat load due to the pump motor inefficiency is calculated for preuprate and          conditions. The overall heat removal capability of the coolers at preuprate conditions is determined an          sensible heat load due to piping is obtained by subtracting pump heat load from cooler heat removal capac            The piping heat load at EPU is scaled due to the increase in fluid temperature and added to the EPU p           heat load to obtain total heat load for the room. The temperature increase across the coolers is calculated Wig Equation 39.6(b) from Reference 13:

Calculation Number: VYC-2279             Revision Number: 0                 CCN: 0          Page 10 of 28
;[Tin -             (Btu/hr) / [cfm
* 1.08 (Btu minl(ft3 hr 0F))]
Both the feedwater and conensate pump             e       aluated in this manner.
1.4      Inputs and Assumptions 1.4.1     Inputs The inputs for this evaluation are the fluid temperature in system piping and ambient room temperature of the areas considered. The pre EPU fluid temperatures are obtained from a PEPSE heat balance based upon current plant operating data adjusted to 5.00" Hg. condenser backpressure to obtain maximum fluid temperatures (Reference 7). For those cases where the fluid temperature is not explicitly listed, the temperature is obtained based upon the pressure and enthalpy listed using ASME steam tables (Reference 9). The EPU fluid temperatures are obtained from the 122% heat balance with a condenser pressure of 5.00" HIg (Reference 8) which provided the highest, and therefore conservative, temperatures.
The current ambient steam tunnel and pump room temperatures are obtained from HVAC system design criteria (Reference 5). The main steam tunnel design temperature is 1307F, the reactor feedwater pump room and condensate pump room design temperatures are 105'F. The current HP and LP heater area temperatures are 125TF as taken from the Environmental Qualification Program Manual (Reference 6, page 11). The 20'F Pre-EPU Vent/Cooling Air AT contained in the Table below for the HP and LP heater area spaces is based upon transfer air at 105TF.
The design conditions for the air handling units are obtained from Reference 4, except for TRU-5 and TSFIA/IB, which have their design conditions specified in Reference 19.
Table 1.4-1 Area                  Equip. ID            Flow, cfm I Tin(&deg;F)      Tout(&deg;F)      AT Reactor Feedwater Pump Room TRU-1,-2,-3,-4                 16,750        105       85            20 Condensate Pump Room                 TRU-5                21,400         105       85            20 Condensate Pump Room                 TSF-IA/1B            5,000          90        105           15 Drywvell                             RRU-l,-2,-3,-4        16,000      [135        97            38 Main Steam Tunnel                    RRU-17A, -17B* 5000               t130       105            25
*Per Reference 18, the coils for these coolers are incorrectly piped as parallel  flow rather than counter flow The design inputs for piping are summarized in the table below:
Table 1.4-2 Pre-EPU pipe Temp      EPU pipe Temp          Pre-EPU ambient air     Pre-EPU Vent/Cig
                                   'F (Ref. 7)           'F (Ref. 8)             Temp 'F                Air AT OF tIP Heater Are ESS to_3 FWHI                                               403.7                  125.0          l       20.0 FWHI Shell                            383.3          03403.7                       125.0          l       20.0 FW to FWH1                            330.7                 346.4                  125.0                  20.0 FW Ivg FWH1                           374.4                 392.6                  125.0                  20.0 Drains lvg FWHI                       343.6                 356.9                  125.0                  20.0

Calculation Number: VYC-2279       Revision Number: 0        CCN: 0        Pagc 11 of 28 Pre-EPU pipe Temp     EPU pipe Temp Pre-EPU ambient air   Pre-EPU Vent/Clg
                             &deg;F (Ref. 7)         &deg;F (Ref. 8)      Temp &deg;F                Air AT &deg;F ESSvto FWH2                    338.0                358.1           125.0                  20.0 FWH2      ell                  338.0                358.1           125.0                  20.0 FW to FW sH                    299.3                312.5           125.0                  20.0 FW lvg FWH2                    330.7                346.4           125.0                  20.0 Drains lvg FWH2                309.7                322.6           125.0                  20.0 LP  hleater Area                              _                                    _
ESS to FWH13                                  _     321.2
__.4             125.0        l         20.0 FWH13 Shell                    305.4                321.2           125.0                  20.0 CND to FWH3                    227.6                238.4           125.0                  20.0 CND lvg FWH3                   296.6            \309.8              125.0                  20.0 Drains lvg FWI13               235.0                24 4            125.0                  20.0 ESS to FWI-I4                  239.0                252.1           125.0                  20.0 FWH4 Shell                     239.0                252.1           125.0                  20.0 CND to FWH4                     167.9               180.9           125.0                  20.0 CND lvg FWH4                   227.6                238.4           12 0                   20.0 Drains lvg FWH4                175.1                189.1           125.0                  20.0 ESS to FWH5                     173.9         _     181.8           125.0                  20.0 FWH5 Shell                      173.9               181.8           125.0                  20.0 CND to FWH5                     135.1               134.7           125.0                 20.0 CND lvg FWH5                    167.9               180.9           125.0        lO Drains lvgFWH5                  145.1               141.8           125.0                  20.0 CND Pump Room                                                                         .-
CND to CNP                      133.8               133.8           105.0                  20.0

Calculation Number: VYC-2279                 Revision Number: 0              CCN: 0            Page 1 2 of 28 Prc-EPU pipe Temp        EPU pipe Temp     Pre-EPU ambient air       Pre-EPU Vent/Clg
                                          &deg;F (Ref. 7)              IF (Ref. 8)         Temp &deg;F                  Air AT OF Lv_ CNP                           133.1                   133.1              105.0                    20.0
   .RlN'\'Pump Room CND to RFP                                296.6          _-SLS                       105.0                    20.0 FW Ivg RFP                               299.3                  312.5                                        20.0 Main Steam ITunnel                                                                                   _
(~   FWIvg FWH1 Main Steam l
I   __
374.4 547.6
_  I    _
392.6 547.6
__       I _   ___
130.0 130.0 I   ___
25.0 25.0 Pump information:
Reactor Feedwater Pumps Table 1.4-3
__          Preuprate (Reference 7)               EPU (Reference 8)
Flow               64                           __        8048044 lb/hr I/of RFPs
* 2                                3        3 per pump           3203763 lb/hr                          26826________
Tin             __6_      . IF                                 309.8 OF_
Tout                  299.31&deg;F                                 312.5 DF Condensate Pumps Table 1.4-4
_                  Preuprate (Reference 7)               EPU (Reference 8)
Flow         -64d75261Ib/hr                               8076444 lb/hr f of CNPs                               3                            3 per pump          2145842 lb/hr                           2692148 lb/hr Tin                   133.8 &deg;F                                         OF Tout                  133.1 OF                                 133.1 Drywell information (Reference 17):
Drywell Cooling Load Summary*
Table 1.4-5 Compon           _                                                                         Btu/lhr Reactor Pressuress1                                                                          459,000 Recirc. Pumps, Valves, anP_                                                                  278,000 Feedwater Pipe & Valves                                                                      124,000 Steam Pipe & Valves                                                                          212,000 Condensate & Instrument Lead                                                                  82,000 Lines                                     ___

Calculation Number: VYC-2279           Revision Number: 0            CCN: 0       Page 13 of 28 Component                                                                     Btu/hl     l Control Rod Drive Pipe              ___                                        {0,400 Control Rod Drive Pipe                             569,000**
Clean-up Demineralizer Pipe &                                                     17,800 Valves                                                                      _
Shutdown Supply Pipe                                                               8,100 Steam Safety/Relief Valves                                                      206,600 Biological Shield (Gamma                                                          16,400 Heating)
Safeguards system Piping (RCIC,                                                   82,000 LPCI, HPCI. and core spray)            /
                                   'Sensible Heat Gain Total, Normal          1,536,300
                                    -Operation Steam Leal;, Valves     /155,000 Latent Heat Gain Total, Normal             155,000 Operation Total Cooling Load, Normal Operation     1,691,300
* Excluding allowance for drywell cooler motors
** Temporary initial load immediately following scram

Calculation Number: VYC-2279               e Revision Number: 0               CCN: 0           Page 14   of 28 1.4.2    Assumptions For the piping in the               el, it is assumed the feedwater piping contributes 1/3 to the total heat gain and the main steam piping contributes 2/3 due to their respective surface areas. This is based upon the fact there are four 18" main steam lines and two 16" feedwater lines. From Reference 14, the external surface area of four 18" pipes is 4
* 4.712 ft2 per foot of pipe 18.848 ft2 per foot of pipe and the external surface area of two 16" pipes is 2
* 4.189 ft2 per foot of pipe = 8.378 ft2 per foot of pipe. No confirmation is required.
It is assumed there is no appreciable change between the preuprate and EPU values for the thermal conductivity, k, of the pipe insulation based on the predicted temperature increases in the various process streams.
1.5       Calculation Piping:
A sample calculation for the estimated temperature rise from the feedwater piping in the steam tunnel is shown below.                                                                                 -
Preuprate pipe temp. (0 F)             374.4            Reference 7 EPU pipe temp. (0 F)                   392.6            Reference 8 Preuprate ambient air temp. (OF)        130             Reference 5 Preuprate Air HIandling Unit          25               Reference 4 (RRU-17A, 130 - 105 = 25)
(AHU) AT (0F)                                        I                                                        I The EPU ambient air temperature is initially estimated at 131 'F. [used for initial iteration and checked later]
The EPU Area Heat Gain Multiplier is obtained using the equation in Section 1.3.
EPU Area Heat Gain Multiplier = (EPU pipe temperature - EPU ambient air temperature)
(preuprate pipe temperature - preuprate ambient air temperature)
EPU Area Heat Gain Multiplier = (392.6 - 131)     IF (374.4 - 130) F EPU Area Heat Gain Multiplier = 1.070 The EPU factor is obtained by subtracting I from the EPU Area Heat Gain Multiplier:
EPU factor = 1.070 -1 = 0.070 From Assumptions section, the feedwater piping contributes 1/3 to the total heat gain in the main steam tunnel, so the EPU factor becomes:
EPU factor= 0.070 *1/3 EPU factor= 0.023

Calculation Number: VYC-2279              Revision Number: 0              CCN: 0         Page 15 of 28 The temperature rise is then calculated by multiplying the EPU factor by the AHU temperature differential.
Temp. rise = 0.023
* 25.0 0 F = 0.586 The difference between estimated EPU and preuprate ambient air temperatures is:
131 - 130 = I (this is not close enough to the calculated temperature rise of 0.586&deg;F)
Try a different EPU ambient air temperature of 130.6 and repeat the process.
Preuprate pipe temp. (F)              374.4            Reference 7 EPU pipe temp. (F)                    392.6            Reference 8 Preuprate ambient air temp. (F)        130             Reference 5 Preuprate AH-U AT (0 F)               25        l      Reference 4 (RRU-17A, 130 - 105 = 25)
The EPU ambient air temperature is estimated at 130.6 'F.
The EPU Area Heat Gain Multiplier is obtained using the equation in Section 1.3.
EPU Area Heat Gain Multiplier = (EPU pipe temperature - EPU ambient air temperature)
(preuprate pipe temperature - preuprate ambient air temperature)
EPU Area Heat Gain Multiplier = (392.6 - 130.6)
(374.4 - 130)
EPU Area Heat Gain Multiplier = 1.072 The EPU factor is obtained by subtracting I from the EPU Area Heat Gain Multiplier:
EPU factor = 1.072 -1 = 0.072 From Assumptions section, the feedwater piping contributes 1/3 to the total heat gain in the main steam tunnel, so the EPU factor becomes:
EPU factor= 0.072 *1/3 EPU factor = 0.024 The temperature rise is then calculated by multiplying the EPU factor by the AHU temperature differential.
Temp. rise = 0.024
* 25.00 F = 0.6 0 F The difference between estimated EPU and preuprate ambient air temperatures is:
130.6 - 130 = 0.6 (this equals the calculated temperature rise of 0.6 0F)
The temperature rises for the remainder of the pipelines was calculated using the same method using an Excel spreadsheet.
The results are shown in Table 1.5-1.

CALCULATION NlUMRER VYC.2279 Rev. 0 CCN O                                                                          Page 16of28 Table 1.5 Design Temperature Differential LA,   - -*                 C
* NA.        410 .. WJ14 wal          VN402)                   01-*             1443i                  444            14443 1                                            #4.440         0n44.W  WM3l .9 b.
10S04-,                   ma1          sin                    136       11        MO                  la              S. C-...                 0073                    Mc 0,1SWd                 IN133         N 0137                   120                 211                 107             S3.C-1                  '073                     210              406 1.721l                 270 l           344                   3210                1.164                4070m           S. C,-I                  0070                    300              44a 214.471.1                3744           M720                 1250                1214                 400G            S. CI-I.                 007                     200                44D                                            -  .
FVM,11            3431           21                   1210                121                 I WA             S. C,-m                  004                     M10              44                                     34103l.
: 4)                      4I:-IPPP.4T-S3     VA 71 Too1.3        3211                  SAO0                4%4                  &#xb6;21 W              .~                      0210                      a1             33                                         3I4..S.l 014                 220           221I                   4In0               131F                 421?            S. C.1                   0217                    210              37?
l.7.4Q32                 0721                   121                 4204                40WA            S..Ce-                   0074                    2100               44 31)?          34d                   lms0                 1204                4Iw0            S.                       0070O                   )D00               44 3,ro39,                           m0                     210                4212                I M)1           &. C-.                   0213                    200              413 30344340e               .321          10__312. _          4200               1211                 ton0            S1. 1..-i              0cam                   .10          -    4 .
v0d                             01 YA0      -              1200          *..                       I M407          S. cl-`1                 DO.            .        210 a              I40 N6 2270~G         204                   4200               1321                 IWO             ObO...3.                 OO                       10 220.O    .   '06.       ..          1200       -        421                  ion             S. CT-i..                000                     M1               44 7210*I M            M4I                   I420.         .4219            -         434             S. C-...                 DM?                     Mo.              I4 MO          204.
M$              .40 MD0. -               3m0             - -   WO     -.      $S.C...           .         W             --     210            .1
                                                 ~MI            Ml                  N                        y          S.C                        2                                      40       4.1          us00.1I me34213                  4720          leis                   Mo21               4          1           121
                                                                                                           '            S. C-               .0144                        M10              232 w460.,S                             Iwo04040       .320                       4INS                 I 3v            S. C*-                   0421.                   210       . .12 4.4,42              I1ll          4210e                  42s0               434465     N  em     4207          ,1011.6....                0207 114                210 144134 MP                  44           4446                - '200               11Th           hIa"                   S.C-                     0W0                      DO              I (417-37                   244T#                                 10                  21                   03               .Cco01                                         0 (4803(4.                   4313         4234                   Kee0                MO                  IWO             S. C-                    04                      210              a00      WA             400 10406404 .44r
!W444344                    W74 I .      37   .l     .     .      7     .           413D                40Iwo               0322                 0074                    20               00.                   INS WI     . W.      44    .            . 34       470      04        210           20G-
                                                       .OT-.-.                                                                  S60 3..70044444430434133(1044033.42.417--                0.                           y~O f   steed           to     __     _  __     _ -,
: 3. 4444         4433415440l344      P,400.3             641.4s-l,4         -      L~      1.63
                                                                                                                -,.-     440 m4.344.P.3c           .''*          -- ITP                                    - -.. , .- ..-. -. -        ...      .. .-

Calculation Number: VYC-2279              Revision Number: 0               CCN: 0         Page 20 of 28 The Total Cooling Load for Nor-mal EPU Operation is: 1,700,675 Btu/br The ratio of cooling loads to differential temperatures is:
q'pt / qp = ATCPU / ATPre ATepi, = (qcu
* ATpr,)/ qprc = (1,700,675 Bt       (135 - 97)0 F) / 1,691,300 BTU/hr AT~pu= 38.21 TF Since ATp,, = 135 - 97 = 38 The increase in dryw        perature due to the higher feedwater temperature is:
AT = ATP           38.2F- 380 F T     'mbient drywell temperature at EPU is 135 TF + 0.2 0F = 135.2 0F Main Steam Tunnel (RRU-17A, -17B )
q (Btufhr) = cfm
* 1.08 (Btu min/ft 3 hr 'F) * [Tin - Tout]0 F q = 5000 cfm
* 1.08 (Btu min/ft 3 hr&deg;F) * [130 - 105]&deg;F q = 1.35 x 105 Btu/hr The main steam tunnel piping is evaluated using the Excel spreadsheet described above. As stated in the assumption section, the temperature rise from the feedwater piping is "weighted" at 33% based upon its surface area ratio when compared to main steam piping. The increased feedwater temperature will result in an ambient temperature increase of approximately 0.60 F to 130.6&deg;F. As previously noted, the service water to RRU-1 7A&B is piped backwards, such that there is parallel flow rather than counter flow (Reference 18). Because of this, both coils are in continuous operation rather than the initial design of one in operation and the other as a back-up. The minimal increase in ambient temperature due to EPU will not adversely affect current operation due to the fact the current peak allowable temperature in the tunnel is 150&deg;F (Reference 6, pg. 10). Also, per Reference 6, Section 7.5.2.5, the high space temperature alarm set point is ] 60&deg;F and the MSIV close/scram set point is 200&deg;F.
Reactor Feedwater Pump Room (TRU-1.-2.-3.-4)
Theat removal capability of the reactor feedwater pump room is:
q (Btu/ = cfm
* 1.08 (Btu min/ft 3 hr &deg;F) * [Tin - Tout]&deg;F q = 4
* 167 ifm
* 1.08 (Btu min/Pt3 hr &deg;F) * [105 - 85]&deg;F q= 1.45x I06            r From above, the preuprate          oad due to two pumps operating is 0.97 x 106 Btu/hr and the EPU heat load due to three pumps operatini 1.47 x 106 Btu/hr.
Using the heat removal capability with a oolers running, the preuprate heat load due to the piping is:
1.45 x 106 Btu/hr - 0.97 x 106 Btu/hr= 0.50 x 106 Btu/
The temperature scaling method using the preuprate and EPU he             ances shows the ambient temperature will increase 1.2&deg;F.
The EPU increase in piping heat load due to higher feedwater temperature can be            imated using:
qpr = mCrATp,. and qppu = mCpATpu\
Since mCp is the same preuprate and EPU, qPU / qpZ = AT~pu / ATp,, or
: q. = qp% *(AT 1,. / ATpre)

Calculation Number: VYC-2279           Revision Number: CCN: 0                                  Page 26 of 28 T 1reas that contain condensate and feedwater piping arc the only areas that will experience an ambient tempere increase during normal operation due to EPU. The following systems will not experience an ambient tempeatture increase during normal operation due to EPU: Reactor Recirculation (RRS), Reactor Core Isolation C ng (RCIC), Residual Heat Removal (RH-R), Reactor Water Clean Up (RWCU), High Pressure Coolant Injec          PCI) and Core Spray (CS) (References 20 through 25).
Reactor Building open areas Temperatures in the open areas of the reac           ilding will not increase during normal operation as a result of EPU (Reference 26).
Control Room As shown in Calculation VYC-1502 (Reference 28), Section 2.1, heat sourl the control room are from electrical equipment, ambient outside air temperature, and personnel. None Uo ese sources will increase at EPU. Therefore, the Control Room HVAC system's ability to provide appropri             temperature and humidity conditions for personnel and equipment during all modes of operation and emergein condition is not impacted by EPU.
1.6       Results The results of this calculation are shown in Table 1.6-1 below.
Table 1.6-1 Area                   EPU Ambient Temperature Increase (OF)
Drywell                                                       0.2 Main Steam Tunnel                                            0.6 LP Heater Area                                                4.1 HP Heater Area                                                1.7 Feedwater Pump Room                                           7.6 Condensate Pump Room                                         3.5 Thencrasein  ainsteam tunnel ambient temperature due to the higher EPU feedwater temperature is The results of the piping evaluation are shown in Table 1.5-1. At normal operating EPU conditions, the am!.tbemperature in the LP heater area will increase approximately 4.10 F to 129.10 F and the HP heater area will in       e 1.70F to 126.7 0F.
The increase in feedWa        mp room ambient temperature due to the higher EPU feedwater temperature is 7.67F to 112.6&deg;F. This res            achieved using both design and empirical information.
The increase in condensate pump roomm              t temperature due to the higher EPU feedwater temperature is 3.5 'F. The ambient temperature in the conden            ump room at EPU based upon design information 0
is 113.2 'F and 122.5 F based upon empirical data.
It is noted that the temperatures obtained in this calculation are cons          emaximum temperatures for the purposes of obtaining bounding temperature increases within the subject ar           Actual EPU maximum temperatures are anticipated to be lower than those calculated.

BVY 05-072 Docket No. 50-271 Exhibit SPSB-C-52-6 Vermont Yankee Nuclear Power Station Proposed Technical Specification Change No. 263 - Supplement No. 30 Extended Power Uprate Response to Request for Additional Information Operation Procedure OP- 2192, Rev. 31 l Total number of pages In this Exhibit I (excludina this cover sheet) Is 5.

VERMONT YANKEE NUCLEAR POWER STATION 62 l            ll     OPERATING PROCEDURE OP 2192 REVISION 31 HEATING. VENTILATING, AND AIR CONDITIONING SYSTEM USE CLASSIFICATION: REFERENCE I                      RESPONSIBLE PROCEDURE OWNER: Manager, Operations REQUIRED REVIEWS                 Yes/No E-Plan                           IOCFR50.54( )  No I                        Security Probable Risk Analysis (PRA)
Reactivity Management 10CFR50.54(p)  No No No LPC    Effective No.       Date 1     10/26/04    App A Pg 15 of 24 2      11/16/04    51 & ADDED 51A of 62 3     01/22/05     60 of 62 4     05/26/05     29 of 62 IImplementation Statement: N/A Issue Date:    07/06/2004 OP 2192 Rev. 31 Page 1 of 62

TABLE OF CONTENTS PURPOSE3 D ISC U SSIO N .................................................................................................................................................... .................
3 ATTACHMENTS ..........                                                                                                                                               6...............
QA REQUIREMENTS CROSS REFERENCE................................................................................................6 REFERENCES AND COMMIT                             NTS ................................                                                                               7 PRECAUTIIONS/LIMITATIONS ................................                                                                                                            9 PREREQ  ISITES...........................................................................................................................................I PROCEDURE ..............................                                                                                                                             12 PROCE URE.......................................        ..... ........................................................................................ 1 Startup, Shutdown and Transfer of Service Building HVAC .12 Startup, Shutdown and Transfer of Turbine Building HVAC .15 Startup, Shutdown and Transfer of Reactor Building HVAC .25 Startup, Shutdown and Transfer of Radwaste Building HVAC .31 Startup and Shutdown of Intake Structure HVAC .33 Startup and Shutdown of Discharge Structure HVAC .....................................                                                                     34 Startup and Shutdown of Switchyard House HVAC .34 Startup and Shutdown of Steam Heat HVAC .37 Startup, Shutdown and Transfer of Advanced Off Gas Building HVAC .38 Operation of Control Room and Cable Vault Battery Room HVAC .............................................                                                  42 Operation of Chiller Units SCH-1 and 2 .44 Cross-Connecting SAC-1 and 2 Chilled Water.Systems .46 Startup and Shutdown of UPS-2A-AC1 HVAC .48 John Deere Diesel Generator Ventilation Controls .50 Startup and Shutdown of Turbine Deck Office HVAC .....................................................................                                     51 Startup and Shutdown of Switchgear Room HVAC .........................                                                                                     52 Loss of Control Room Ventilation (UND98080) .........................                                                                                      53 Local Manual Operation of EDG Room A(B) Exhaust Fan TEF-2(3) (Use VYOPF 2192.01) (ER990738_01) .........................                                                                                             53 Startup and Shutdown of Steam Heat to the CST .........................                                                                                    55 SCH-1-1 and SCH-1-2 Reset After Low Pressure Trip .........................                                                                                56 Temporary Space Heating .........................                                                                                                          57 FINAL CONDITIONS .........................                                                                                                                          62 OP 2192 Rev. 31 Page 2 of 62

The Turbine Lube Oil Room fire dampers are controlled by a local hand switch located outside the Lube Oil Room door.
The air is exhausted to the atmosphere by wall exhaust fans, roof exhaust fans or through the station stack. Several areas, such as the Control Room and office spaces recirculate air on a continuous cycle.
The system serving the Control Room is designed to provide cooling during the summer and heating during the winter. Air is circulated through a chilled water cooling coil, steam preheat coil, a steam reheat coil and ductwork by one of two system fans. Fresh air is normally drawn into the system mixing with the recirculated flow. A humidistat, on the west hallway wall in the first fan room, controls the relative humidity between 20% and 50% with a humidifier located in SRHC-1. It is operated with instrument air and controls humidity by spraying steam into the air flow. Two exhaust fans in the North wall of the Control Room, kitchen and rest room, serve to exhaust these rooms. The "Control Room H and V" switch on CRP 9-25 is provided to allow the operator to isolate the Control Room and Computer Room by closing the fresh air dampers and the Control Room kitchen and bathroom exhaust vents during off normal conditions. This is accomplished by moving the switch from "NORMAL" to "EMER". If a report was received of a toxic gas release which could affect Control Room personnel, the operators would don the self-contained breathing apparatus located in the Control Room. If Contol oolng ooms copleelylost emrgecy cooling can be initiated using portable smoke ejectors. (UND98OSO)
The Control Room and Service Building chilled water cooling coils are located in the SAC-1 and SAC-2 supply fan housings respectively. The cooling coils are cooled by dual compressor refrigeration units SCH-1 and 2 that cycle as necessary to maintain chilled water temperature. Demineralized water from the chilled water pumps circulates through the chiller heat exchanger and gives off its heat to the chiller units. The cooled water passes from the chiller units to the cooling coils. The amount of chilled water flowing into the cooling coils is controlled by the mixing valves. Each mixing valve is controlled by a thermostat that senses supply air temperature. The Services Building chilled water piping system can be valved into the Control Room HVAC because both the NNS piping system and the current SC3 piping system, respectively, were designed and built to the same specifications (i.e., non-seismic), therefore, failure mechanism(s) are the same for each system. In addition, isolation of the NNS and SC3 systems can be accomplished because the valves are in a mild environment area. Continued operation of the Control Room HVAC by valving in the Service Building HVAC to supply chilled water is consistent with the Safety Class Manual and the HVAC DBD.
Air for air compressor cooling is drawn through a wall louver located in the outside wall.
This louver also supplies any required room ventilation air. Discharge dampers which exhaust into the room are located on the air compressor discharge duct to allow for air recirculation.
Two oil-fired steam boilers supply steam for the heating coils and some of the unit heaters. Other unit heaters are run electrically.
All RRUs and TRUs utilize service water as the cooling medium except the drywell RRU-1 through RRU-4 which use RBCCW. (MOOID9502 14)
OP 2192 Rev. 31 Page 4 of 62
: 7.       For minimum ECCS RRU comer room support operation, RRU 7 must be operable for the Northeast comer room and RRU 8 must be operable for the Southeast comer room.
RRUs 7 and 8 can be removed from service for maintenance and the associated Core Spray Pump/RHR Pump may remain operable. Refer to OP 2181, Service Water/Alternate Cooling Operating Procedure, Precautions and Limitations, for administrative requirements and actions necessary to maintain operability.
: 8.       SP-1, SCH-1 and SAC-1 supply air conditioning for the Control Room. This is a Safety Class 3 system and requires special consideration for its timely repair.
: 9.      To prevent the possibility of reverse air flow, building ventilation should be secured in the following order:
: a.      Radwaste Building
: b.       Turbine Building
: c.       Reactor Building
: d.       Service Building
: 10. Minimize the time that TEF-2/3 are operated in manual. The UFSAR specified minimum Turbine Building design temperature, which includes the DG rooms, is 50 0F and the AS-2 battery load calculation assumes an electrolyte temperature of 50 0F.
(ER960055_01)
: 11. Securing TRU-5 with the condensate pumps running will result in a condensate pump bearing temperature rise. Planned maintenance on TRU-5 should be coordinated such that the time the unit is out of service is minimized.
: 1. Control Room temperatures in excess of 78&deg;F are indicative of a need for corrective action. Corrective actions need to be completed prior to exceeding 95&deg;F to ensure the Control Room does not reach 120&deg;F (upper temperature operability limit for Control        /
Roomistruentation). (UND98080)/
: 13.      HS-139 and HS-140 in the Reactor Building must remain locked closed during plant operation due to House Heating Steam, High Energy Line Break concerns. These valves may only be opened with the permission of the Design Engineering Department.
(ER96-0482, TAG_REVIEW_9703-26)
: 14. The Main Station Battery Room must be maintained at >600F. The main station battery load calculations are based upon this minimum temperature.
: 15. One of the two standby gas treatment trains should be placed in service whenever normal Reactor Building ventilation is secured.
: 16. In order to meet environmental qualification program requirements, the RCIC room fan or alternate ventilation from the Reactor Building must be operable and the RCIC room temperature must be less than 1120F. However, to satisfy station blackout analysis, the RCIC Room temperature must be maintained less than 104 0F. (EPCQ9504)
OP 2192 Rev. 31 Page 10 of 62
: 2. Shutdown
: a.      Secure the West S teaRoom           exhaust fan SWGR-EF-1A.
: b.      If     emperature cannot be maintained in the required range, notify the Operations Manager of the need to initiate actions to provide supplemental heg<     coling.
Q. Loss of Control Room Ventilation (UND98080)
: 1. If a loss of normal Control Room ventilation occurs, refer to Section J and place SAC-IB in service.,
: 2. If Control Room cooling is lost, perform applicable action:
: a.       Refer to Section L and cross-connect chilled water from the Service Building to the Control Room.
: b.       Perform the following:
: 1)      In the Control Room back panel area, remove 11 full size ceiling tiles.
: 2)      Open Control Room panel doors.
: 3)      Notify Security and Shift Engineer that Control Room doors will be opened.
: 4)       Open Control Room doors.
: 5)     Using two smoke ejectors or other portable cooling equipment to
     \                   ~create temporary air flow paths, ventilate Control Room./
: 6)      Imnplement the administrative requirements of AP 0077.
R. Local Manual Operation of EDG Room A(B) Exhaust Fan TEF-2(3) (Use VYOPF2192.01) (ER990738_01)
: 1. Ensure the selected EDG Room temperature is             [RATS-1A(B)].
: 2. Obtain Shift Manager permnissio         ocal manual operation of TEF-2(3).
: 3. Obtain key#D
: 4.       t ion a dedicated Auxiliary Operator at the selected Diesel Generator Room area.
OP 2192 Rev. 31 Page 53 of 62

BVY 05-072 Docket No. 50-271 Exhibit EMEB-B-138-1 Vermont Yankee Nuclear Power Station Proposed Technical Specification Change No. 263 - Supplement No. 30 Extended Power Uprate Response to Request for Additional Information Prediction of Unsteady Loading on A Circular Cylinder in High Reynolds Number Flows using Eddy Simulation.
Total number of pages in this Exhibit I (excludina this cover sheet) is 9. l

Proceedings of OMAE 2005:
24 th International Conference on Offshore Mechanics and Arctic Engineering June 12 - 16, 2005, Halkidiki, Greece OMAE 2005-67044 PREDICTION OF UNSTEADY LOADING ON A CIRCULAR CYLINDER IN HIGH REYNOLDS NUMBER FLOWS USING LARGE EDDY SIMULATION Sung-Eun Kim ---                                              L. Srinivasa Mohan Fluent Incorporated, Lebanon, N.H.                                            Fluent India 03766, U.S.A.                                                     Pune, India ABSTRACT                                                             of large-scale turbulent structure around a circular cylinder, which in turn affects unsteady loading on the cylinder.
Large eddy simulations were carried out for the flow            Apparently, turbulence modeling is an issue here.
around a hydrodynamically smooth, fixed circular cylinder at              There are largely three approaches being explored by CFD two Reynolds numbers, one at a subcritical Reynolds number           practitioners to modeling high-Re turbulent flows around (Re = 1.4 x 105) and the other at a supercritical Reynolds           circular cylinders and bluff bodies in general. One approach number (Re= 1.0 x 106). The computations were made using a           employs unsteady Reynolds-averaged Navier-Stokes (URANS) parallelized finite-volume Navier-Stokes solver based on a           equations supplemented by turbulence models as the governing multidimensional linear reconstruction scheme that allows use        equations.      URANS-based approach, despite its low of unstructured meshes. Central differencing was used for            computational cost mainly due to less demanding mesh discretization of both convection and diffusion terms. Time-         resolution requirement, is fundamentally ill equipped to capture advancement scheme, based on an implicit, non-iterative              large-scale turbulent structure present in the flows. Results fractional-step method, was adopted in conjunction with a            obtained using URANS computations typically underpredict three-level, backward second-order temporal discretization.          the amplitudes of fluctuating forces.           Obviously, this Subgrid-scale turbulent viscosity was modeled by a dynamic           shortcoming has a serious negative implication to accurate Smagorinsky model adapted to arbitrary unstructured meshes           prediction of VIV.
with the aid of a test-filter applicable to arbitrary unstructured        Fundamentally more viable than URANS-based approach meshes. The present LES results closely reproduced the flow          for bluff-body flows is large eddy simulation (LES). LES is features observed in experiments at both Reynolds numbers.           designed to directly resolve large eddies, with the effects of The time-averaged mean drag coefficient, root-mean-square            unresolved smaller eddies, namely subgrid-scale turbulence, on force coefficients and the frequency content of fluctuating          the resolved large eddies accounted for by turbulence models.
forces (vortex-shedding frequency) are predicted with a              However, LES is computationally very expensive, often commendable accuracy.                                                prohibitively, especially when thin turbulent boundary layer is a predominant feature of the flow in question to be accurately INTRODUCTION                                                         resolved in a LES. Resolving near-wall turbulence with infinitesimal length and time scales requires extremely fine Unsteady loading on a circular cylinder caused by the flow      mesh and small time-step size. For that matter, LES is more is a precursor to its vortex-induced vibration (VIV). The major      feasible for free flows such as jets, mixing layer, and wakes, difficulty of computing flows around circular cylinders              and massively separated flows.
originate from the fact that circular cylinder flows of practical         There are very few LES studies published in the literature interest are high Reynolds number (Re) flows, often involving a      of flows around circular cylinders at high Reynolds numbers.
laminar-to-turbulent transition in various regions of the flows      Breuer [I) was the one of the very few who tackled the case of such as boundary layer, separated shear-layer, or near-wake.         a high subcritical Reynolds number, Re = 1.4 x 105, at which The location of transition and the extent of laminar or turbulent    the experiment of Cantwell and Coles [2] was conducted.
flow regime, in real flows, depend on such factors as Reynolds       Using a multi-block, structured-mesh-based finite-volume number, freestream turbulence, surface roughness, span-              solver with an explicit time-marching scheme, Breuer was able diameter ratio (LID), and blockage, among others. The state of       to predict the global parameters of the flow, and the mean flow the flow in those regions dictates the formation and evolution       and turbulence with a commendable accuracy. More recently, I                                  Copyright &#xa9; 2005 by ASME

Catalano et al. [3] attempted LES for three Reynolds numbers      triangular,   tetrahedral, pyramidal,    prismatic,   and   hybrid in critical to supercritical Reynolds regime (Re = 0.5 x 106, 1.0 meshes. The solution gradients at cell centers that are needed x 106, 2.0 x 106).         Their predictions of the global flow   to compute convective and diffusive fluxes are determined parameters were in reasonable agreement with the experimental     applying the Green-Gauss theorem [7]. Diffusive fluxes are data, although the skin-friction and the Reynolds number-         discretized using central differencing.         Discretization of dependency of the mean drag coefficient were poorly predicted. convective fluxes requires great caution in LES.          Upwind-A more recent trend in turbulence modeling of bluff-body    biased schemes such as second-order upwind scheme have been flows is to employ what may be called "hybrid" approaches in      widely used for RANS computations. Unfortunately, numerical which one attempts to adjust turbulence models to local mesh      diffusion introduced by upwind schemes, which might be resolution in one way or another. In one such approach            acceptable in RANS computations for high-Re flows where referred to, in the literature, as detached eddy simulation       eddy-viscosity is larger than molecular viscosity by orders of (DES), either a RANS-based or a subgrid-scale (SGS)               magnitude, can overvhelm physical diffusion that is typically turbulence model is invoked depending on whether or not the       much smaller in LES; subgrid-scale turbulent viscosity is filter-length (local mesh size) is larger than the local integral smaller than RANS-based eddy viscosity by orders of length-scale of turbulence.         In DES, turbulence models     magnitude. For this reason, a central differencing scheme 113]
essentially reduce to RANS models in near-wall region or when     was employed for the discretization of convection terms in this the local mesh size is too coarse to explicitly resolve energy-   study.
containing eddies. One fundamental criticism about DES comes down to the lingering question of how one can possibly            TINME-ADV'ANCENMENT SCIIENMES reconcile a RANS turbulence model and a subgrid-scale turbulence model, two very different models in terms of their          An implicit fractional-step method (FSM) [8] in spectral content, at the common interface. It should also be      combination with a second-order. accurate, three-level realized that, as a consequence of falling back to a RANS         backward-differencing scheme for time-discretization was model in near-wall region, the fidelity of DES would be solely    employed to advance the solution in time. In this algorithm, the determined by that of the embedded RANS model. Quite a few        momentum equations are decoupled from the continuity studies employing DES have appeared recently on circular          equation using an approximate factorization of the coupled cylinder flows. Among others, Travin et al.[4], Vatsa and         Navier-Stokes equations. For incompressible flows, the FSM Singer[5], and more recently Pandya et al. [6] all employed       preserves the formal second-order temporal accuracy without DES based on an eddy-viscosity transport model to study the       having to perform, at each time-step, costly outer iterations to flow around a smooth circular cylinder at subcritical and         couple velocity and pressure. The FSM thus provides a highly supercritical Reynolds numbers. The results of these DES are      efficient algorithm for CPU-intensive transient computations mixed, insofar as some of the global flow parameters such as      like LES.
the r.m.s. force coefficients and the mean flow in the near-wake       Consider a semi-discrete form of the Navier-Stokes were predicted poorly.                                            equations in "pressure-correction" (+p"+ = p"+I - p") form, The objective of this study is to assess the feasibility of LES for high Reynolds number flows around a fixed, smooth circular cylinder. To that end, two Reynolds numbers, one in                                                                      (I) subcritical (Re = 1.4 x 105) and the other in supercritical (Re =                               ID 0J4,')0)
L Ox 106) regime were deliberately chosen so that they straddle the critical Reynolds number (- 3 x 105). The main concern is     where unt and r are the velocity vector and the momentum how well LES can reproduce the known characteristics of the       source vector, respectively, and the integer n the time level, flow at the two Reynolds numbers and the changes in the global    with n+1 refers to the current time level. A is the coefficient flow parameters such as drag coefficient, r.m.s. force            matrix defined in terms of the discrete convective and diffusive coefficients, and vortex-shedding frequency as the Reynolds       operators and the time-advancement scheme chosen, and G and number increases.                                                 D the discrete gradient and divergence operators, respectively.
The results of the LES will be compared whenever                 The coupled system of equations in Equation (I) is possible, with the experimental data and other LES and DES        extremely difficult to solve as it stands, since the matrix A has results reported in the literature.                               to be inverted for every iteration. In the FSM, the original coupled equations in Equation (I) are approximated by NUMERICAL METHOD The computations were carried out using the segregated solver in FLUENT, a general-purpose CFD software. The
[A iF'  0 ID AtDGJLIO AGl( u.,r"(r+[O(A, 2A "P~,.        0)IoLJ (2) details of the numerical method are described in the following.
Factorizing equation (2), we have a series of split operations SPATIAL DISCRETIZATION SCIIENMES                                  that consist of FLUENT employs a cell-centered finite-volume method                                          AO = -GP' + r based on a multidimensional linear reconstruction scheme that                                                                     (3)
AtDG~p/+' = DOi permits use of computational elements (cells) with arbitrary polyhedral topology, including quadrilateral, hexahedral,                                          U"   = 6 -AGJ" 2                                   Copyright &#xa9; 2005 by ASME

On a per-iteration basis, the series of operations in                    Ta,  I'l r,=-CVZ11                                        (5)
Equation (3) closely resemble the SIMIPLI.C scheme. The difference is that, in the iterative SIMPLEC scheme, the series of operations in Equation (3) are repeated in a loop until the              We determine the model constant, C,, using the dynamic solutions converge, while the FSM needs only one sweep. To            Smagorinsky model (DSM) originally proposed by Germano et preserve second-order accuracy with the FSMr, however, sub-           al. [12]. To implement the dynamic procedure for the present iterations are needed for the set of three momentum equations         finite-volume solver requires a special test-filter applicable to and individual scalar equations to account for the nonlinearity       arbitrary unstructured meshes. The dynamic procedure employed in and coupling among the individual equations and high-order         in the present study is "local" in the sense that it does not require source terms. Yet, the non-iterative FSM is takes much less           the existence of any statistically homogeneous directions, being CPU time than the iterative SIMPLEC scheme.                           applicable to arbitrary complex three-dimensional flows. The The system of discretized governing equations are solved         details of the implementation of the DSM in the framework of using a point-implicit, Gauss-Seidel relaxation along with an         unstructured mesh based finite-volume solver can be found in algebraic multigrid (AMG) method to accelerate solution               Kim [13]. The DSM has been validated for a number of wall-convergence. The solver and the subgrid-scale turbulence              bounded flows such as a square cylinder and a sphere.
model are fully parallelized.
As a validation for the spatial discretization schemes and the transient algorithm, we computed laminar flow around a            SOLUTION DOMAIN AND COMPUTATIONAL MESH circular cylinder at the Reynolds number of 100 with both non-iterative FSM and iterative SIMPLEC scheme.                At this          Our choices of the solution domain and the computational Reynolds number, the flow exhibits a periodic vortex-shedding.        mesh were guided by the findings in Breuer's LES study [1].
A C-type structured mesh with 48,000 cells was used for the           What one should keep in mind in determining the size of the solution domain is that the spanwise extent of the domain computation. A dimensionless time-step size of = 0.04 (non-a, dimensionalized by D/UI D being the cylinder diameter, U0 the         should be larger than the spanwise correlation length of freestream velocity) was used, with which one period of the           turbulence. Fortunately, the spanwise correlation length is vortex-shedding was resolved with approximately 150 time-             known to decrease as the Reynolds number increases. We took advantage of this fact, having decided to use a spanwise length steps. The predicted mean drag coefficient (CD), r.m.s. lift of 2.OD for both the subcritical and the critical Re numbers.
coefficient (c), and Strouhal number (St) are summarized in           The top and bottom boundaries of the domain were placed at Table 1. The FSM gives practically the same predictions as the         10.5D from the cylinder axis. Thus, the blockage ratio of our iterative SIMPLEC scheme, and the global parameters                   numerical model (HID, where H is the height of the domain) is predicted by both methods are seen to closely match the other         approximately 4.8 %. The upstream inlet and the downstream predictions and the experimental data.                                exit boundaries are located at 8.5D upstream and 20.5D downstream of the cylinder axis, respectively.
A locally refined, hexahedral mesh with a total of 6.8 Table 1. Summary of the global flow parameters predicted              million cells was used for the computations for both Reynolds for laminar flow (Re = 100) - Norberg [101 compiled                   numbers. The overall mesh resolution is comparable to case numerical results whose references can be found in ref. 1101          "Bl" in Breuer 111. The near-wall mesh resolution is such that the distance from the cylinder surface is 104D at the wall-
                                     .CD      CL          St          adjacent cells, which translates to y4(=u,)/I) well under 1.0 for Present - FSM                   1.336        0.28       0.165         the subcritical flow, and below 6.0 for the supercritical flow.
Present - SIMPLEC               1.345        0.28       0.166         For the subcritical flow, the boundary layer was resolved with 10- 18 cells in the laminar region.
Park etal. [9]                   1.33        0.33       0.165 Norberg [10]                       -      0.23/0.29        -          BOUNDARY CONDITIONS AND OTHER DETAILS OF WVilliamson [Il]                   -           -        0.164         COMPUTATION On the cylinder wall, we employ a law-of-the-wall that SUBGRID-SCALE TURBULENCE MODELING                                     invokes proper wall-laws depending on the mesh resolution, namely, y 4 at the wall adjacent cells. The top and bottom wall For incompressible flows, the filtered Navier-Stokes          boundaries were treated as a slip (zero-stress) wall. A equations can be written as                                           periodicity condition was applied on the pair of lateral boundaries in the span-wise direction. At the upstream inlet boundary, the freestream condition was specified. At the ai7, +ai =,-l ap -Irl              a Vaui               (4)'    downstream exit boundary, we extrapolated the solution at    ax,       p ax,    ax, axet axJ                          variables from the adjacent interior cells in a mass-conserving where r,, = uju, -      J is the subgrid-scale turbulent stress. In manner.
The time-step size (4t) used for the computations is 0.005 this study, the subgrid-scale (SGS) stress is modeled using           in a dimensionless unit for both Reynolds numbers. The time-isotropic eddy-viscosity as                                           step size was determined based on the estimate of the 3                                   Copyright 0 2005 by ASME

characteristic length and time scales ot the smallest resolved eddies, -r=l/d', where ( was taken as 0.05D, and d as               Table 2. Summary of the global nowr parameters predicted by 0.2U0. From these rough estimates and with the dimensionless        the present LES for Re = 1.4 x 10 , compared Xith other time-step size of 0.005, one turnover time of the smallest          numerical reults (Breuer [1]; )lES-TS - Travin et a/.14]; DES-LS resolved eddies will be resolved with 50 time steps.                - Travln et al.[4]; and experimental data (CC - Cantivell &
Coles[2], WA - WVest & Apelt 1141, SB: Szepessy & Bearnian [15],
To shorten the initial transient period of the solution and to NO - Norberg 110], ZD: Zdravkovich [161) quickly attain a statistically stationary state, we took, as the initial condition, a partially converged steady RANS solution with pseudo-random fluctuating velocity components                                       CD        C,         C c        rL,      St superimposed on the mean velocity field.                              LES (Present)     1.13     0.59         1.20      0.67    0.205 The computations were carried out on a 16-CPU Inx86              LES (Breuer)      1.45        -         1.76      0.34    0.204 cluster with AMD Opteron processors and Infiniband                    DES-TS           0.59      0.06         0.67       1.2    0.31 interconnects. With the non-iterative FSM, the computation            DES-LS            1.08     0.29         1.04       1.1    0.21 took approximately 12 CPU-seconds per a time-step, which              Exp.(CC)          1.24        -         1.21      0.5     0.18 translates into being able to complete roughly 7,200 time-steps       Exp.(WA)           1.3     0.58           -         -       -
in 24 hours.
Exp.(SB)          1.35      0.5         1.05-s Exp.(NO)            -      0.52           -            0.18/0.195 RESULTS                                                               Exp. (ZD)         1.15    0.5/0.6     1.05/1.2         0.18/0.21 Subcritical Reynolds number (Re = 1.4 x10 5 )
Figure I depicts the flow structure around the cylinder.            The present LES gave the mean CD around 1.13, which Although the near-wake flow appears chaotic, the figure clearly     falls within the scatter of the experimental data (1.1 - 1.35) for shows the existence of a large-scale motion, namely alternate       smooth cylinders. Interestingly, our prediction came out vortex-shedding known to occur in the subcritical Re number         substantially lower than Breuer's prediction despite the range. The color map of the resolved turbulent kinetic energy       comparable mesh resolution, the same formal order of spatial in Figure I also indicates that the boundary layer remains          accuracy (second-order) and the same dynamic SGS turbulence laminar up to the separation point, and transition occurs in the    model used in both computations.
separated shear layer.
4 d
Figure 1. The flow structure behind the circular cylinder at                                   400      450        500 Re = 1.4 x 105 - iso-surface of the second invariant of                                   Non-dimensional lime, 1*(= tUJD) velocity deformation tensor, colored by the resolved turbulent kinetic energy                                            Figure 2. Time histories of the drag and lift coefficients predicted by the present LES for Re = 1.4 X 105 The global flow parameters predicted by the present LES are summarized in Table 2 along with the LES prediction of               The DES-TS result [4] is shown to severely underpredict Breuer [1] and the DES predictions of Travin et al. [4], and the    the mean CD, better matching the data obtained with rough experimental data. As for Breuer's LES result, among the            cylinders [17J, which is consistent with the DES predictions by cases involving different mesh resolution and SGS turbulence        others [5, 6]. It has been found that DES, run in a normal models, we picked the result of case "B`1," whose mesh and          mode with finite freestream turbulence, essentially yields a SGS turbulence model (dynamic Smagorinsky mode) closely             tripped boundary layer, leading to a premature transition to match ours.        In selecting the experimental data for the       turbulence and an early drag crisis even at a subcritical comparison, we gave more weight to the ones measured on             Reynolds number. The DES-LS result [4] based on the so-smooth cylinder surface, and with sufficiently large span-wise      called "tripless" approach, in which laminar flow is essentially lengths (LID > 5) and smaller blockage-ratio (HID < 0.1).           enforced in the boundary layer, better predicts the mean CD.
4                                    Copyright &#xa9; 2005 by ASME

10 10 I0 l.0 1-.o 0   2.0    3.0    4.0    5.0    6.0     7.0  8.0
       &deg; .0   0.2   0.4  0.6   0.8     1.0* 1.2   1.4  1.6   1.8 2.0                                    x/D St Figure 3. Power spectral density          Of CL time-history for Re  Figure 5. Time-averaged axial velocity profile in the wake
= 1.4 x 10o                                                          predicted by the present LEIS for Re = 1.4 X 105 The r.m.s. lift coefficient predicted by the present LES is         Figure 5 depicts the predicted time-averaged axial velocity within the scatter of the experimental data, although it is closer   profile along the wake centerline at the mid-span. The overall to the upper bound. The CL history Figure 2 clearly shows            agreement between the prediction and the measurement is quite that, despite modulation of the amplitude, there is a distinct       good. The rapid relaxation of the mean axial velocity in x/D =
frequency of vortex-shedding, which is consistent with what          1.0 - 3.0 shown in the measurement is an indication of has been known for smooth cylinders in the subcritical regime.       vigorous mixing of momentum taking place in the near-wake The Strouhal number corresponding to the main shedding               region. That our prediction closely captures the recovery of the frequency was found to be 0.205 as shown in Figure 3.                mean axial velocity verifies that the dynamics of the large-scale structure in the near-wake is well predicted in the present LES.
2.0 The length of the time-averaged recirculation bubble (see also Table 2 for Lr) from our LES was found to be slightly larger than what Cantwell and Coles [2] reported. Yet the present
                           -Present                                  LES prediction is much closer to the data than the DES
: 1.                    a Exp. (Cantwell &Coles,1983)
* Exp. (Sarioglu & Yavuz, 2002)         predictions. Considering that hot-wire measurements become increasingly difficult and less reliable near the recirculation region, our LES prediction is considered quite good.
   ;-  0.0  -
5.0 -                                   Lower surface
      -2.
0       30        60         90       120      150     180 0 (deg.)
Figure 4. Mean static pressure (Cp) distribution on the                    0.0 cylinder surface for Re = 1.4 x 105 The time-averaged pressure distribution on the cylinder
                                                                          -5.0 surface is shown in Figure 4, along with the experimental data
[2, 18] obtained at the same Reynolds number. The negative peak predicted by the present LES is larger than the measured                  0       30      60        90       120      150     180 0 (deg.)
ones. However, the pressure level in the separated region - the plateau after 0 = 90 degrees, and the mean base pressure (UPB)       Figure 6. Scaled instantaneous skin-friction coefficient are all accurately predicted by the present LES.                     distribution on the cylinder surface at Re = 1.4 x 105 5s                                  Copyright &#xa9; 2005 by ASME

                                                                                                                 . .  - . ^U w  -
peaks at 3.0 around 0= 50 degrees, which is closely reproduced by the present LES.
Supercritical Reynolds number Case (Re = 1.0 x10 6 )
0 (deg.)
Figure 7. Resolved turbulent kinetic energy distribution at 0.0001D above the cylinder surface, predicted the present Figure 8. The flow structure behind the circular cylinder at LES at Re = 1.4 x 105                                                  Re = 1.0 X 106 - an iso-surface of the second invariant of velocity deformation tensor, colored by the resolved As mentioned earlier, the experimental evidence [19]              turbulent kinetic energy indicates that, in the subcritical regime, the boundary layer remains laminar until it finally separates from the cylinder surface. Therefore, it would be of much interest to see how                 Figure 8 portrays an impression of the instantaneous flow structure. Compared to the flow structure for the subcritical closely LES can reproduce the physics associated with the separation and the transition. Figures 6 and 7 depict the              flow shown in Figure 1, the wake is much narrower and more chaotic without any trace of a large-scale alternate vortex circumferential variations of two quantities along the cylinder surface that shed some light on how the LES predicts the               shedding, suggesting a delayed turbulent separation. This is typical of the flow in the supercritical regime. The color map separation and transition.
on the iso-suraface - representing resolved turbulent kinetic Figure 6 shows a plot of the instantaneous skin-friction energy - also indicates that the laminar boundary layer is coefficient - rescaled to match the non-dimensionalization sustained farther downstream than the subcritical flow.
adopted by Achenbach [19] - against the circumferential angle.
First, it is noticeable that the skin-friction distribution exhibits a Table 3. Summary of the global quantities predicted by LES for considerable top-bottom asymmetry. This is a clearly the               Re = 1.0 x 10', compared with other LES result and the (upstream) influence of the large-scale, alternate vortex-             experimental data; CA - Catalano et al.[3]; SZ - Szechenyl.120];
shedding. What is most interesting in the figure is the                SHI - Shih et al. [211; ZD - Zdravkovich [16])
appearance of small transient separation bubbles on both upper and lower surface starting as early as 70 - 75 degrees.                                          CD   I    C;.L      I        CpE     St Interestingly, this range largely coincides with the range of            LES (present)          0.27      0.12               0.28      -
separation angles reported by many others [2, 19] deduced from           LES (CA)               0.31        -                0.32    0.35 the inflection point of mean Cp curve; 77 degree by Cantwell             Exp.(SZ)               0.25                                   -
and Coles [2], and 78 degrees by Achenbach [19] at Re = 1.0 x            Exp.(SH)               0.24        -                0.33      -
105 . Although not shown here, the separation angle determined           Exp. (ZD)            .0.2/04   0.1/0.15          0.2/0.34 0.18/0.5 based on the time-averaged velocity field obtained from the present LES was found to be much larger than the ones determined from Cp distribution, reaching around 98 degree,                 The global flow parameters predicted by the present LES which was also found by Breuer [1].                   Based on the     are summarized in Table 3. Our prediction of the mean drag instantaneous skin-friction distribution, we surmise that a            coefficient (CD) closely matches the data of Szechenyi [20] and transient laminar separation occurs much earlier than the mean         Shih et al. [21] interpolated at the present Reynolds number.
separation angle.                                                      We consider their data more reliable than others in view of the The distribution of the resolved turbulent kinetic energy        sufficiently large spanwise lengths used (LID = 9.3 and 8.0, depicted in Figure 7 supports the possibility of an intermittent       respectively), relatively low blockage ratio (7% and 11%,
laminar separation around 70 - 75 degrees, insofar as the              respectively), surface smoothness, and relatively low turbulent kinetic energy is still quite low there. The complete        turbulence intensity of the wind tunnels - 0.08 % for Shih et transition occurs a little'downstream as indicated by the rapid        al.'s data [21]. The r.m.s. lift coefficient prediction by the increase in the resolved turbulent kinetic energy between 75           present LES falls within the scatter of the experimental data and 85 degrees. The magnitude of the skin-friction coefficient         compiled by Zdravkovich [16]. The base pressure is also seen predicted by the present LES is also in good agreement with            to be in the range of the experimental data.
Achenbach's data [19] measured at Re = 1.0 x 105.
Achenbach's data show that the scaled skin-friction coefficient 6                                       Copyright 0 2005 by ASME

As before, we plotted the circumferential variations of the instantaneous skin-friction and the resolved turbulent kinetic energy in Figures II and 12. The top-bottom asymmetry of the skin-friction distribution is seen to be much less pronounced than for the subcritical flow. Figure 12 suggests that the boundary layer sustains being laminar down to 90 degrees. The separation angle of the time-averaged velocity field from the present LES was found to be around 108 degrees, which is close to 106 degrees estimated by Shih et al.[21] using their U~~'
1"'1'jI                                                   experimental data on a smooth cylinder at approximately the 4            40        *'*'
VI              4
       -0      I ._. -      .           ...   ;         ..       ,I       same Reynolds number. All this taken together, our LES result suggests that the boundary layer undergoes transition starting as
       -0.2-                                                              early as 90 degrees, before it finally separates - "turbulently" -
at around 108 degrees.
              -04 420             440             460             480 Non-dimensional time, l* (= tU 0/D)
Figure 9. Time histories of drag and lift coefficients for the circular cylinder at Re = 1.0 x 106                                                                                           surface       -Upper 5.0 -                                  Lowersurfacel The time histories of drag and lift coefficients are shown in Figure 9. Consistent with the flow structure portrayed in Figure             0.
8, the CL-history at this supercritical Reynolds number is more random than that for the subcritical flow. Indeed, no single frequency of vortex-shedding could be identified at this Reynolds number, as shown in the plot of the power spectral                     -50 density of CL in Figure 10. This finding is consistent with the others' experimental observations, for instance, Shih et al. [21 ]
who found that coherent vortex shedding disappeared on a smooth cylinder beyond Re = 4 x 105.
* 0         30       60        90      120      150      180 0 (deg.)
          -02 Figure 11. Scaled instantaneous skin-friction coefficient distribution on the cylinder surface at Re =1.0 x 106 10'           t U
o     6 10 I,t 10 0          0.2  0.4   0.6   0.8     1    1.2   1.4   1.6    1.8 2 Strouhal Number Figure 10. Power spectral density of the lift coefficient (CL) for the critical Reynolds number (Re = 1.0 x 106) 0 (deg.)
For a smooth cylinder, the experimental data collected by          Figure 12. Resolved turbulent kinetic energy distribution at Zdravkovich [16] and Achenbach's data [19] indicate that,                 0.OOO1D above the cylinder surface, predicted the present above the Reynolds number of 1.5 x 106 referred to as "TrBL4"             LES at Re= 1.0 x 10O in [16], transition clearly precedes the boundary layer separation. The exact location of the transition depends on such factors as freestream turbulence level and surface                         The skin-friction prediction, including the location of the roughness. At Re = 1.0 x 106 falling in the range referred to as          peak and the magnitude, is also in good agreement with the "TrBL3", however, both sets of data seem to suggest that                  experimental data of Achenbach [19] measured at Re = 8.5 x transition occurs near the separation.                                     10 and Re = 3.6 x 106. At these two Reynolds numbers, 7                                    Copyright C 2005 by ASME

Achenbach's data show that the scaled skin-friction                 University in Turkey for kindly allowing me to use his distributions have the maximum values in the range of 3.5 - 4.0     experimental data.
occurring around 0= 60 - 65 degrees. As shown in Figure I1, our LES prediction closely matches Achenbach's data. This is        REFERENCE in a sharp contrast with the LES prediction by Catalano et al.
[3], which significantly overpredicted the overall level of the     [I] Breuer, M., 2000, "A Challenging Test Case for Large skin-friction on the front half of the cylinder. Contrary to what        Eddy Simulation: High Reynolds Number Circular is suggested by the experimental evidence, their LES seem to             Cylinder Flow," Int'l J. Heat and FluidFlow, 21, pp. 648 -
yield too early a transition, giving a skin-friction level typical       654.
of a turbulent boundary layer from immediately downstream of        [2] Cantwell, B. and Coles, D., 1983, "An Experimental Study the front stagnation point. As they pointed out [3], one most            of Entrainment and Transport in the Turbulent Near-Wake likely cause for the discrepancy is the substantially coarser            of a Circular Cylinder," J. Fluid Mlech., 136, pp. 321 - 374.
mesh (yV around 50) used in their LES predictions.                  [3] Catalano, P., WVang, M., Iccarino, G. and Moin, P., 2003, "Numerical Simulation of the Flow Around a Circular

==SUMMARY==
AND CONCLUSIONS                                                  Cylinder at High Reynolds Numbers," Int'l J. Heat and Fluid Flow, 24, pp. 463 - 469.
Turbulent flow around a smooth fixed circular cylinder         [4] Travin, A., Shur, M., Strelets. M. and Spalart, P., 1999, was computed using large eddy simulation (LES) at two                    "Detached-Eddy Simulation Past a Circular Cylinder,"
Reynolds numbers (Re = 1.4 x 105, 1.0 x 106).                  The       Flow, Turbulence and Combustion, 63, pp. 293 - 313.
computations were carried out using a second-order accurate,        [5] Vatsa, V. N. and Singer, B. A., 2003, "Evaluation of a parallelized finite-volume Navier-Stokes solver capable of               Second-Order Accurate Navier-Stokes Code for Detached handling arbitrary unstructured meshes. An implicit, non-                Eddy Simulation Past a Cricular Cylinder," AIAA Paper iterative fractional-step method was employed in favor of its            2003-4085.
high efficiency in LES for incompressible flows.                    [6] Pandya, M. J., Frink, N. T., Abdol-Hamid, K. S., and The LES results for both Reynolds numbers predicted,                Chung, J. J., 2004, "Recent Enhancements to USM3D with a good accuracy, the global flow parameters such as mean            Unstructured Flow Solver for Unsteady Flows," AIAA drag coefficient, r.m.s. lift coefficient, and the Strouhal number.      Paper 2004-5201.
Furthermore, the salient features of the subcritical and the        [7] Kim, S. E., Makarov, B., and Caraeni, D., 2004, "Multi-supercritical flows experimentally observed and measured on a            Dimensional Reconstruction Scheme for Unstructured smooth cylinder, such as the separation angle, mean velocity in          Meshes," AIAA Paper 2004-2548.
the wake, length of the recirculation bubble, and transition        [8] Kim, S. E. and Makarov, B., 2005, "An Implicit location, are largely reproduced by the LES. Despite the use of          Fractional-Step Method for Efficient Transient Simulation a large number of computational elements (6.8 million cells),            of Incompressible Flows," To be presented at 17Th AIAA the solution turnaround time was quite reasonable.                       Computational Fluid Dynamics Conference, June 6 - 9, Toronto, Ontario.
To summarize what has been achieved through this study:        [9] Park, J., Kwon, K., and Choi, H., 1998, "Numerical Solutions of Flow Past a Circular Cylinder at Reynolds
* The spatial discretization method (finite-volume on              Numbers up to 160," KSAIE nt. J., 12, pp. 1200 - 1205.
unstructured meshes) and the solution algorithm            [10]Norberg, C., 2003, "Fluctuating Lift on a Circular (implicit fractional-step method) employed in this study        Cylinder: Review and New Measurements," J. Fluids and have been shown to be sufficiently robust and accurate          Structures, 17, pp. 57 - 96.
to be used in LES for high Reynolds number flows.          [I I]Williamson, C.H.K., 1989, "Oblique and Parallel Modes of
* The dynamic Smagorinsky model adapted to                         Vortex Shedding in the Wake of a Circular Cylinder at unstructured meshes using a novel test-filter [13] has          Low Reynolds Numbers," J. Fluid Mech., 206, pp. 579 -
been shown to work robustly and accurately for                  627.
complex, high Reynolds number turbulent flows.             [12]Germano, M/, Piomelli, U., Moin, P., and Cabot, W. H.,
1991, "Dynamic Subgrid Scale Eddy Viscosity Model,"
* The present LES capability has been shown to predict the salient features of turbulent flow around a smooth          Physics of Fluids A, 3, 19, pp.1760- 1765.
circular cylinder at both subcritical and supercritical    [13]Kim, S. E., 2004, "Large Eddy Simulation Using regimes with a good accuracy and reasonable                     Unstructured Mesh and Dynamic                  Subgrid-Scale Turbulence Models," AIAA Paper 2004-2548.
computational cost.
[14]West, G. S. and Apelt, C. -J., 1993, "Measurements of fluctuating Pressures and Forces on a Circular Cylinder in It is planned to continue this work to include higher the Reynolds Number Range 104 to 2.5 x 105," J. Fluids Reynolds numbers.
and Structures, 7, pp.227 - 244.
[15] Szepessy, S. and Bearman, P. W., 1992, "Aspect Ratio and End Plate Effects on Vortex Shedding from a Circular ACKNOWLEDGEMENT                                                          Cylinder, J. FluidMech., 234, pp. 191 - 217.
[16]Zdravkovich, M. M., 1997, "Flow Around Circular The authors acknowledge that FLUENT and GAMBIT                      Cylinders, Fundamentals, Vol. 1, Oxford Univ. Press software were used for this study. The authors are also (Chapter 6).
thankful to Dr. Mustafa Sarioglu at Karadeniz Technical 8                                   Copyright &#xa9; 2005 by ASME

117]Roshko, A., 1961, "Experiments on the Flow Past a Circular Cylinder at a Very High Reynolds Number," J.
Fluid Alech., 10, No. 3, pp. 345 - 356.
[18]Sarioglu, M. and Yavuz, T., 2002, "Subcritical Flow Around Bluff Bodies," AMAA J. Vol. 40, No. 7, July.
[19]Achenbach, E., 1968, "Distribution of Local Pressure and Skin-Friction Around a Circular Cylinder in Cross-Flow up to Re = 5 x 106,,, J. Fluid Mech., 34, Part4, pp. 625 - 639.
[20]Szechenyi, E., 1975, "Supercritical Reynolds Number Simulation for Two-Dimensional I-low Over Circular Cylinders," J. Fluid. Mech., 70, pp. 529 - 542.
[21IShih, W. C. L., Wang, C., Coles, D., and Roshko, A., 1993, "Experiments on Flow Past Rough Circular Cylinders at Large Reynolds Numbers," J. Wind Eng. And Indus.
Aerodynamics, 49, pp. 351 - 368.
9- Copyright &#xa9; 2005 by ASME

BVY 05-072 Docket No. 50-271 Exhibit EMEB-B-138-2 Vermont Yankee Nuclear Power Station Proposed Technical Specification Change No. 263 - Supplement No. 30 Extended Power Uprate Response to Request for Additional Information Large Eddy Simulation of Confined Swirling Coaxial Jets.
Total number of pages in this Exhibit I (excludina this cover sheet) is 10. l

                                                                                                 . D  .. .D +-   +  D Proceedings of ASME-FED:
2nd  Symposium on Measurements and Modeling of Large-Scale Turbulent Structures June 19 - 23, 2005 FEDSM 2005-77085 LARGE EDDY SIMULATION OF CONFINED SWIRLING COAXIAL JETS Sung-Eun Kim, Fluent Inc.,              Xuelei Zhu, Fluent Inc., Lebanon,            Stefano Orsino, Fluent Inc.,
Lebanon, NH 03766, U.S.A.                          NH 03766, U.S.A.                   Lebanon, NH 03766, U.S.A.
ABSTRACT                                                           concentration were much less satisfactory.            The major discrepancy was found in the core region, where the mean Large eddy simulations (LES) have been carried out for        species concentration was grossly underpredicted by the RANS confined swirling coaxial jets discharged into a suddenly          computations. Evidently, the RANS model employed gives expanded pipe, which was studied experimentally by Roback          rise to an excessive mixing. Brankovic et al. [31 investigated and Johnson [1, 21. The computations were made using a             the sensitivity of their RANS predictions to such parameters as parallelized finite-volume-based Navier-Stokes solver that is      the turbulent Schmidt number and the inlet turbulence level.
second-order accurate in time and space, and permits use of        However, the influence of these parameters was negligibly unstructured meshes. The computational domain starts from an       small. The poor prediction of the species mixing has been inlet placed upstream of the swirl generator, which makes the      attributed to the inability of RANS-based turbulence models to inlet boundary condition easy to specify.          Subgrid-scale   accurately represent the mixing by large-scale turbulent turbulent stresses were modeled using a dynamic Smagorinsky        strmctuirc (or large eddies).
model applicable to complex three-dimensional flows without              Playing a major role in mixing, large eddies are also any statistically homogeneous directions.          Subgrid-scale   responsible for undesirable yet somewhat subtler phenomena turbulent scalar flux is modeled using a constant Schmidt          such as combustion instability and acoustic excitation inside number in conjunction with the dynamically computed subgrid-       combustors. To address these issues properly, one has to turn scale turbulent viscosity. The LES predictions were found to       to high-level turbulent simulation like LES. Akselvoll and closely reproduce the salient features of the flow and the         Moin [5] used LES to compute the flow and the species species concentration downstream of the expansion. One of the      concentration in non-swirling coaxialjets. Pierce and Moin [6]
conclusions was that a good resolution of the mean flow and        attempted LES for the same confined coaxial swirling jet as is turbulence in the upstream region is crucial in accurately         considered here. These studies demonstrated that LES can predicting the mixing downstream of the expansion.                 indeed predict the flow and the species concentration in confined coaxial jets with a commendable accuracy. LES also allows one to directly compute r.m.s. velocity fluctuations and INTRODUCTION                                                       r.m.s. species fluctuation which are important quantities in the context of turbulent combustion.
Confined swirling co-axial jets discharging into a suddenly        The present study concerns a LES computation for the expanded pipe [1, 2] have been studied numerically by several      swirling coaxial jets using a finite-volume solver that permits others using the Reynolds-Averaged Navier-Stokes (RANS)            use of unstructured meshes.      Unlike the study of Pierce and equations [3, 4] and large eddy simulation [6]. The flow in        Moin [6], however, our computational domain starts from an question consists of a non-swirling jet in the center, and an      inlet boundary placed upstream of the swirl generator. Despite outer annual jet with a swirl imparted by a 8-vane, variable-      the larger solution domain and the implied increase in the angle swirl generator, with a swirl number of approximately        computational resource, having the inlet boundary upstream of 0.41.                                                              the swirl generator makes it straightforward to specify the inlet Earlier numerical studies using RANS-based turbulence         boundary conditions. Meshing the computational domain models [3, 4] have shown that the mean velocity field in the       including the swirl generator, which could become a difficult mixing region .- the shear layer between the jets and the inner    and time-consuming task with structured meshes, is made recirculation region - was predicted with a reasonably good        easier by the flexibility offered by unstructured mesh allowed accuracy. However, the predictions of the mean species             by the present finite-volume solver.
I                                  Copyright C)2005 by ASME

This paper is, in many aspects, a pi( ..         epoIl aboutl an ongoing study whose ultimate goal is to finbl awnoplimal                                                       Aii=-Gp" + r' strategy based on the technique of LES fur modeling turbulent (3) flow and species mixing in coaxial conmbustors with or without swirl.                                                                                                      fi   = a, - AYG,"-
The paper is organized as follows. F:irst, we briefly discuss the numerical method and the subgrid-scale turbulence models                On a per-iteration basis, the series of operations in for stresses and scalar flux. Special emphasis is laid on a            Equation (3) closely resembles SIMPLEC schiemc.                 The transient algorithm whose efficiency significantly benefits the        difference is that, in the iterative SIMNIP.LEC scheme, the present study. This is followed by the details of the LES              operations in Equation (3) are repeated in an o0.teI loop until computations regarding the choices of the solution domain,             the all the solution variables converge, whereas the FSMI needs mesh (resolution), boundary conditions, timne-step size, and the       only one sweep. To preserve second-order accuracy) with the overall solution strategy. The results are then presented.             FSM, however, sub-iterations are needed for the set of three momentum equations and individual scalar eqtuations lo account for the nonlinearity in and coupling among the NUMERICAL METHOD                                                       individual equations and high-order source terms. Yet, non-iterative FSM is takes much less CPU time than iterative The computation were carried out using the segregated             SIMPLEC scheme.
solver in FLUENT, a general-purpose CFD software.                           The system of discretized governing equations are solved FLUENT employs a cell-centered finite-volume method based              using a point-implicit, Gauss-Seidel relaxation along with on a multi-dimensional linear reconstruction schemne that              algebraic multi-grid (AMG) method to accelerate solution permits use of computational elements (cells) with arbitrary           convergence. The N-S solver and the SGS turbulence model polyhedral cell topology including quadrilateral, hexahedral,          are fully parallelized.
triangular, tetrahedral, pyramidal, prismatic, and hybrid meshes. The solution gradients at cell centers that aIe needed to compute convective and diffusive fluxes are obtained by             SUBGRID-SCALE TURBULENCE MODELING applying Green-Gauss theorem [9]. Diffusive and convective fluxes are discretized using central differencing [8J.                      For incompressible flows,            the filtered  Navicr-Stokes An implicit fractional-step method (I SM) [10] in                equations can be written as conjunction with a second-order accurate, three-level backward- differencing scheme for time-discrelizition was employed to advance the solution in time. In this algorithill, the momentum equations are decoupled from the continuity                           aw, +2a~717   =_iI DJP arqij a [ ,aI, TX,-_-
ax,      -
(4) equation applying an approximate factorization of the coupled Navier-Stokes equations. For incompressible flows, the FSMI preserves the formal second-order temporal accuracy without            where -rv=u~u,-j            is the subgrid-scalc turbulent stress. In having to perform, at each time-step, costly outer iterations to       this study, the subgrid-scale stress is modeled using isotropic couple velocity-field and pressure. '[he FSNI thus provides a          eddy-viscosity as highly efficient algorithm for CPJU-intensive transiet computations like LES.
Consider the semi-discrete form of the Navier-Stokes                     TO,  110 -ra = -2C,25ZlMSij                               (5) equations in "pressure-correction" (4y",I = p"/' - p") form:
We determined the model constant, C,, using the dynamic
[A   G] u"-`
ID 0] 4,j; )
                                                 -=(r 0
(I)
Smagorinsky model (DSM) originally proposed by Germano et al. [7, 8].
by The filtered transport equation for a passive scalar is given where un`t and r are the velocity vector and the momentum source vector, respectively, and the integer n is the time level.
A is the coefficient matrix defined in terms of the discrete                          3+ 2aii= a (savP q                                (6) convective and diffusive operators and the timc-advancement scheme chosen, and G and D are the discrete gradient and a,     ax        axe    ax. i divergence operators, respectively. The coupled system of equations in Equation (1) is extremely difficult to solve as it is,    where q is the subgrid-scale turbulent flux of the species since the matrix A has to be inverted for every iteration. In the      concentration (s7), and a the molecular diffusivity. The subgrid-.
fractional-step method, the original coupled equations in scale flux is modeled using Equation (I) are approximated by
[A ID o1 AiDG IO' AZG1(
I u _=(r JV0+)t)
                                             + [O(A2)]            (2)
SC,   ax, (7)
Factorizing equation (2), we have a series of split operations as 2                                       Copyright &#xa9;2005 by ASME

where Sc, is the subgrid-scale tobtulent Schmidi     nmibecr. In this study, we used a constaln vaILue (f 0.9.
To implement the dynamic procedtlre for the present finite-volume solver requires a test-filter applicable to arbitrary unstructured meshes. The test-filter finally adopted for this study is a top-hat filter involving a volume comprising the cell in the center plus the neighboring cells that share the faces with the center cell. To make the dynamic procedure tractable, an approximation was made that is tantamount1l to a non-varying C, over the test-filter volume. Thus, the (Irilalic procedure employed in this study is "approximately local" in the sense that, despite the ad hoc assumption, it does not require an existence of any statistically homogeneous directions, being applicable to Figure 2. A sectional view of the compltational mesh complex three-dimensional flows. To avoid numerical instability          (medium-sized mesh wiith 2.7 niilliori lements) used in the likely to be caused by a large fluctuation of the model constant,        computations.
we smoothed the model constant by applying the test-filter on it, and also clipped it so that the effective viscosity remains positive.
The additional details of the implementation of the DSM in the                 The computational meshes were built using Gambit. We framework of an unstructured mesh based finite-volume solver             used three progressively refined, unstructured hexahedral can be found in the reference 18]. The DSNI has been validated           meshes. for the computations. The three meshes have 1.1 for a number of wall-bounded flows such as fully-developed               million (coarse), 2.7 million (niedium), and 4.8 million (fine) channel flows and flows around bluff bodies such as a square             elements, respectively. A sectional view of the medium-size cylinder and a sphere.                                                   mesh is shown in Figure 2. The resolution of the medium mesh is such that, in the mixing region, it can resolve the integral DETAILS OF COMPUTATIONS                                                  length-scale estimated around 20 mnm based on the experimental observation [1, 2], which is equivalent to D/6, with Solution domain, swirl vane geonmetry and meshes                         approximately 20 elements.           It can resol ve the smallest observable eddies of a size around 6 mm, which were observed A partial view of the computational domain is depicted in          in the experiment [1, 2], with 6 - 7 elements.
Figure 1, along with the coordinate system. The inlet boundary                 It is worth mentioning here that the mesh resolution in and is placed at LO.D upstream of the sw"irl generator, where D is           around the inner and the annular tubes and the swirl vanes is the downstream pipe diameter (D = 122 mm). The exit                      too coarse to resolve the energy-containing eddies originating boundary is at 15D downstream of the pipe expansion. The                 from and transported in that upstream region. We recognize computational domain thus has the swirl generator in it. The             that poor resolution of the turbulent structure in the upstream entire inlet tube, 8 swirl vanes, and the downstream pipe were           region will negatively impact the prediction in the downstream, modeled without using any periodic boundaries.                           particularly in the shear layers of the mixing region, inasmuch as the energy-carrying eddies from upstream feed the turbulent structure being developed in the downstream mixing region.
pipecexpaelmionl Furthermore, none of the three meshes are fine enough to accurately resolve the viscous sublayer on the pipe wall C-'
downstream of the expansion. This 'as deemed justifiable in light of a relatively passive role played by the wall downstream swirl vane of the expansion in the mixing occurring near the central core.
Boundary conditions, time-step size, and solution strategy On the annular jet inlet, a uniform axial velocity of 1.667 n/s was specified according to the data given by Roback and Johnson [1). For the velocity boundary condition at the inner inlte L     Z                                                        jet inlet, we used 0.797 m/s, a value derived from the volumetric flow-rate (6.2 gallon per minute) given in Ref. [1],
Figure 1. A partial view of the computational domain                     instead of the inlet velocity mentioned in the same report (0.52 mi/s). The inlet velocity of 0.52 n/s quoted in the report does The swirl generator consists of 8 identical vanes mounted          not match the given volume flow-rate, and is inconceivably too on the hub with an equal spacing in ihe azimuthal direction.             low considering that the measured axial velocity immediately The blade-section has a camber and a vane-angle that change              downstream of the expansion is around 0.8 n/s.
with the radius.         The blade geometry was taken from the                 On wvall boundaries, we employed a blended law-of-the-design data given in the Ref. [1]. A NURB surface was built              wall that invokes proper wall-laws depending on the local mesh based on the digitized surface geometry.                                 resolution, namely y/ at the wall-adjacent cells.             The downstream exit boundary was treated as a "pressure-outlet" boundary offered in FLUENT.              In essence, the solution 3                                     Copyright &#xa9; 2005 by ASME

variables are extiapol;died in a mass-conserving manner on this          'The LiS was slarle(l osing a steady RANS solution as the boundary.                                                          initial guess. To accelerate the solution to a statistically
        'rhe influence ol the exit boundary condition was          stationary'state, we superimposed a pseudo-random fluctuating discussed at length by Pierce and Moin (6]. They found that        velocity-field on the mean velocity field taken from the RANS the usual convective outtlow boundary condition applied on an      solution. Before tile statistics are collected, tile LES exit boundary with a cross-section of the downstream pipe          computation was run until the initial transients in the solution yielded a central recirculation zone that is far smaller than what are completely washed out, which typically took 2 - 3 flow-through times; one flow-through time is taken to be LlUb, where L is the length of the downstream pipe, and Ub the axial mean bulk velocity.
RESULTS Figure 4. Instantaneous velocity vectors in the r-z plane
     -E   *.    .
                                             . .- ,T k'.
                                                              ~
Figure 3. Instantaneous velocity vectors at four axial locations - top-left, z = 25 mm; top-right, z= 51 mm; bottom-left, z=1 02 mm, hottom-right, z = 203 mm was observed in the experiment. They were able to bring their LES prediction much closer to the experimental data by putting a second expansion before the exit boundary, which was conceived based on an argument that it better represents the settling chamber used in the experiment.
We also investigated the effect of the second expansion, comparing the result to what was obtained without he               Figure 5. Time-averaged velocity vectors in the r-z plane expansion.      However, we could not notice any appreciable difference between the two.           The insignificantly small difference found in our computations is perhaps due to the way           To obtain stable statistics of the solution, the transient the solution variables are extrapolated at "pressure-outlet"       computations were continued for a sufficiently long period of boundary that could be different from their treatment of the       time, typically for more than 7 - 8 flow-through times, until convective boundary.                                               the time-averaged velocity and species concentration fields The time-step size used in the present LES was determined    largely recover an axisynimletry.
based on the estimated characteristic time-scale of the smallest         Regarding the mesh-dependency of the solutions, it was eddies to be resolved in the LES. We took v = 1.0 m/s and .1       found that the LES results from the meditim mesh (2.7 million
= 0.05D (D = 0.122 m) as the characteristic velocity-scale and     cells) and the fine mesh (4.8 million cells) showed little length-scale of the smallest eddies. These estimates give an       difference, while the coarse mesh result deviates a little farther eddy-tumover time of v =e/v= 0.006 second. It was finally          from other two. Unless stated otherwise, the results presented decided to use a time-step size of At = 0.0002 second, which       in this paper are for the meditim mesh.
will resolves one eddy-turnover time (T) of the smallest resolved eddies roughly in 30 time steps.
4                                   Copyright &#xa9; 2005 by ASME

the penetration length of the center jet with high species concentration (colored pink in the figure in Figure 7) was found to be around 50 nmnm,                which is closely reproduced by our LES results as shown in Figure 7.
Figurc 6. Jnstantaneous velocity vector plots at two axial locations - left, z = 25 mrzs; right, z= 51 mmi Overall flo'llv Structure Figures 3 and 4 show the instantaneous velocity vectors projected on four crossflow (r-0) planes and a r-z plane. The vector plots portray turbulent eddies with widely varying length-scales throughout the mixing region. Small eddies are     Figure 8. Vortical flow structure in the mixing region shown to form in the shear layers between the inner jet and the  visualized by the iso-surface of the second-invariant of the outer annular jet, growing in size in the downstream direction. velocity deformation tensor, colored by velocity magnitude.
At z = 25 mnt, small eddies are confined near the shear layer between the inner and the outer annual jets. Yet, larger eddies           Figure 8 gives an overall impression of the turbulent are also seen to have formed in the annular recirculation region vortical structure in the mixing region. This figure indicates behind the step. One can visually tell from the figure that the  that the flow in this coaxial jets are well mixed.
smallest structure at z = 25 mm resolved in the LES is roughly D/20 (D = 122 mm), which is close to the size of the smallest    \'elocitv field eddy, 6 mm, observed by Roback and Johnson [I]. The size of the largest eddy at z = 25 mm, which is about two-thirds of the           Figure 9 shows the mean axial velocity profile along the step-height as shown in the figure, also closely matches the     centerline of the pipe. The LES prediction quite closely experimentally observed value, 20 mm, quoted in ref. [1].        reproduces the general trend such as the rapid drop of the mean Figure 5 depicts the mean velocity vectors on a r-z plane.       axial velocity immediately downstream of the expansion and the gradual recovery further downstream. However, the recirculation zone predicted by the present LES appears to be shifted slightly downstream compared to what is indicated by the measurement.
I1 .. .   .     . . .   .  .  . .   .  . . . . . . . . .   .
                                                                                       -      U.FS (present) 1.01          {0    Measured (Roback and Johnson, 193)1 i
                                                                  .9 C.                                                0          0 U
0   0 I
0, 0.0           2\                  a;~0 0\
                                                                                                   &deg;
                                                                               .. 0..1....
Figure 7.          Contours of the instantaneous         species        - I,.
concentration in r-z plane                                                                 10(           200          300       400      500 Axial distance from the expansion (mm)
Figure 6 and 7 depict the contours of the instantaneous     Figure 9. Mean axial velocity along the centerline of the species concentration on two crossflow planes and a r-z plane. pipe.
The contours on the crossbow planes give an idea of the length-scale of the turbulent structure. The observations from these-figures are consistent with those from the instantaneous velocity vector plots discussed earlier. In the experiment [1],
5s                                                 Copyright C 2005 by ASME

                                                                                                            -     e wo      -o z = 5 miii                                                                       z = 5 mm 1.5 Io   Measured (Roback and Johnson, 1983)
                                                                                -      Present LES 00 1OF-I      0        00 0
U                                                  0
(.51-U                              00 O. I.-        0   /
I                     I 0.2            0.4                   0.6      0.8 r/rO z = 25 mm e
                                                                    .3 I-U r/rO z=5l mm                                                                         z=51 mm 1.0 Measured (Roback and Johnson, 1983)
                                                                                 -     Prescnt LES oOIO        0   0
                                                                   >C 0.5 C
0    00        _0 M     0.
T13                0   000   o      8
                                                                  .el0.
[is,
                                                                        -U.Jz             0.2            0.4                   0.6       0.8 rlrO Figure 10. Mean axial and azimuthal velocity profiles at two axial locations (z = 25 mm, z = 51 mm)
Figure 10 shows the radial profiles of the mean axial and      obtained by averaging the radial profiles taken at four azimuthal velocity components at three axial locations (z = 5      azimuthal locations (0= 0&deg;, 9 0 &deg;, 1800, 2700). The predictions mm, 25 mm, 51 mm). The profiles shown in the figure were           are seen to capture the overall trends and the peak values of the 6                                                     Copyright &#xa9;2005 by ASME

velocity components. However, it is cle;rl) noliceable that the    used in this study, including tile fine mesh (4.8 mS1illion cells),
locations of the peaks predicted by the LES arc generally          are tine enough to accurately resolve the enerev-meollaijlling shifted toward the centerline. This implies that the outer         eddies generated in the upstream region.
annular jet expands or opens up less than it does in reality.           The r.m.s. axial velocity fluctuation shown in Figure I I This also means that the annular recirculation zone predicted by   seems to support our reasoning. The r.m.s. axial "elocity the present LES is longer (in the axial direction) than in the     fluctuation at the centerline (r = 0) at z = 25 inn: show'n in the experiment. The underprediction of the jet angle is also           top-figure, and the one at z = 5 min (not shown here) are correlated with the downstream shift of the central recirculation  severely underpredicted in the LES computation. As shown in zone discussed earlier.                                            the bottom-figure for z = 51 mm, further downstreami ol the jet exit, the r.m.s. axial velocity fluctuation calchies up with the z = 25 mm                           data, as the turbulent eddies generated in tile shear layer become full-fledged.
The impact of the incoming turbulent eddies on the mixing in the downstream as suggested above and the cost implication of using an extremely fine mesh in the upstream region beg a E 0.6 question; what would be the best practice that can be adopted to obtain a sufficiently accurate prediction of the fow and species mixing occurring in a coaxial jet combustor with LES? We
 , 0.4                                                             will ponder a little upon this question at the very end.
Species concentration 0A c:- 0.21                                                               Figure 12 shows the profile of the mean species concentration along the pipe centerline. The LES prediction closely reproduces the trend - the plunge of the mean species concentration occurring near z = 50 mint.            One noticeable discrepancy between the prediction and the mcasurenlent is in the length of the inviscid core for the species concentration. The LES z=51 mm                             yields an inviscid core for the species concentration almost down to z = 35 mm, whereas the experimental data indicates that the mean species concentration starts being diffused away almost immediately, after the expansion. We think that this is again, in a major part, due to the poor resolution of the incoming turbulent eddies which would take part in "tearing" the inner jet.
 .0 U2 T/ro Figure 11. rams. axial velocity fluctuation at two cross-flow planes (z = 25 mm, z = 51 m)
We surmise that, among others things, the most likely culprit for this discrepancy is the lack of mesh resolution in the region upstream of the expansion, in and around the inner and the outer annular pipes, and the swirl vanes. As mentioned         Figure 12. Mean species concentration along the centerline earlier, turbulent eddies coming from the upstream region feed     of the pipe the shear layers developed downstream of the expansion, enhancing the mixing of the momentum and the species concentration in the inner and the annular jets with the                The radial profiles of the mean species concentration at two axial locations are shown in Figure 13. The predictions are in surrounding flow, which will lead to an increase in the jet angle. In our LES computations, these energy-feeding eddies        good agreement with the measurements at both locations. The are almost missing, since none of the computational meshes         mean species concentration profile at z = 25 mm predicted by the LES computation shows a sign of being "under-diffused", insofar 7                                    Copyright 3 2005 by ASME

as it has a fuller profile than the   leastie(l one. The prediction at r.m.s. value at z = 25 nmm, located in tile sheal layerl between the z = 51 mm is somewhat lower than the measurement near the              inner and the outer annUlar jets, is ;lso considerably centerline. However, the overall agreement at this location is         underpredicted.      Apparently, the LIES utnderpredicts the remarkably good in view of the steep change of the mean species        entrainment of the ambient fluid occurrineg at this location. At z concentration near z = 51 mm, as can be seen in Figure 12. It          = 51 mm, the prediction conies milc h closer to the should be noted, in passing, that the RANS predictions reported        measurement, as the r.m.s. velocity luc1tluationr did at the same in the literature [3] severely under-predicled the mean species        axial location.
concentration at this axial location.
z = 25 nini                                                           z = 25   mini C
0 C
0 U
2 CJ z=51 mm                                                              z = 51 nini r/r11                                                                 r/ro Figure 13.        Mean species concentration at two axial              Figure 14. r~m.s. species concentration at two axial locations locations (z = 25 mm, 51 mm)                                           (z = 25 mm, 51 mm)
Equally important - probably even more important than the             The same remarks given regarding what might have caused mean species concentration in the context of modeling                  the discrepancy in the r.m.s. velocity fluctuation (Figure 11) turbulent combustion, is the fluctuation of species                    largely apply to the results for the r.m.s. fluctuating species concentration. The usual RANS-based turbulence models                  concentration.
cannot directly predict the r.m.s. fluctuating species concentration, unless the transport equation for the variance of fluctuating species concentration is explicitly solved. One            PRELMINARY RESULTS WITH LOCALLY REFINED obvious benefit from LES is that one can directly compute it.          MESHES Figure 14 depicts the r.m.s. fluctuating species concentration at the two crossflow planes. At z = 25 mm, as in the case of the               The fact that the fine mesh (4.8 million cells) offers a r.m.s. axial velocity fluctuation, the LES result grossly              meager improvement over the medium mesh (2.7 million cells) underpredicts the r.m.s. value in the core region. The peak            in terms of the mesh resolution (spacing) warrants additional 8                                      Copyright ( 2005 by ASMIE

Computations vsith siuhbanitially finer mesh, for instance, by halving the grid sspacing in all three (axial, radial, and         CONCLUSION azimuthal) directions.            However, doubling the number of elements in the three directions results in an eight-fold increase     The flow and tile species transport in confined swirling in the total cell counts, which becomes unwieldy.                  coaxial jets were conimpited using LES. An unstructured mesh-based finite-volumne solver was employed for the computations.
A highly efficient timc-advancement scheme was used in z=51 mm                      conjunction with second-order accurate temporal and spatial
: 2.  .-
discretization schemes.       A dynamic Smagorinsky model M acasurcd (Rotack & Johnson IlIl)              adapted to general three-ditnensional flows was employed as 2.0 _    -   ILES (2.7NM ccli mesh)                           the subgrid-scale turbulence model.
              ---  LES (2.7NM cell & locaily refined)
For LES, it is evidently a bold attempt and a costly E 1.5i                                                           proposition to include the upstream components of a combustor like the swirl generator and the upstream tubes in the I .0                                                          computational domain. Nevertiheless, the present LES a;                                                                computations closely reproduce the salient features of the flow C
M 0.51                                                          and the species concentration in the mixing region. For an accurate prediction of the mixing in the downstream (e.g.,
V                                                                combustion chamber), a good resolution of the mean flow and 0.0 0    CS                                       the turbulence in and around the inner and the outer annular tubes, and the swirl vanes turned out to be more important than
     .0.5                                                          originally thought. The numerical evidence found from this D          0.2         0.4          0.6      0.8       study indicates that the three globally refined meshes used in r1r,                      the present study, despite the largest cell counts reaching up to 4.8 million cells, still cannot provide a sufficient resolution of z=51 mm                      the upstream region. We believe that the lack of mesh I .0
                . Measured                                        resolution is responsible for the discrepancy between the
              -    LES (2.7NM ccll mesh)                           predictions and the measurements, most notably the overall
              --- LES (2.7M cell mesh & locally refined)           shift of the velocity peaks toward the centerline.
Finally, a preliminary result was presented which was 1   0.5 I    0o~            -                    obtained with a new mesh locally refined in the upstream part
 .22:1                                                             of the domain. The significant improvement from this locally V                                                   00           adapted mesh supports our conclusion, and at the same time, provides us with an avenue to improving the accuracy of LES 1>                                                               prediction for the subject flow.
C ri 0.0 4!
ACKNOWLEDGEMENT I            I     .                 I          The authors acknowledge that FLUENT and GAMBIT 0.2         0.4          0.6      0.8       software were used for this study. The LES-cluster in Fluent r/ro                      Inc. was used for the computations.
Figure 15. Mean axial and azimuthal velocity profiles at two axial locations - predicted with a locally refined mesh        REFERENCES To keep the mesh size under a tractable limit, the medium    [I] Roback, R. and Johnson, B. V., "Mass and Momentum mesh was locally refined only in the domain upstream of the             Turbulent Transport Experiments with Confined Swirling expansion. Furthermore, only the cells within a specified               Coaxial Jets," NASA CR-168252, 1983.
distance from the wall, roughly 0.07D in this study, were          [2] Johnson, B. V. and Roback, R., "Mass and Momentum refined to further save the cell counts. The local refinement           Turbulent Transport Experiments with Confined Swirling resulted in a total of 6.5 million cells. The computation has           Coaxial Jets - Part I," AIAA-84-1380, Presented at been carried out on this mesh. The results are shown in Figure          AIAAISAE/ASME 20 'h Joint Ppropulsion Conference, 15, being compared with the medium mesh results, for the                Cincinnati, OH, June II - 13, 1984.
radial profiles of the mean axial and azimuthal velocity           [3] Brankovic, A., Ryder, Jr. R. C., and Syed, S. A., 1998, components at z = 51 mm. Clearly, the predictions with the              "Mixing and Combustion Modeling for Gas Turbine adapted mesh are significantly improved over the predictions            Combustors Using Unstructured CFD Technology," AIAA with the medium mesh. The overall shift of the velocity peaks           Paper 98-3854, Presented at 34th AIAA/ASME/SAE/ASEE is now far smaller than what we saw earlier, and the predictions        Joint Propulsion Conference & Exhibit, July 13 - 15, 1998, capture the profiles of the mean velocity components much               Cleveland, 011.
more closely.
9                                  Copyright &#xa9;2005 by ASME

14J l in. C.- A., "IModeling a Contined Swirling Coaxial Jet,"
Annzual Ressearcht Brief, 1998, Center for Turbuilenice Rtesearchl., Stanford University, ]998.
15] Akselvoll, K. and Moin, P., 1996, ''Large-Eddy-Simullationl of Turbulent Confined Coannuair Jets," J. fluid ine(chI.,
315,pli. 3 87-41 16] P'ierce, C. D. and Moin, P., 1998, "Large Eddy Simulation of a Confined Jet with Sw~irl and Heat Release," AIAA Paper, 9S-2S892.
171 Germano, M/, Piomelli, U., Moin, P., and Cabot, XV. II.,
1991, "Dynamnic Subgrid Scale Eddy Viscosity Model,"
Physics of Fluidls A, 3, 19, pp. 1760 - 1765.
18] Kim, S. E., 2004, "Large Eddy Simulation Using U~nstructured      Mesh and Dynamic Subgrid-Scale Turbulence Models," AIAA Paper 2004-2548.
191 Kim, S. E., Makiarov, B., and Caraeni, D., 2004, "Mulfli-.
Dimensional Reconstruction Scheme for Unstructured Meshes," AIAA Papser 2004-2548.
[10] Kim, S. E. and Makarov, B., 2005, "An Implicit Fractional-Step Method for Efficient Transient Simulation of Incompressible Flowvs," To be presented at 17'h AIAA Computational Fluid Dynamics Conference, June 6 - 9, Toronto, Ontario.
10 Copyright C)2005 by ASME

BVY 05-072 Docket No. 50-271 Exhibit EMEB-B-138-3 Vermont Yankee Nuclear Power Station Proposed Technical Specification Change No. 263 - Supplement No. 30 Extended Power Uprate Response to Request for Additional Information Large Eddy Simulation Using Unstructured Meshes and Dynamic Subgrid-Scale Turbulence Models.
l Total number of pages in this Exhibit I (excludinc this cover sheet) is 17. l

34th AIAA Fluid Dynamics Conference and Exhibit. June 28 - July 1, Portland Oregon Large Eddy Simulation Using Unstructured Meshes and Dynamic Subgrid-Scale Turbulence Models Sung-EunL Kim*
Fluent Inc, Lebanon, Nrew Hampshire, 03766, U.S.A.
This paper concerns development of a large eddy simulation (LES) capability based on a finitc-volume solver that permits use of unstructured meshes. The solver employs a cell-centered scheme along with a multi-dimensional linear reconstruction. Convection and diffusion terms are discretized using a second-order central-differencing scheme. A three-level second-order scheme is used for temporal discretization. Subgrid-scale turbulent stresses are closed using dynamic Smagorinsky model and dynamic turbulent kinetic energy transport model. A test-filter was designed for the dynamic procedure which is applicable to unstructured meshes ofarbitrary cell-topology. The dynamic procedure also avails itself to three-dimensional flows without any statistically homogenous directions. W\'all boundary conditions are imposed using a wall-function approach that applies appropriate wall-laws depending on near-wall mesh resolution. The LES capability is validated for a wide range of wall-bounded flows. We present here the results for a fully-developed channel flow and two bluff-body Blows. The predictions are in good agreement with direct numerical simulation (DNS) results and the experimentalaiata
: 1. Introduction IITE encounter many industrial applications of computational fluid dynamics (CFD) where the flows are dominated VWby unsteady, large-scalc coherent structures. Those large-scale structures impact, to a great cxtcnt, various aspects of the flows such as energy consumption, safety, comfort, and noise. The ramification of whether or not one can harness the large-scale structures is therefore quite significant. Attempts to numerically predict such flows using unsteady Reynolds-Averaged Navier-Stokes (URANS) equations have been met with limited success. Some of the better RANS models seem to be capable of capturing the "Jargest" scale occurring often in the form of alternate vortex-shedding. However, the remaining coherent structures are left largely unresolved.
Large eddy simulation (LES) is fundamentally suited to the task of predicting coherent structures. The major obstacle in using LES for practical high Reynolds-number (Re) flows, from the practitioners' standpoint, is its high cost incurred by an unwieldy number of computational elements and painfully long solution time. Yet today's ever-increasing computing power is rapidly making LES feasible. Another difficulty often encountered when attempting LES for industrial applications comes from complex geometry. In CFD, meshing for industrial applications involving complex geometry by itself can become a grand challenge. It is hugely time-consuming or often impossible to generate high-quality structured meshes for complex configurations, which has led industrial CFD practitioners to opt for unstructured meshes. Although unstuctured meshes are routinely used today in RANS computations for industrial applications, attempts to conduct LES with unstructured meshes are just starting to appear in the literature.1- 5 As yet the efficacy of unstructured meshes for LES for practical high-Re flows has not been fully established. Among the issues yet to be addressed are numerical accuracy, stability, and subgrid-scale turbulence modeling.
This paper is concerned with evaluating a LES capability developed using a finite-volume solver based on second-order numerics. Permitting use of unstructured meshes, the resulting LES capability can easily handle complex ge-ometries encountered in industrial applications. In addition, it lends itself to local mesh adaptation that can be utilized
   'Principal Engineer, Fluent Inc., Lebanon N.H., Menmber AIAA.
Copyright 0 2004 by Fluent Inc.. Published by the American Institute of Aeronautics and Astronautics, Inc. with permission.
I of 17 Arnerican institute of Aeronautics and Astronautics Paper 2004-2548

to efficiently allocate computational cells, substantially reducing the computational cost. Adequacy of second-order spatial discretization for LES has often been questioned, and there are some misgivings about using it for LE-S. It will be shown in this paper, however, that the second-ordercentral differencing scheme adopted in the present work yields a commendable accuracy for LES. For subgrid-scale (SGS) turbulence closure, we implemented two dynamic SGS vis-cosity models in the framework of unstructured meshes, namely, the dynamic Smagorinsky model originally proposed by Germano et al.,6 and Lilly' and the dynamic turbulence kinetic energy transport model of Kim and Menon. 8 9 For the dynamic procedure, a test-filter readily applicable to unstructured meshes was designed. The resulting dynamic SGS models can be used for three-dimcnsional flows without any statistically homogeneous directions.
The paper is organized as follows. We start with a brief description of the two dynamic subgrid-scale turbulence models and the details of their implementations. This will be followed by an overview of the numerical methods and algorithms adopted in this work. Finally, validations will be presented for a selected number of wall-bounded flows ranging from a fully-developed channel flow to a couple of bluff-body flows including one around a squarc-cylinder with salient edges and another past a smooth sphere.
II. Filtered Navier-Stokes Equation and Subgrid-Scale 71urbulencc Modeling' A. Implicit filter with finite-volume discrelization The governing equations for LES are generally obtained by filtering the Navier-Stokes equations in either Fourier (wave-number) or physical space. In the present work, the filtering operation (denoted by an overhar) is defined as a spatially convoluted integral of the variable in question as
                                                      &#xa2;(X) = l       (y)G(x,y)dy                                           (1) where D is the computational domain, y E VD and G is the filter function.
With the cell-centered finite-volume discretization and the linear reconstruction scheme employed in this work, the discrete solution variable at a cell-center (co) can be written as C(co) =      f,    (Y)d, y C        V                                    (2) where V is the volume of a computational cell.
The definition in Equation (2) of a discrete solution variable at cell center can be interpreted as a filtering operation
                                             &#xa2;(x) -(CO)       = V         (y)dy,y E V                                      (3)
The implied filter function, G(x )), is then a top-hat filter G(x,y) = {/V                 for Ix -     El                                   (4) 0           otherwise Using the filtering operation in Equation (3), the filtered Navier-Stokes equations for incompressible flows (as-sumed for brevity) can be written as
_+ a            _       ap arij + a                                               (5) at      axj -           Tpx,a5x-       axj     axjJ anti where   'rij is the subgrid-scale stress defined by; 7jj-U_7j
                                                                       - Rj                                                (7) which is unknown and needs a closure.
Thus, we used in this work grid and finite-volume discretization as an implicit filter.
2 or 17 American Institute of Aeronautics and Astronautics Paper 2004-2548

B. Subgrid-scale (SGS) stress models The SGS models based on the concept of isotropic eddy-viscosity compute the SGS turbulent stress from Tjj - 3Tkk 5 iJ = -2VS3j                                                       (8) where v, is the SGS eddy-viscosity, and S3j the resolved rate-of-strain tensor defined by
                                                          -      I / aTh         a~\
ii      2       xj      DxiJ The task of SGS turbulence modeling is to express SGS viscosity, v,, as a functional of known quantities. In the present work, we employed two dynamic SGS eddy-viscosity models. They are described below.
: 1. Dynamic Smagorinsky AModel (DSAM)
The underpinning of DSM is the algebraic eddy-viscosity model originally proposed by Smagorinsky.' 0 In the original Smagorinsky's model, SGS eddy-viscosity is computed from v, = c          jsj 1sl                                                     (9) where Cv is a model constant (Cv 0.1 0.2), SI                              _       the/ modulus of rate-of-strain tensor of the resolved scales, and a VI/ 3.
The subgrid-scale stress can then be written as Tij -    .tA       =-2cvA          IsI si                                         (10)
Despite its simplicity, this model has several shortcomings. The most problematic one, from a practical standpoint, has to do with the model constant, Cv. There is no single value of the constant which is universally applicable to a wide range of flows. Another serious drawback is that the SGS viscosity model in Equation (9) with a constant \,'lue of Cv is not applicable to transitional flows where the flow in question is laminar either locally or intermittently, since Equation (9) always gives a finite SGS viscosity even in laminar region as long as there is velocity gradient.
Germano et aL6 and subsequently Lilly 7 conceived a procedure that resolves these problems. In this procedure, Cv is dynamically computed during LES, on-the-fly, using the information provided by the smaller scales of the resolved (known) fields. To separate the smaller scales from the resolved field, the dynamic procedure needs a so-called "test-filter" having a width (A) that is larger than the grid-filter width (A). Denoting the test-filtered variables by a tilde and putting the "grid-filtered" Navier-Stokes equations through the test-filter, we obtain "test-filtered" Navier-Stokes equations as G-i +       -          --            I
                                                                            -+-             v-I                                     (11) at       axi             TX,       aXJ aXj          A )
                                                                   -     =0                                                         (12) ax, where Tj is now a "subtest-scale" stress defined by Tij =1iftiJ-Thlij                                                      (13)
The underpinning of the dynamic modeling is a similarity concept that Tij, the subtest-scale stress, can be written as a functional of the larger resolved scales in a manner similar to Tij, 6 which leads to 3 of 17 American Institute Of Aeronauhics and Astronautics Paper 2004-2548

T j "jT         =-2
                                                                 =-           S Sij                                    (14)
T2v T where Q is the test-filter width. Note that the same model coefficient, C., is used in the expressions for both tij and Tij.
Equation (14) alone is not helpful at all in determining Cv, because Tij is not known in LI-S. The breakthrough came from the realization that 'Tij and Tjj are related to each other by 6
                                                      =    Uij-U      in(tj-    j-I-izxj
                                                      =    uizj-iij_                                                   (15)
The stress, Lij, which might be called subtest-scale Leonard stress, can be interpreted as the stress associated with the smaller resolved scales between the test-filter scale (a) and the grid-filter scale (a). Since Lij can be directly computed from the resolved scales in LES, the identity given by Equation (15) can be used to detenninc C.,.. Thus, we have Ljj- -} Lkk      = ajCv - OiJCv                                       (16) where aj=          -2A2 S Sj                                           (17) 2 1SISij                                        (18)
One predicament that makes it difficult to determine Cv from Equation (16) is the fact that C, in the second term on the right-hand side of the equation is under the test-filtering operator. As it stands, Equation (16) is an integral equation for Cv as discussed at length by Ghosal et al.1 1 This difficulty can be avoided by taking out Cv from the Lij- 8jLkk = Cv (aijr-ij)                                             (19)
We followed this rather ad hoc approach in this work despite its mathematical inconsistency, which amounts to as-suming that Cv remains constant in the computational cells associated with the test-filter.
Since there are more equations in Equation (19) than the unknowns, the model constant Cv is obtained by seeking for Cv which minimizes the error norm defined by E== (Lij-8j_        Lkk-CVMij)                                         (20) where A1j =aij    -     =-2
                                                           -i      (a2S        _-   2 I [)
Taking aE/aCv and setting it zero, we obtain
                                                              -      UM i=                                                (21)
Cv determined in this way varies with time and space. In fact, it varies in a wide range, often taking either large negative or positive value. Although negative Cv and consequently negative eddy-viscosity is often interpreted as representing "back-scatter" (flow of energy from smaller to larger scales), too large a negative eddy-viscosity causes numerical instability, ultimately leading to divergence of numerical solutions. The usual remedy for this numerical difficulty is to average Cv in statistically homogeneous directions. Obviously, this workaround can be exploited only when there are such homogeneous directions, which is a rarity in practical applications. Even if there are any statisti-cally homogeneousdirections, the averagingbecomesextremelycumbersomewith unstnictured meshes. In the present 4 of 17 American Instilute of Aeronautics and Astronautics Paper 2004-2548

work, we simply "condition" the C, computed by Equation (21) by test-filtering it. This simple volume-weighted aver-aging better preserves the locality of the model constant. To ameliorate the potential numerical difficulty caused by Cv being negative for an extended period of time, we evaluated two alternative approaches. In one approach, we simply clip Cv at 0 when it becomes negative. This option therefore rules out any chance for the model to mimic backscatter.
In another, effective viscosity (v+v,) instead of Cv is clipped at zero, permitting small negative SGS viscosity to happen. As yet we do not have any conclusive evidence that supports superiority of one approach to another in terms of prediction accuracy, except an indication that computations with teie first approach appear numerically more robust.
For this reason, we used the first approach in the computations presented in this paper.
As mentioned earlier, the dynamic procedure described above requires a test-filter. The most important criterion the test-filter should satisfy specifically for this work is that it should be applicable to unstructured meshes of arbitrary cell topology without incurring unduly high cost. The test-filter finally adopted is a top-hat filter that involves a volume comprising the cell itself plus the neighboring cells that share the cell faces with the center cell. Thus, the test-filter operation amounts to a volume-weighted averaging of the variable in question, which is easily implementable and takcs advantage of the data structure of the underlying finite-volume solver. With hexahedral mcshcs, the ratio of the test-filter to the gird-filter scale (A/A) is approximately 2.1 (= 91/3). The ratio for tetrahedra is smaller, being around 1.7(= 51/ 3 ).
: 2. Localized Dynamic Kinetic Energy Model (LDKEAI)
The dynamic Smagorinsky's model described so far is essentially an algebraic model in which subgrid-scale stresses are modeled using the resolved velocity field. A more elaborate SGS stress model would be the one which is directly based on SGS turbulence and can be used to parametrize SGS stresses. The most widely used ones among others are what can be called "one-equation" models in which SGS turbulent kiinetic energy, kg, = ( - -i)/2,          ik2 is explicitly computed by solving its transport equation. 8' 9' 11,12 In the present work, a localized dynamic k,,,-equation model of Kim and Menon 8 9 was chosen in favor of its overall efficacy.13 14 In the LDKEM, the subgrid-scale eddy viscosity, v,, is computed from vI =   Ck     A                                               (22)
Consequently, SGS stress is can be written as 2rj-3k,, = -2Cxk:,g, ASjj                                        (23) k.., is obtained by solving its transport equation"' 9 a,
                                      +1 aTx,
                                                   =   _Tij axi
_C~_ - + a TXj      a,
                                                                                              +V )   axi
                                                                                                         ](24)
The only difference between the original formulation and the present one is that the contribution from the molecular diffusion of keg, is included in this work-I As shown above, there are three model parameters appearing in these equations, namely, Ck, Cc, and ckk, which need to be specified. In the current implementation of the LDKEM, the first two are determined from the dynamic procedures to be described in the following, whereas Ok is simply set to a constant value of 1.0. The underpinning of the dynamic procedure employed in the LDKEM is the hypothesis corroborated by the experimental evidence 15 ' 6 that there is a strong correlation between the subgrid scale stress, rij, and the subtest-scale Leonard's stress, Loj. In place of parameterizing Tj and utilizing the Germano's identity as was done in the DSM, the LDKEM models Lij directly as L;}-      Lkk = -2CkAk-,,        Sij                                    (25) 5 of 17 American Institute of Aeronautics and Asironamilics RPa.r 2004-2548

where ke,, is the resolved turbulent kinetic energy associated with the scales between the test-filtcr (A) and the grid-filter (A). It can be directly computed from kl.%= 214    (Z I k 14       itt-ttl                                    (26)
The model parameter, C*, can then be determined from Equation (25) by minimizing the error norn as in the DSM.
Consequently wc have CA _(27) with Alij defined by
                                                         >1;j = -2Ak-,/Ssr Sij                                             (28)
The model parameter, CE, of the dissipation term in Equation (24) is also determined by a dynamic procedure, whose underpinning is the hypothesis that the dissipation-rate of kle (e) can be expressed in the same functional form as the dissipation-rate of k 1-3/2 e = Ci ""                                                    (29)
The dissipation-rate of klet can also be computed from e=(V+Nv,)          D~-aX an)       1 (30)
(vv(axi acti         axi ax-i)
From Equation (29) and Equation (30), C' is given in a closed form A(V+vt) (ai,, ai a h'-                                                 (31) keg          axj axj axj a.rj The DKEM has several desirable attributes that the DSM lacks. First, as a consequence of parameterizing Lj directly, the dynamic procedure in the LDKEM does not involve any test-filter operation on the model parameter, C*.
Thus, unlike the DSM model, C'k is a genuine, local quantity free from any mathematical inconsistency. Secondly, Ck in the LDKEM behaves numerically more benignly than Cv in the DSM, having much less fluctuation. As a small premium, one can even save a small amount of computational effort with the LDKEM, inasmuch as the test-filtering on the SGS stress done on in the DSM is not needed. Lastly and probably most important, the LDKEM enjoy~sthe benefits of a high-order turbulence model. Adopting SGS turbulent kinetic energy to parametrize the SGS stress renders the LDKEM better suited to non-equilibrium flows. By accounting the ktg,-budget more rigorously, backscatter of kinetic energy is allowed in the LDKEM on a much sounder physical basis than in the DSM.
III.      Boundary conditions WVall boundaries are the most crucial and yet difficult ones to handle in LES. In LES resolving all the way down to viscous sublayer, no-slip would be clearly the choice for wall boundary condition for the resolved velocity field.
However, the cost of such "wall-resolving" LES is prohibitively high for practical flows involving high-Re flows. A practical alternative is to use the law-of-the-wall bridging the wall and the first grid point (cell center) off the wall. The simplest implementation of the wall-law can done using U'      +'y            for y+ < ),                                     (32)
                                                          '=
ln(Ey+)         fory+>y+
where E = 9.793, K = 0.419, y+-=Uy/v,               _1u/i, yv+ is the "cross-over"at which the two wall laws intersect. This
                                                    =+
demarcation of the entire inner layer into the two distinct layers is apparently a simplification which is at odds with the 6 of 17 American Institute of Aeronautics and Astronautics Paper 2004-2548

presence of a buffer region in reality between the viscous sublayer and the log-layer. However, this is much better than exclusively relying on no-slip condition. One can mimic the presence of the buffer layer by blending the linear and the logarithmic laws using an adequate blending function. The blending has some merits, inasmuch as, besides giving a smooth transition between the two layers (numerically more stable) and representing the mean velocity profile in the buffer layer more accurately, the blending also allows the respective laws in the two regions to be independently modified or extended to take into account other effects such as pressure gradient, surface roughness, and transpiration.
In the present work, we employed the blending function suggested by Kader. 17 In the finite-volume discretization adopted in this work, the blended wall-law is employed to compute wall-shear, essentially diffusion flux at wall. To that end, the wall-law is applied to the parallel components of the resolved velocity at wall-adjacent cells to compute the friction-velocity (Far) in an iterative manner.
The LDKEM needs a wall boundary condition for k                  To that end, either a Dirichlet-type boundary condition (ksg, = 0) or a Neumann boundary condition (Dk-g,/1n = 0, in view of kgi )92) is conceivable. In the present work, we simply set the diffusion flux of k.,,a at walls to zero, which is essestially equivalent to the Neumann boundary condition.
I\'     Flow solver The present work was carried out using FLUENT, a general-purposc CFD code. FLUENT employs a cell-centcrcd finite-volume method based on a multi-dimensional linear reconstruction scheme, which permits use of computational elements (cells) with arbitrary olyhedral ppology, including quadrilateral, hexahedral, triangular, tetrahedral, pyra-midal, prismatic, and hybrid meshes. There are several choices of the solver algorithms in FLUENT including coupled explicit, coupled implicit, and segregated solvers. For the computations presented in this paper, we used the segregated solver exclusively.
Inthe segregated solver, the goverm-g e-quationsa-re solv-ed s-equ-enitally.-Sev-eraditerent solution--algoritihms are offered including SIMPLE, SIMPLEC, PISO, and fractional-step method (FSM). The temporal discretization in the segregated solver employs a fully-implicit, threc-level second-order scheme. Time-accurate solutions can be obtained using either iterative time-advancement (ITA) scheme or non-iterative time-advancement (NITA) scheme. The NITA scheme greatly saves CPU time, since the costly outer iterations arc not needed. Unless stated otherwise, wc used the fractional-step method in conjunction with the NITA scheme in this study.
Accurate spatial discretization is crucial in LES. The spatial discretization schemes employed in this work are based on a multi-dimensional linear reconstruction scheme.18 20         - Diffusive fluxes are discretized using central differ-encing. Discretization of convective fluxes requires caution in LES. Upwind-biased schemes such as second-order upwind, QUICK, and third-order MUSCL schemes have been most widely used for RANS computations. Unfortu-nately, numerical diffusion introduced by upwind schemes, which might be acceptable in RANS computations for high Reynolds-number flows, is detrimental to LES. This is because, in LES, numerical diffusion, however small it is, can easily overwhelm physical diffusion. This is the case even with high-order upwind schemes. For this reason, for LES, central-differencing schemes have been preferred for their meritoriously low - or zero in ideal conditions -
numerical diffusion. Thus we added a second-order central differencing (CD) scheme for discrctization of convective terms specifically for LES. 21 Unfortunately, any pure CD schemes are susceptible to producing unphysical oscillations in the solution fields, which becomes especially pervasive in high Peclet-number situations - low diffusivity and coarse mesh which is almost a norm in LES for industrial applications. The usual remedy is to add a modicum of numerical dissipation, either explicitly or implicitly, to suppress the oscillations at the price of sacrificing spatial accuracy. In our implementation of CD, however, no numerical dissipation was explicitly added. And, unless stated otherwise, the CD scheme was used in this work. To back up the CD scheme in case it fails, we also developed what may be called a bounded central differencing (BCD) scheme that essentially detects in the solution fields any wiggles with a wave length of 2Ax or less (%< 2Ax) and suppress them by switching to upwind schemes of varying orders depending on the severity of the wiggles, while retaining the CD elsewhere. It should be emphasized that the BCD scheme significantly differs from the often-employed hybrid schemes blending central differencing and upwind schemes with a fixed ratio.
The BCD scheme is reserved for industrial applications involving high-Reynolds number flows and less-than-ideal 7 of 17 American Institute of Aeronautics and Astronautics Paper 2004-2548

meshes.
The discretized algebraic equations are solved using a point-wise Gauss-Seidel iterative algorithm. An algebraic multi-grid (AMG) method is employed to accelerate solution convergence. The solver is fully parallelized, which is crucial in LES for industrial applications.
V     Validations The LES capability described so far has been validated for a wide range of wall-hounded flows from simple. to complex ones. In this paper, we present the results of a fully-developed channel flow and two bluff-body flows. The channel flow case is a fundamentally important case whose subtlety offers an opportunity to critically evaluate various aspects of SGS turbulence modeling such as the dynamic procedure used to determine the model constants.
The bluff-body flows include the one around a cylinder with square cross-section at a moderately high Reynolds number and the one around a sphere at two Reynolds numbers. Deliberately chosen, both involve large-scale, co-herent structures around the bodies and in the wake, representing typical bluff-body flows encountered in industrial applications.
A. Ftilly-developed channel flow at Ret         =   180 A fully-developed channel flow was computed for the Reynolds number of Ret = 180 (ReH = 3,300) using the two dynamic models. The computational domain is a box of the size [2iTH x 2H x 7tH] in the axial, normal, and spanwisc directions, respectively. The computational domain is bounded by two walls on the bottom and the top of the channel, two pairs of periodic boundaries in the axial and the spanwise directions. The computations w"re carried out using two hexahedral grids; a coarse grid with 36 x 36 x 36 cells and a globally refined mesh with 72 x 72 x 72 cells. The resolutions of the meshes are such that, with the coarser mesh, y+ value at the wall-adjacent cells is approximately 0.6, and the cell size is dy+ = 27 near the channel centerplane. The channel walls are treated effectively as no-slip boundaries due to sufficiently low y+ values at the wall-neighboring cells. On the pair of periodic boundaries in the axial direction, a pressure-drop across the pair derived using the given wall-shear (Xs, = pu2) was specified, with the flow-rate determined as a part of the solution. The time-step size of At+ = 0.3 was used, where At+ = At 112/v. The CD scheme was used for the discretization of convection terms.
The mean axial velocity (U+) and the three r.m.s.velocity components (uf+, V+, iV+) predicted using the two dy-namic models are shown in Figure 1 on the following page along with the DNS results of Kim et al.22 The predictions with the coarse mesh are seen to overpredict U+ by about 8 - 12% near the center of the channel. The peak in the W'+
profile is also overpredicted with the coarse mesh, whereas the peaks in the profiles of Vi+and sw1 are underpredicted.
Our results with the coarse mesh show largely the same trends as found by others who employed grids of somewhat finer resolutions (32 x 64 x 32 mesh,2 3 65 x 65 x 65 mesh'), yet closely matching their predictions despite the coarser mesh employed in this work. However, our predictions of the r.m.s. velocity components near the channel center is relatively poor. We surmise that the much larger grid spacing near the channel center (Ay+ = 27) is responsible for that.
The predictions improve greatly with the fine mesh. Both dynamic models reproduce the DNS results remarkably well. The mean axial velocity and the r.m.s. fluctuating velocity components are accurately predicted throughout the entire range of y+. Particularly noteworthy is the excellent agreement with the DNS data in terms of the peaks values and their locations of the r.m.s.fluctuating velocity components. Overall, the results obtained with the fine mesh compare favorably with other results mentioned earlier." 23 For instance, our predictions are substantially closer to the DNS data than the results of Haworth and Jansent who computed the same channel flow using LES on a 65 x 65 x 65 node mesh using the Lagrangian dynamic Smagorinsky's model.
* The present results are promising, inasmuch as they demonstrate that the second-order CD scheme in conjunction with the dynamic models is able to accurately predict this fundamentally important wall-bounded flow carrying an intricate near-wall physics. Regarding the impact of SGS modeling, we did not find any significant difference between the results from the two dynamic models. This should not come as a big surprise, however, since the fully-developed channel flow is near equilibrium in the mean.
                                                              .8 of 17 American Institute of Aeronautics and Astronautics Paper 2004-2548

(a)Mean velocitv (Ut I nredictions                                         (b) r.m.s. vclocitv (u"' nredictions t>         5//.                                                              +0.6   .v 0.4  Ft    /       -    Prese, DSNI (36x36x36 meh)               _                                  Presnt. DSM (36x3
           /        ----   Present. DSM (72x72x72 mesh)                         0.4            --.-    Present. DSM (72x7 02 _
Present, LDKEM (36x36x36 mesh)
Present LDKEM (72x72x2 mesh)
DNS (Kim et al., 1987)                               0.2 3           ------ Present, LDKEM (3.
Prcsent LDKEM (7 o DNS (Kinet al.,.....
192 0               50              100             150           200                            0              100 Y                                                                          y (c) r.m.s. velocity (VI+) predictions                                      (d) r.m.s. velocity (m" ) prr edictions Figure 1. The results of LES using two dynamic SGS turbulence models for Ret = 180 9 of 17 American Institute of Aeronautics and Astronautics Paper 2004-2548
: 13. Flow around a square cylinder The flow past a square cylinder measured by Lyn et al.2 4 was considered. The Reynolds number based on the freestream velocity (Uo) and the width of the cylinder (H) is 22,000. The subject flow is featured by a massive flow separation accompanied by unsteady large-scale structures of widely varying length scales. As such, it aptly represents turbulent flows around bluff bodies with sharp edges. Some others have also tackled this flow using LES.13 25 The domain size and the mesh resolution were chosen in reference to the earlier studies by others. 1'325 A more comprehensive study using different mesh resolutions -andUoma-nsizes (san-wis m p-artic-ular) is deTerre-dfr-a-future study. Our objective here is to evaluate the efficacy of the present LES capability by comparing the predictions with other results based on meshes with comparable resolution. The computational domain is bounded by an upstream inlet boundary, top and bottom boundaries located 7.0 H from the center of the cylinder, and an exit boundary at 20 H.
from the cylinder axis. Freestream conditions were specified at the inlet. The top and the bottom boundaries were treated as symmetry planes (frictionless walls). The exit boundary was modeled as a pressure boundary where the solution variables are extrapolated in a mass-conserving manner. A pair of periodic boundaries separated with a span of 3.0H were used in the spanwise direction. The computational domain is filled with a hexahedral mesh with 660,000 elements. NVe took advantage of our unstructured mesh capability, embedding a block of locally refined mesh around the cylinder to better resolve the near-wall and wake regions as depicted inJgure 2 on the next page. The averaged wall-distance at the wall-adjacent cells is 0.012 11. The time-step size (At) used for the present computations is 0.02 time unit (H/Uo) which is comparable to that used by others.' 3 The CD scheme was used for the discretization of convective terms. The statistics were obtained during the LES for a sufficiently long period of time, typically for more than several scores of thne units.
Figure 2 on the following page shows the profiles of the mean axial velocity (U/Uo) and the r.m.s.velocity fluctua-tions (' and Vl)along the centerplane (y = 0) in the wake predicted using the two dynamic models. The time-averaged axial velocity distributions show that the length of the recirculation bubble behind the cylinder is predicted by the two dynamic models to be around L, = 0.9 which agrees remarkably well with the experimental value (L, ; 0.9). The negative peak and the recovery of the mean velocity in the near-wake are also captured very closely. However, the predictions start to deviate from the measurement forx/H > 2.0, reaching 0.8 Uo asymptotically in the far-wake. As shown in the figure, this value is considerably larger than what the measurement indicates (0.62Uo).24 Interestingly, others who computed the same flowt 3' 25 also have grossly overpredicted the recovery of the axial mean velocity. Our LES predictions of the mean axial velocity in the far-wake with both dynamic models were found to be largely com-parable to the prediction by Sohankar et al.13 based on their dynamic one-equation model denoted by "OEDSMA" in their paper. However, the present LES predictions reproduce the mean velocity profile in the recirculation bubble more accurately than others. Particularly noteworthy is that our DSM yields somehow a much better prediction than the DSM model used by Sohankar et al. in terms of the recirculation bubble size and the asymptotic value of the axial mean velocity in the far-wakc, which begs a question of what could possibly contribute to this sizable difference.
One possible cause is the effectively finer mesh used in the present computations which was made possible by the embedded region of fine mesh around the cylinder. It is also quite likely that the differences in the details of the DSM implementation is responsible. In this regard, it should be noted that Sohankar el aL average Cv in the spanwise (homogeneous) direction, whereas the DSM used here does not.
The r.m.s.fluctuating velocity components are also predicted with a reasonable accuracy by the present LES. The peak values are appreciably underpredicted. However, the locations of the peaks are closely captured. Another obser-vation worthy of mentioning is that V is relatively poorly predicted by both dynamic models.
Table I on page 12 summarizes other global parameters predicted by the present computations, along with the results predicted by others.13 '25 Our LES predictions of the mean drag coefficient, Strouhal number, and r.m.s.lift coefficient well match the measurements and the predictions by others.
lo of 17 American Institute of Aeronautics and Astronautics Raper 2004-2548

1.0 0.4 0.2                -      LES with DSNI (present)
                                                                                                      ----- LES with LDKEMI (present) 0.0         = =           o         ettal. (I 995)
HILM
_O             2_0    _      40             6.0       8.0   10.0 X1l1 (a) Mesh with an embedded fine nmesh                                              (b) Axial mean velocity 00.3 -                    000                                                0.6-0.2-04
                              -LES    with DSMI (present)
LES with LDKEM (present)                                              -       LES with DSM (present) 0.1                      o    Lyn et al. (1995) t------                     0.2                             LES with LDKEM (present)     -
Ia       Lyn et al. (1995)
_ ___-                                                                                               I               I         II
                    ]!J.0            5.0                            10.0                     2.0            4.0            6.0       8.0   10.0 (c) rm.s. axial fluctuating velocity                                      (d) rm.s. vertical fluctuating velocity Figure 2. Mesh used for the flow around a square cylinder and the rnean axial velocity and normal stress distributions along the wake centerline Il of 17 American Institute of Aeronautics and Astronautics Paper 2004-2548

Table 1. Summary of the LES predictions of the global quantities for the square-cylinder case (Reif = 22,000)
Methods                                   L,            St           CD        CIL DSM (present)                             0.9           0.133         2.19     1.19 DSM (Sohankaretal.13 )              ~0.6                0.126         2.03     1.23 DSM (Furcby et aL 25)                     0.83          0.132         2.0      1.34 LDKEM (present)                           0.9           0.131         2.14     1.17 OEDSM (Sohankaretal.'3 )              :0.6              0.132         2.32     1.54 LDKEM (Fureby etal.25 )                   0.74          0.130         2.10     1.32 Measured (Lyn et aL.24)                   0.9           0.13      : 2.1        1.2 C. Flow around a sphere
: 1. Direct simulationforRCD = 300 Before tackling the turbulent flow cases, laminar flow at ReD = 300 was computed on a hybrid unstructured mesh using direct numerical simulation. At this Reynolds number, the flow exhibits a weak unsteadiness leading to oscillations in drag and lift forces. The hybrid mesh has a total of 860,000 cells, consisting of prismatic cells in the near-wall region grown from the surface triangles on the body surface and tetrahedral cells filling the rest of the solution domain. The time-step size of 0.04 D/U was used. Several others have computed this flow to validate their numerics.26- 2 8 Table 2. Summary or thc prediction for the laminar flow over asphere (kCD.!QQ09)
Methods                                 St                     CD Present                                 0.133                  0.667i Tomboulides e: al7-                     0.136                  0.67 1 Johnson andd';Iitc_26                     .0137__._ __         0.6561 Measured 2 9                            0.15 - 0.16            0.629:
The results are summarized in Table 2 along with others' predictions. 2 62 8 Tomboulides et aL27 and Johnson and Patel2 6 used, respectively, a high-order spectral element method and a second-order upwind finite difference scheme on high-quality structured meshes. As shown, our predictions agree well with others' results, which is remarkable considering that a hybrid unstructured mesh was used in this work in conjunction with the second-order discretization scheme.
2.Tu rbulentflows We considered two Reynolds numbers (ReD = 1.0 X 104, 1.14 x 106), one being in sub-critical and the other in super-critical regime. The subcritical flow case has been numerically studied by several others,2 8' 30 while the supercritical case was studied experimentally by Achenbach. 31 In this work, a hybrid unstructured mesh was deliberately used for both Re cases. The hybrid mesh has in total 2.46 million cells, consisting of 0.6 million prismatic cells in the near-wall region and tetrahedral cells filling the rest of the solution domain, with a large fraction of the total cell counts clustered in the near-wake region (see Figure 3).
The mesh quality is not exceptionally high and yet quite reasonable, except the rapid expansion of cell size around the 12 of 17 American Inslilute of Aeronautics and Astronatimics Paper 2004-2548

(a) Hybrid mesh consisting of prismatic and tetrahedral elements                     (b) Enlarged view of the near-wall mesh Figure 3. Hybrid unstructured mesh used for the flow around a sphere cell-clustered region. The average distance from the wall at the wall-adjacent cells is around 1.1 x 10-3 D. For the lower-Re case, the near-wall mesh is sufficiently fine to resolve the boundary layer, with they+ values at the wall cells below y+ = 1.0 for most part of the wall. For the higher-Re case, however, the near-wall mesh is far from being fine enough to accurately resolve the boundary layer which is much thinner than the lower-Re case, and the y+ values at the wall cells increase by almost two orders of magnitude. Thus, the wall adjacent cells are most likely to penetrate the fully turbulent region (log-layer) on a significant portion of the wall, especially near 0 = 90&deg; where the skin-friction reaches a maximum. The mesh being not ideal, the higher-Re case offers a good opportunity to assess the wall-function based approach adopted in this work. Partial views of the mesh are shown in Figure 3.
Attempts to use the pure CD scheme have not been successful for this case. Numerical oscillations were observed sporadically in a few spots rather remote from the body where the cell size increases rapidly, being accompanied by abnormally large velocity magnitude. Although the oscillations were not catastrophic and affected the global quantities very little, the subsequent computations were carried out using the BCD scheme discussed earlier. A time-step size of 0.02D/U was used for both Reynolds numbers. The data were collected for more than hundreds of time units.
Figure 4 shows the time-histories of CD for the lower-Re case recorded during the LES using the two dynamic models.
Table 3. Summary of the LES prediction for turbulent flow past a sphere (ReD = 10,000)
Methods                                       CD              St LES with DSM (present)                        0.438         0.182         86- 87        86 - 88 LES with LDKEM.(present)                      0.433        0.185          86- 87        86 - 88 DES (Pelacz et aL)30                          0.430            -              _            _
DES (Constantinescu et aL)28                  0.397        0.200          84 - 87     93 - 108 LES (Constantinescu e aL.)28                  0.393        0.195          84 - 86       86 - 88 Measured 3 1' 32                           A:0.40       0.185 - 0.19 Correlation33                                 0.46            -
13 of 17 American Institute of Aeronautics and Astronautics Paper 2004-2548

The global quantities predicted for the lower Re case are summarized in Table 3 on the page before along with others' results. The predicted CD values (0.438 and 0.433 for the DSM and LDKEM, respectively) are in fair agreement with the ofteen-quoted experimental value of 0.4 measured in 1920's 34 and other predictions. Constantinescu et al 28 predicted CD, at around 0.4 using LES and detached eddy simulation (DES) on a structured hexahedral mesh having 450,000 nodes. Our predictions are closer to the value obtained by Pelaez et al.30 (C) ; 0.43) who carried out a DES on an unstructured mesh with 770,000 nodes. The Strouhal numbers predicted by tie two dynamic models came out very close to each other, matching the measured one quite closely in view of the scatter in the experimental data.
The data of Achenbach 3t and Kim and Durbin35 favor lower Strouhal number around St = 0. 15, whereas Sakamoto's data 32 suggests a higher value between 0.18 and 0.19. The locations of flow separation (0,) predicted by the two dynamic models are nearly identical and were found in 860 870. The locations of laminar-lo-turbulent transition
(&#xa2;,), which were obtained in this work by reading off the angle beyond which v, or /g, increases rapidly, were found in 860 88&deg;for both dynamic models, which compare well with the LES predictions by Constantinescu e al28 103 0.s5           I                                    I t 1t51Hilil1 1 1Experimental .
102o                Io24 1!1ll llaLES with DSM LESwithLDKEM         i .
0.50                                                                                  l~~~~~~~                    :U~......
minar calc.i
                                                                                                                                 -.. 1 10
: 10.       1        X             I              1_1 II 1500       200       ' 30           400        500         600            t            i'               1.   )
I UJD                                                                        Reg)
(a)Time histories of CD) for Reg = 10.000                      (b) Comparison of the predicted time-averagcd CD values (ReD =
300,10,000,1.14 x I06) and the experimental mean CD-curve for a range of Reynolds number Figure 4. Time histories of drag coefficient (CD) and and the predicted Mean drag coeflicients for the sphere With the higher-Re case, the flow in reality has already undergone the "drag crisis", and the flow structure has changed drastically from those of subcritical regime. The change in the flow structure can be seen from Figure 5 and Figure 6. Depicted in these figures are the iso-contours of the second-invariant of the deformation tensor, (f5ijQij -
SijSj)/2. Both figures aptly portray the hairpin-like vortical structures in the wake observed in experiments. The wake for the higher-Re case (Figure 6) is much narrower than in Figure 5, which is the consequence of the delayed onset of flow separation. The CD values predicted by the DSM and the LDKEM are 0.139 and 0.142, respectively. These values are fairly close to the range of values (CD = 0.12 0.14) measured by Achcnbach. 3 1 For the higher-Re case, the predictions of the locations of the separation and the onset of transition were much less satisfactory. At Re = 1.14 x 106, the experimental results 3t show that the transition occurs near 970 98&deg; well before the separation occurring near 120&deg;. The present predictions failed to reproduce this experimental finding. The present results exhibit too early a separation at around 1000 and a delayed transition. This discrepancy is most likely due to the use of too coarse a mesh in this work to resolve the very thin boundary layer for the high-Re case.
The drag coefficients predicted in this study for the three Reynolds numbers are plotted in Figure 4 along with the mean experimental curve.
14 of 17 American Institute of Aeronautics and Astronautics Paper 2004-2548

Figure 5. Vortical structure in the near-wake of the sphere for ReD = 10,000 - visualized using the iso-contour or the second-invariant or the velocily, deormation tensor Figure 6. Vortical structure In the near-wake of the sphere forReD    =  1.14 x 106 - visualized using the Iso-contour of the second-invariant of the velocity deformation lensor 15 of 17 American Institute of Aeronautics and Astronautics Riaper 2004-2548

VI.      Summary and Conclusion A large eddy simulation capability based on a finite-volume solver has been developed and validated for a num-ber of wall-bounded flows. The finite-volume solver employs second-order numerics and permits use of unstructured meshes, thus being able to easily handle industrial applications involving complex geometry. Turbulence closure for subgrid-sale stresses is effected using two dynamic subgrid-scale viscosity models, namely, the dynamic Smagorin-sky model (DSM) and the localized dynamic k-equation model (LDKEM). These two dynamic models allow one to compute arbitrary three-dimensional flows without any statistically homogeneous directions. The validations demon-strated that the present LES capability is capable of predicting the wall-bounded flows of varying complexity with a commendable accuracy, having a potential to provide a practical tool for high-level simulation of turbulent flows encountered in industrial applications.
VII.       Acknowledgments The author gratefully acknowledges use of Fluent Inc.'s software and thanks the members Of the development group at Fluent Inc. Special thanks go to Sunil Vijay Unaune at the Fluent India who contributed the mesh used for the computation of the flow past a sphere.
References t
Haworth, D. C. and Jansen, K.. "Large-Eddy Simulation on Unstructured Deforming Meshes: Towards reciprocating IC Engines," Computers Fluids. Vol. 29,2000. pp. 493-524.
2K. Mahesh, G. Constantinescu, P.M., "Large Eddy Simulation of Gas Turbine Conibustors:' Annual Reserach Briefs, Center for Turbulence Research, NASA Ames/Stanford University, 2000, pp. 219-229.
3 Urbin, G. and Knight, D.. "Large Eddy Simulation of a Supersonic Boundary Layer Using an Unstructured Grid' AIAA Journal, Vol. 39, No. 7, July 2001, pp. 927-935.
4 K. Jansen, T. M. and Reba, R., "Finite-Elemnent Based Largc-Eddy Simulation of the Near-Nozzle Region of a Compressible Round Jet:'
Aiaa paper 2002-2358, 2002.
5S. Benhamadouche, K. M. and Constantinescu, G., "Colocated Finite-Volume Schemes for Large Eddy Simulation on Unstructured Grids:'
Proceedings of the Summer Program,Center for Turbulence Research, NASA Aames/Stanford University, 2002, pp. 143-154.
6 M. Germano, U. Piomelli, P. M. and Cabot, W. II., "Dynamic Subgrid Scale Eddy Viscosity Model:' Physics of Fluids A. Vol. 3, 1991, pp. 1760-1765.
7 Lilly. D. K., "A Proposed Modification of the Germano Subgrid Scale Closure Method:' Physics of FluidsA, Vol. 4, No. 3,1992, pp. 633-635.
SKim, W.-W. and Menon, S., "A New Dynamic One-Equation Sub-Grid Scale Model for Large Eddy Simulations:' Aima paper 95-0356.
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9 Kim, W.-W. and Menon, S., "Application of the Localized Dynamic Subgrid.Scale Model to Turbulent Wall-Bounded Flows," Aiaa paper 97-0210, 1997.
1'Smagorinsky, J., 'General Circulation Experiments with the Primitive Equations, part I The Basic Experiment," Mlonthly Weatlher Review, Vol. 91, 1963, pp. 99-165.
1IS. Ghosal, T. Lund, P. M. and Akselvoll, K., "A Dynamic Localization Model for Large Eddy Simulation of Turbulent Flows:' Journalof FluidMechanics, Vol. 286, 1995, pp. 2 2 9 - 2 5 5 .
12 Davidson, L., "Large Eddy Simulation: A Dynamic One-Equatio Subgrid Model for Three Dimensional Recirculating Flow:' Proceedings of the 11th InternationalSymposium Turbulent Shear Flows, ol. 3,1997, pp. 26.1-26.6.
13A. Sohankar, L. D. and Norberg, C., 'Large Eddy Simulation of Flow Past a Square Cylinder Comparison of Different Subgrid Scale Models:' Journalof Fluids Engineering, Vol. 122, No. 1.2000, pp. 376-404.
14 Krajnovic, S. and Davidson, L., "Large Eddy Simulation or the Flow Around a Bluff Body:' AIAA Journal, Vol. 40, No. 5, May 2002, pp. 927-935.
15S. Liu, C. M. and Katz, J., "On the Properties of Similarity SUbgrod-Scale Models as Deduced from Measurements in a Turbulent Jet,'
Journalof Fluid Mfechanics, Vol. 275, 1994, pp. 83-119.
16S. Liu, C M. and Katz, J., Experimental Study of Similarity Subgrid-Scale Models of Turbulence in the Far-Fieldofa Jet, Direct and Large Eddy Simnulation, Kluwer, 1994.
t7 Kader, B. A., 'Temperature and Concentration Profiles in Fully Turbulent Boundary Layers:' Int. J. Heat fass Transfer, Vol. 24, No. 9, 1981,pp. 1541-1544.
      ' t Mathur, S. R. and Murthy.J. Y., "A Pressure-Based Method for Unstructured Meshes:' NumericalHeat Transfer,Vol. 31,1997. pp. 195-215.
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      "9 S.-E. Kim, S. R. Mathur. J. Y. M. and Choudhury, D., "A Reynolds-Averaged Navier-Stokes Solver Using Unstructured Mesh-Based Finite-Volume Scheme.' Aiaa paper 98-0231. 1998.
2tS.-E. Kim, B. MI.and Caraeni, D., "A Multidimiensional L.inear Reconstnrction Scheme for Arbitrary Unstructured Grids:' Aiaa paper 2003-3990, Jan. 2003.
21L. Davidson, D. Cokijat, J. F. M. A. L. C. KI.W. R.. "LESFOIL: Large Eddy Simulation of Flow Around a High Lift Airfoil:' Afores on Numerical Fluid lechanics and Alutidisciplinary Design - I'vlunie 83, Springer, 2002.
22 J. Kim, P.M. and NMoser, R.,"Turbulent Statistics in Fully Developed Channel Flow at Low Reynolds Number," Journal of Fluid lMechanics, Vol. 177,1987, pp. 133-166.
23D.Ch!opiD._Prasad, M. W. and Pierce, C., "Evaluation of an Industrial CFD code for LES ApplicationsProceed/ingsof rhie 2000 Suinijer Program. Center for Turbulence Research, NASA AamneslStanford University, 2000, pp. 221-228.
24 D. A. Lyn, S. Einav, W. R. and Park, J. H., "A Laser Doppler Velocitimetry Study of Ensemble Averaged Characteristics of the Turbulent Near-Wake of a Square Cylinder," Journal of Fluid Alechanics, Vol. 304, 1995, pp. 285-319.
25C. Fureby, G. Tabor.      G1.
G. W. and Gosman. A. D.. "Large Eddy Simulation of the Flow Around a Square Prism' AMAA Journal. Vol. 38.
No. 3, May 2000, pp. 442-452.
26 johnson, T. A. and Patel, V. C., "Flow Past a Sphere up to a Reynolds Number of 300:' Journal of Fluid Mechanics, Vol. 378, 1999, pp. 19-70.
27A. G. Tomboulides, S. A. Orszag. G. A. K., "Direct and Large-Eddy Simulations of Axisymmetric Wakes:' Aiaa paper 93-0546, Jan. 1993.
28G. Constantinescu, NM. C. and Squires, K., "Turbulence Modeling Applied to Flow over a Sphere' AIAA Journal, Vol. 41, No. 9, 2003, pp. 1733-1742.
     "Roos, F.W. and Williams, W. W., "Some Experimental Results on Sphere and Disk Drag.: AIAA Journal, Vol. 9, No.2, 1971. pp. 285-291.
30J. Pelaez, D. J. M. and Kandil, O., "Unsteady Analysis of Separated Aerodynamic Flows Using an Unstructured Multigrid Algorithm:' Aiaa paper200108gThr20 3
tAchenbach E., "Experiments on the Flow past Spheres at Very Iligh Reynolds Numbers:'Journal of FluidAechanics, Vol. 54, No.3. 1972, pp. 565-575.
32Sakamoto, H. and lianiu, 11.,"A Study on Vortex Shedding from Spheres in an Unifonm Flow:' Journal of Fluids Fngineering. Vol. 112, 1990. pp. 386-393.
33 White, F. M., Viscous Fluid Flow, McGraw-Hill, 1974.
34Schlichting, H., Boundary-Layer Theory, NMcGraw-llill, 1979.
35Kim, K. J. and Durbin, P.A., "Observation of the Frequencies in a Sphere Wake and Drag Increase by Acoustic Excitation:' Physics of Fluids, Vol. 31, No. 11, 1988, pp. 3260-3265.
17 of 17 American Institute of Aeronautics and Astronautics Paper 2004-2548

BVY 05-072 Docket No. 50-271 Exhibit SRXB-A-1 8-1 Vermont Yankee Nuclear Power Station Proposed Technical Specification Change No. 263 - Supplement No. 30 Extended Power Uprate Response to Request for Additional Information "RS-001, BWR Template SE" l Total number of pages In this Exhibit I (excludina this cover sheet) Is 8.                    I

(2) draft GDC-9, insofar as it requires that the reactor coolant pressure boundary shall be designed and constructed so as to have an exceedingly low probability of gross rupture or significant leakage throughout its design lifetime; 2.8.5 Accident and Transient Analyses 2.8.5.1    Decrease In Feedwater Temperature, Increase in Feedwater Flow, Increase In Steam Flow, and Inadvertent Opening of a Main Steam Relief or Safety Valve Reaulatorv Evaluation Excessive heat removal causes a decrease In moderator temperature which Increases core reactivity and can lead to a power level Increase and a decrease In shutdown margin. Any unplanned power level Increase may result Infuel damage or excessive reactor system pressure. Reactor protection and safety systems are actuated to mitigate the transient. The NRC staff's review covered (1) postulated Initial core and reactor conditions, (2)methods of thermal and hydraulic analyses, (3) the sequence of events, (4) assumed reactions of reactor system components, (5)functional and operational characteristics of the reactor protection system, (6) operator actions, and (7)the results of the transient analyses. The NRC's acceptance criteria are based on (1)draft GDC-6, insofar as It requires that the reactor core be
\desi nezto function throughout its design lifetime without exceeding acceptable fuel damage imits,(Ydraft GDC-14 and 15, Insofar as they require that the core protection system be designed to act automatically to prevent or suppress conditions that could result In exceeding acceptable fuel damage limits and that protection systems be pr*,ided for sensing accident situations and Initiating the operation of necessary ESFs; and (X) draft GDC-27 and 28, Insofar as they require that at least two reactivity control systems be provided and be capable of making and holding the core subcritical from any hot standby or hot operating condition sufficiently fast to prevent exceeding acceptable fuel damage limits. Specific review criteria are contained In SRP Section 15.1.1-4 and other guidance provided In Matrix 8 of RS-001.
Technical Evaluation
[insert technical evaluation. The technical evaluation should (1)clearly explain why the proposed changes satisfy each of the requirements In the regulatory evaluation and (2) provide a clear link to the conclusions reached by the NRC staff, as documented In the conclusion section.]
Conclusion The NRC staff has reviewed the licensee's analyses of the excess heat removal events described above and concludes that the licensee's analyses have adequately accounted for operation of the plant at the proposed power level and were performed using acceptable analytical models. The NRC staff further concludes that the licensee has demonstrated that the reactor protection and safety systems will continue to ensure that the               s nd the RCPB pressure limits will not be exceeded as a result of these events. Based on tis, the NRC sta               Hi concludes that the plant will continue to meet the requirements of draft GDC6,A4, 5, 27,           and 28 following implementation of the proposed EPU. Therefore, the NRC staff finds the proposed EPU acceptable with respect to the events stated.
INSERT 8 FOR SECTION 32. BWR TEMPLATE SAFETY EVALUATION DECEMBER 2003

(2) draft GDC-9, insofar as it require$ that the reactor coolant pressure boundary shall be designed and constructed so as to have an exceedingly low probability of gross rupture or significant leakage throughout its design lifetime; 2.8.5.2 Decrease In Heat Removal by the Secondary System 2.8.5.2.1     Loss of External Load; Turbine Trip; Loss of Condenser Vacuum; C,losure of Main Steam Isolation Valve; and Steam Pressure Regulator Failure (Closed)
Reoulatory Evaluation A number of initiating events may result in unplanned decreases In heat remov val by the secondary system. These events result in a sudden reduction In steam flow arid, consequently, result In pressurization events. Reactor protection and safety systems are actt Wated to mitigate the transient. The NRC staffs review covered the sequence of events, the anmdlytical models used for analyses, the values of parameters used in the analytical models, and the results of the transient analyses. The NRC's acceptance criteria are based on (1) draft GDO :6, insofar as it requires that the reactor core be desIgned to funojon throughout Its design lifeltime without exceeding acceptable fuel damage lirni, and (tYdraft GDC-27 and 28, Insofar as they require that at least two reactivity control systems be provided and be capable of making and holding the core subcritical from any hot standby or hot operating condition sufficiently fast to prevent exceeding acceptable fuel damage limits. Specific review criteria are contained in SRP Section 15.2.1-5 and other guidance provided in Matrix 8 of RS-001.
Technical Evaluation
[insert technical evaluation. The technical evaluation should (1) clearly explain why the proposed changes satisfy each of the requirements In the regulatory evaluation and (2) provide a clear link to the conclusions reached by the NRC staff, as documented in the conclusion section.]
Conclusion The NRC staff has reviewed the licensee's analyses of the'decrease in heat removal events described above and concludes that the licensee's analyses have adequately accounited for operation of the plant at the proposed power level and were performed using acceptable analytical models. The NRC staff further concludes that the licensee has demonstrated that the reactor protection and safety systems will continue to ensure that the(KAMb F9)and the RCPB pressure limits will not be exceeded as a result of these events. Based on -Ws, the NRC sta               q concludes that the plant will continue to meet the requirements of draft GDC-6, 27, and 28 following Implementation of the proposed EPU. Therefore, the NRC staff finds the proposed EPU acceptable with respect to the events stated.
INSERT 8 FOR SECTION 3.2- BWR TEMPLATE SAFETY EVALUATION DECEMBER 2003

(2) draft GDC-9, insofar as it requires that the reactor coolant pressure boundary shall be designed and constructed so as to have an exceedingly low probability of gross rupture or significant leakage throughout its design lifetime; 2.8.5.2.2 Loss of Nonemergency AC Power to the Station Auxiliaries Reaulatory Evaluation The loss of nonemergency ac power is assumed to result in the loss of all power to the station auxiliaries and the simultaneous tripping of all reactor coolant circulation pumps. This causes a flow coastdown as well as a decrease In heat removal by the secondary system, a turbine trip, an Increase In pressure and temperature of the coolant, and a reactor trip. Reactor protection and safety systems are actuated to mitigate the transient. The NRC staffs review covered (1) the sequence of events, (2) the analytical model used for analyses, (3) the values of parameters used In the analytical model, and (4) the results of the transient analyses. The NR~C~'s acceptance criteria are based on (1) draft GDC-6, insofar as it requires that the reactor core bedesined to function throughout its design lifetime without exceeding acceptable fuel amage7limitsand (draft GDC-27 and 28, Insofar as they require that at least two reactivity control systems be provided and be capable of making and holding the core subcritical from any hot standby or hot operating condition sufficiently fast to prevent exceeding acceptable fuel damage limits. Specific review criteria are contained in SRP Section 15.2.6 and other guidance provided in Matrix 8 of RS-O001.
Technical Evaluation
[Insert technical evaluation. The technical evaluation should (1) clearly explain why the proposed changes satisfy each of the requirements in the regulatory evaluation and (2) provide a clear link to the conclusions reached by the NRC staff, as documented In the conclusion section.]
Conclusion The NRC staff has reviewed the licensee's analyses of the loss of nonemergency ac power to station auxiliaries event and concludes that the licensee's analyses have adequately accounted for operation of the plant at the proposed power level and were performed using acceptable analytical models. The NRC staff furtherconcludes that the licensee hs.Lemonstrated that the reactor protection and safety systems will continue to ensure that thelAFiand the RCPB pressure limits will not be exceeded as a result of this event. Based on this, the NRC staff               9 concludes that the plant will continue to meet the requirements of draft GDC-6, , and 28 following Implementation of the proposed EPU. Therefore, the NRC staff finds the proposed EPU acceptable with respect to the loss of nonemergency ac power to station auxiliaries event INSERT 8 FOR SECTION 3.2 - BWR TEMPLATE SAFETY EVALUATION DECEMBER 2003

(2) draft GDC-9, insofar as it requires that the reactor coolant pressure boundary shall be designed and constructed so as to have an exceedingly low probability of gross rupture or significant leakage throughout its design lifetime; 2.8.5.2.3 Loss of Normal Feedwater Flow Regulatory Evaluation A loss of normal feedwater flow could occur from pump failures, valve malfunctions, or a LOOP.
Loss of feedwater flow results Inan increase In reactor coolant temperature and pressure which eventually requires a reactor trip to prevent fuel damage. Decay heat must be transferred from fuel following a loss of normal feedwater flow. Reactor protection and safety systems are actuated to provide this function and mitigate other aspects of the transient. The NRC staffs review covered (1)the sequence of events, (2) the analytical model used for analyses, (3) the values of parameters used In the analytical model, and (4)the results of the transient analyses.
The NRC's acceptance criteria are based on (1)draft GDC-6, Insofar as it requires that the reactoro         signed tQ function throughout its design lifetime without exceeding acceptable fuel damage limitsand (4Y'draft GDC-27 and 28, insofar as they require that at least two reactivity control systems be provided and be capable of making and holding the core subcritical from any hot standby or hot operating condition sufficiently fast to prevent exceeding acceptable fuel damage limits. Specific review criteria are contained In SRP Section 15.2.7 and other guidance provided In Matrix 8 of RS-001.
Technical Evaluation lnsert technical evaluation. The technical evaluation should (1) clearly explain why the proposed changes satisfy each of the requirements In the regulatory evaluation and (2) provide a clear link to the conclusions reached by the NRC staff, as documented In the conclusion section.]
Conclusion The NRC staff has reviewed the licensee's analyses of the loss of normal feedwater flow event and concludes that the licensee's analyses have adequately accounted for operation of the plant at the proposed power level and were performed using acceptable analytical models. The NRC staff further concludes that the licensee has demonstrated that the reactor protection and safety systems will continue to ensure that the(tDs and the RCPB pressure limits will not be exceeded as a result of the loss of norrneecwater flow. Based on this, the NRC staff                       9 concludes that the plant will continue to meet the requirements of draft GDC-6, 7, and 28 following implementation of the proposed EPU. Therefore, the NRC staff finds the proposed EPU acceptable with respect to the loss of normal feedwater flow event.
INSERT BFOR SECTION 3.2 - BWR TEMPLATE SAFETY EVALUATION DECEMBER2003

(2) draft GDC-9, insofar as it requires that the reactor coolant pressure boundary shall be designed and constructed so as to have an exceedingly low probability of gross rupture or significant leakage throughout its design lifetime; 2.8.5.3 Decrease In Reactor Coolant System Flow 2.8.5.3.1 Loss of Forced Reactor Coolant Flow Reaulatorv Evaluation A decrease in reactor coolant flow occurring while the plant Is at power could result In a degradation of core heat transfer. An increase in fuel temperature and accompanying fuel damage could then result I            are exceeded during the transient. Reactor protection and safety systems are actuated to mitigate the transient. The NRC staff's review covered (1)the postulated initial core and reactor conditions, (2)the methods of thermal and hydraulic analyses, (3)the sequence of events, (4) assumed reactions of reactor systems components, (5)the functional and operational characteristics of the reactor protection system, (6) operator actions, and (7) the results of the transient analyses. The NRC's acceptance criteria are based on (1) draft GDC-6, Insofar as it requires that the reactor core be designed to function throughout its design lifetime without exceeding acceptable fuel damage limitswand (;)draft GDC-27 and 28, Insofar as they require that at least two reactivity control systems be provided and be capable of making and holding the core subcritical from any hot standby or hot operating condition sufficiently fast to prevent exceeding acceptable fuel damage limits. Specific review criteria are contained In SRP Section 15.3.1-2 and other guidance provided In Matrix 8 of RS-001.
Technical Evaluation
[Insert technical evaluation. The technical evaluation should (1) clearly explain why the proposed changes satisfy each of the requirements In the regulatory evaluation and (2) provide a clear link to the conclusions reached by the NRC staff, as documented In the conclusion section.]
Conclusion The NRC staff has reviewed the licensee's analyses of the decrease In reactor coolant flow event and concludes that the licensee's analyses have adequately accounted for operation of the plant at the proposed power level and were performed using acceptable analytical models.
The NRC staff further concludes that the licensee has demonstrated that the reactor protection and safety systems will continue to ensure that the(~ and the RCPB pressure limits will not be exceeded as a result of this event. Based on this,the NRC staff concludes that the plant         q will continue to meet the requirements of draft GDC-6V27, and 288tollowing implementation of the proposed EPU. Therefore, the NRC staff finds the proposed EPU acceptable with respect to the decrease in reactor coolant flow event.
INSERT8 FOR SECTION 3.2 -BWRTEMPLATE SAFETY EVALUATION DECEMBER 2003

I (2)draft GDC-9, insofar as it requires that the reactor coolant pressure boundary shall be designed and constructed so as to have an exceedingly low probability of gross rupture or significant leakage throughout its design lifetime; 2;8.5.5   Inadvertent Operation of ECCS or Malfunction that Increases Reactor Coolant Inventory Regulatory Evaluation Equipment malfunctions, operator errors, and abnormal occurrences could cause unplanned Increases Inreactor coolant Inventory. Depending on the temperature of the Injected water and the response of the automatic control systems, a power level increase may result and, without adequate controls, could lead to fuel damage or ovdrpressurization of the RCS. Alternatively, a power level decrease and depressurization may result. Reactor protection and safety systems are actuated to mitigate these events. The NRC staffs review covered (1)the sequence of events, (a)the analytical model used for analyses, (3)the values of parameters used Inthe analytical model, and (4)the results of the transient analyses. The NRC's acceptance criteria re based on (1) draft GDC-T,insofar as it re uire            the reactor core be designed t function throughout its design ietime without exce ding acceptable fuel damage limitsand (%Ndraft GDC-27 and 28, Insofar as they require that at least two reactivity control systems be provided and be capable of making and holding the core subcritical from any hot standby or hot operating condition sufficiently fast to prevent exceeding acceptable fuel damage limits. Specific review criteria are contained in SRP Section 15.5.1-2 and other guidance provided in Matrix 8 of RS-001.
Technical Evaluation
[Insert technical evaluation. The technical evaluation should (1) clearly explain why the proposed changes satisfy each of the requirements in the regulatory evaluation and (2) provide a clear link to the conclusions reached by the NRC staff, as documented in the conclusion section.]
Conclusion The NRC staff has reviewed the licensee's analyses of the Inadvertent operation of ECCS or malfunction that Increases reactor coolant Inventory and concludes that the licensee's analyses have adequately accounted for operation of the plant at the proposed power level and were performed using acceptable analytical models. The NRC staff further concludes that the licensee s demonstrated that the reactor protection and safety systems will continue to ensure that theAFD nd the RCP pressure limits will not be exceeded as a result of this event.                     +
Based on is, the NRC staff concludes that the plant will continue to meet the requirements of draft GDC                                                                                 re, the R staff finds the proposed EPU acceptable with respect to the Inadvertent operation of ECCS or malfunction that increases reactor coolant inventory.
INSERT 8 FOR SECTION 3.2 - BWR TEMPLATE SAFETY EVALUATION DECEMBER 2003

(2) draft GDC-9, insofar as It requires that the reactor coolant pressure boundary shall be designed and constructed so as to have an exceedingly low probability of gross rupture or significant leakage throughout its design lifetime; 2.8.5.4.3   Startup of a Recirculation Loop at an Incorrect Temperature and Flow Controller Malfunction Causing an Increase In Core Flow Rate Regulatory Evaluation A startup of an Inactive loop transient may result in either an Increased core flow or the introduction of cooler water into the core. This event causes an Increase In core reactivity due to decreased moderator temperature and core void fraction. The NRC staffs review covered (1)the sequence of events, (2)the analytical model, (3)the values of parameters used Inthe analytical model, and (4)the results of the transient analyses. The NRC's acceptance criteria arbased on (1 draft GDC-6, insof a require that the reactor core be desigr ed to function throughout its design lifetime without exceeding accepa e el damage limits,(2draft GDC-14 and 15, Insofar as they require that the core protection systems be designed to act automatically to prevent or suppress conditions that could result In exceeding acceptable fuel damage limits and that protection systems be provided for sensing accident situations and Initiating the operation of necessary ESFs; Updraft GPC-32, Insofar as It requires that limits, which Include considerable margin, be placed on the maximum reactivity worth of control rods or elements and on rates at which reactivity can be Increased to ensure that the potential effects of a sudden or large change of reactivity cannot (a)rupture the reactor coolant pressure boundary or (b) disrupt the core, its support structures, or other vessel internals sufficiently to Impair the effectiveness of emergency core cooling; and (Wfdraft GDC-27 and 28, Insofar as they require that at least two reactivity control systems be provided and be capable of making and holding the core subcritical from any hot standby or hot operating condition sufficiently fast to prevent exceeding acceptable fuel damage limits. Specific review criteria are contained In SRP Section 15.4.4-5 and other guidance provided In Matrix 8 of RS-001.
Technical Evaluation
[Insert technical evaluation. The technical evaluation should (1) clearly explain why the proposed changes satisfy each of the requirements In the regulatory evaluation and (2) provide a clear link to the conclusions reached by the NRC staff, as documented In the conclusion section.]
Conclusion The NRC staff has reviewed the Scensee's analyses of the increase in core flow event and concludes that the licensee's analyses have adequately accounted for operation of the plant at the proposed power level and were performed using acceptable analytical models. The NRC staff further concludes that the licensee has demonstrated that the reactor protection and safety systems will continue to ensure that                  Dand the RCPB pressure limits will not be exceeded as a result of this event. Based on this, he NRC staff concludes that the plant will continue to meet the requirements of draft GDC b,'14, 15, 27, 28, and 32 following Implementation of the proposed EPU. Therefore, the NRC staff finds the proposed EPU acceptable with respect to the increase in core flow event.
INSERT 8 FOR SECTION 3.2 - BWR TEMPLATE SAFETY EVALUATION DECEMBER 2003

(2)draft GDC-9, Insofar as it requires that the reactor coolant pressure boundary shall be designed and constructed so as to have an
              -exceedingly low probability of gross rupture or significant leakage throughout its design lifetime;
   .8.5.6 Decrease In Reactor C6olant Inventory 2.8.5.6.1 Inadvertent Opening of a Pressure Relief Valve Regulatorv Evaluation The Inadvertent opening of a pressure relief valve results In a reactor coolant Inventory decrease and a decrease in RCS pressure. The pressure relief valve discharges into the suppression pool. Normally there Is no reactor trip. The pressure regulator senses the RCS pressure decrease and partially closes the turbine control valves (TCVs) to stabilize the reactor at a lower pressure. The reactor power settles out at nearly the Initial power level. The coolant Inventory is maintained by the feedwater control system using water from the condensate storage tank via the condenser hotwell. The NRC staffs review covered (1)the sequence of events, (2)the analytical model used for analyses, (3) the values of parameters used Inthe analytical model, and (4) the results of the transient analyses. The NRC's acceptance criteria are based on (1) ft GDC-6 Insof as           uires tha              r e be designed to function throughout Its desig eimeswithout exceeding acceptable fuel damageimitand (kpdraft GDC-27 and 28, insofar as they require that at least two reactivity control systems be provided and be capable of making and holding the core subcritical from any hot standby or hot operating condition sufficiently fast to prevent exceeding acceptable fuel damage limits. Specific review criteria are     I contained in SRP Section 15.6.1 and other guidance provided In Matrix 8 of RS-001.
Technical Evaluation
[insert technical evaluation. The technical evaluation should (1) clearly explain why the proposed changes satisfy each of the requirements In the regulatory evaluation and (2) provide a clear link to the conclusions reached by the NRC staff, as documented In the conclusion section.]
Conclusion The NRC staff has reviewed the licensee's analyses of the inadvertent opening of a pressure relief valve event and concludes that the licensee's analyses have adequately accounted for operation of the plant at the proposed power level and were performed using acceptable analytical models. The NRC staff further concludes that the licensee has demonstrated that the reactor protection and safety systems will continue to ensure that the(A9FDLI-and the RCPB pressure limits will not be exceeded as a result of this event. Based on Ris-The NRC staff concludes that the plant will continue to meet the requirements, of draft GD C6in                   u.
following implementation of the proposed EPU. Therefore, the NRC staff finds the proposed EPU acceptable with respect to the inadvertent opening of a pressure relief valve event.
INSERT8 FORSECTION 3.2 BWRTEMPLATESAFETY EVALUATION DECEMBER 2003

BVY 05-072 Docket No. 50-271 Attachment 12 Vermont Yankee Nuclear Power Station Proposed Technical Specification Change No. 263 - Supplement No. 30 Extended Power Uprate Response to Request for Additional Information GE Affidavits for Attachment 9 Total number of pages in Attachment 12 (excluding this cover sheet) is 6.

General Electric Company AFFIDAVIT I, Robert E. Gamble, state as follows:
(1) I am Manager, ESBWR, General Electric Company ("GE") and have been delegated the function of reviewing the information described in paragraph (2) which is sought to be withheld, and have been authorized to apply for its withholding.
(2) The information sought to be withheld is contained in Enclosure 2 to GE letter GE-VYNPS-AEP-395, Michael Dick (GE) to Craig Nichols (ENOI), Responses to NRC RAIs SPSB-51 and SRXB-7, 8, 9, 10, 13, 14, 15, 16, 17, 19, 20, 21, 22, and 23, dated July 29, 2005. The Enclosure 2 proprietary information, Responses to NRC RAIs SPSB-51 and SRXB-7, 8, 9, 10, 13, 14, 15, 16, 17, 19, 20, 21, 22, and 23, is delineated by a double underline inside double square brackets. The header of each page in Attachment 2 carries the notation "GE Proprietary Information. (3)." In each case, the superscript notationt refers to Paragrph (3) of this affidavit, which provides the basis for the proprietary determination.
(3) In making this application for withholding of proprietary information of which it is the owner, GE relies upon the exemption from disclosure set forth in the Freedom of Information Act ("FOIA"), 5 USC Sec. 552(b)(4), and the Trade Secrets Act, 18 USC Sec.
1905, and NRC regulations 10 CFR 9.17(a)(4), and 2.390(a)(4) for "trade secrets" (Exemption 4). The material for which exemption from disclosure is here sought also qualify under the narrower definition of "trade secret", within the meanings assigned to those terms for purposes of FOIA Exemption 4 in, respectively, Critical Mass Energy Proiect v. Nuclear Regulatory Commission. 975F2d87l (DC Cir. 1992), and Public Citizen Health Research Group v. FDA, 704F2d1280 (DC Cir. 1983).
(4) Some examples of categories of information which fit into the definition of proprietary information are:
: a. Information that discloses a process, method, or apparatus, including supporting data and analyses, where prevention of its use by General Electric's competitors without license from General Electric constitutes a competitive economic advantage over other companies;
: b. Information which, if used by a competitor, would reduce his expenditure of resources or improve his competitive position in the design, manufacture, shipment, installation, assurance of quality, or licensing of a similar product;
: c. Information which reveals aspects of past, present, or future General Electric customer-funded development plans and programs, resulting in potential products to General Electric;
: d. Information which discloses patentable subject matter for which it may be desirable to obtain patent protection.
GBS-04-12-AFAf GE-VY-AEP-395.docAffidavit Page I

The information sought to be withheld is considered to be proprietary for the reasons set forth in paragraphs (4)a., and (4)b, above.
(5) To address 10 CFR 2.390 (b) (4), the information sought to be withheld is being submitted to NRC in confidence. The information is of a sort customarily held in confidence by GE, and is in fact so held. The information sought to be withheld has, to the best of my knowledge and belief, consistently been held in confidence by GE, no public disclosure has been made, and it is not available in public sources. All disclosures to third parties including any required transmittals to NRC, have been made, or must be made, pursuant to regulatory provisions or proprietary agreements which provide for maintenance of the information in confidence. Its initial designation as proprietary information, and the subsequent steps taken to prevent its unauthorized disclosure, are as set forth in paragraphs (6) and (7) following.
(6) Initial approval of proprietary treatment of a document is made by the manager of the originating component, the person most likely to be acquainted with the value and sensitivity of the information in relation to industry knowledge. Access to such documents within GE is limited on a "need to know" basis.
(7) The procedure for approval of external release of such a document typically requires review by the staff manager, project manager, principal scientist or other equivalent authority, by the manager of the cognizant marketing function (or his delegate), and by the Legal Operation, for technical content, competitive effect, and determination of the accuracy of the proprietary designation. Disclosures outside GE are limited to regulatory bodies, customers, and potential customers, and their agents, suppliers, and licensees, and others with a legitimate need for the information, and then only in accordance with appropriate regulatory provisions or proprietary agreements.
(8) The information identified in paragraph (2), above, is classified as proprietary because it contains detailed results and conclusions from evaluations of the safety-significant changes necessary to demonstrate the regulatory acceptability for the constant pressure power uprate for a GE BWR, utilizing analytical models and methods, including computer codes, which GE has developed, obtained NRC approval of, and applied to perform evaluations of transient and accident events in the GE Boiling Water Reactor ("BWR"). The development and approval of these system, component, and thermal hydraulic models and computer codes was achieved at a significant cost to GE, on the order of several million dollar.
(9) Public disclosure of the information sought to be withheld is likely to cause substantial harm to GE's competitive position and foreclose or reduce the availability of profit-making opportunities. The information is part of GE's comprehensive BWR safety and teclmology base, and its commercial value extends beyond the original development cost. The value of the technology base goes beyond the extensive physical database and analytical methodology and includes development of the expertise to determine and apply the appropriate evaluation process. In addition, the technology base includes the value derived from providing analyses done with NRC-approved methods.
The research, development, engineering, analytical and NRC review costs comprise a substantial investment of time and money by GE.
GBS-04-12-AFAf GE-VY-AEP-395.docAffidavit Page 2

The precise value of the expertise to devise an evaluation process and apply the correct analytical methodology is difficult to quantify, but it clearly is substantial.
GE's competitive advantage will be lost if its competitors are able to use the results of the GE experience to normalize or verify their own process or if they are able to claim an equivalent understanding by demonstrating that they can arrive at the same or similar conclusions.
The value of this information to GE would be lost if the information were disclosed to the public. Making such information available to competitors without their having been required to undertake a similar expenditure of resources would unfairly provide competitors with a windfall, and deprive GE of the opportunity to exercise its competitive advantage to seek an adequate return on its large investment in developing these very valuable analytical tools.
I declare under penalty of perjury that the foregoing affidavit and the matters stated therein are true and correct to the best of my knowledge, information, and belief.
Executed on this      9    day of    Thl'.                2005.
Robert E. Gamblc General Electric Company GBS-04-12-AFAf GE-VY-AEP-395.docAffidavit Page 3

General Electric Company AFFIDAVIT I, Robert E. Gamble, state as follows:
(1)      I am Manager, ESBWR, General Electric Company ("GE"), have been delegated the function of reviewing the information described in paragraph (2) which is sought to be withheld, and have been authorized to apply for its withholding.
(2)      The information sought to be withheld is contained in Enclosure 1 of GE letter, GE-VYNPS-AEP-396, Responses to NRC RAIs SRXB-24-58, dated August 1, 2005. The proprietary information in Enclosure 1, Responses to NRC RAIs SRXB-24-58 (Proprietary),
is delineated by a double underline inside double square brackets. Figures and large equation objects are identified with double square brackets before and after the object. In each case, the superscript notationt31 refers to Paragraph (3) of this affidavit, which provides the basis for the proprietary determination.
(3)      In making this application for withholding of proprietary information of which it is the owner, GE relies upon the exemption from disclosure set forth in the Freedom of Information Act ("FOIA"), 5 USC Sec. 552(b)(4), and the Trade Secrets Act, 18 USC Sec.
1905, and NRC regulations 10 CFR 9.17(a)(4), and 2.390(a)(4) for "trade secrets" (Exemption 4). The material for which exemption from disclosure is here sought also qualify under the narrower definition of "trade secret", within the meanings assigned to those terms for purposes of FOIA Exemption 4 in, respectively, Critical Mass Energy Project v. Nuclear Regulatorv Commission, 975F2d871 (DC Cir. 1992), and Public Citizen Health Research Group v. FDA, 704F2d1280 (DC Cir. 1983).
(4)      Some examples of categories of information which fit into the definition of proprietary information are:
: a.      Information that discloses a process, method, or apparatus, including supporting data and analyses, where prevention of its use by General Electric's competitors without license from General Electric constitutes a competitive economic advantage over other companies;
: b.      Information which, if used by a competitor, would reduce his expenditure of resources or improve his competitive position in the design, manufacture, shipment, installation, assurance of quality, or licensing of a similar product;
: c. Information which reveals aspects of past, present, or future General Electric customer-funded development plans and programs, resulting in potential products to General Electric;
: d.       Information which discloses patentable subject matter for which it may be desirable to obtain patent protection.
The information sought to be withheld is considered to be proprietary for the reasons set forth in paragraphs (4)a., and (4)b, above.
(5)     To address 10 CFR 2.390 (b) (4), the information sought to be withheld is being submitted to NRC in confidence. The information is of a sort customarily held in confidence by GE, and is in fact so held. The information sought to be withheld has, to the best of my knowledge and belief, consistently been held in confidence by GE, no public disclosure has been made, and it is not available in public sources. All disclosures to third parties including any required transmittals to NRC, have been made, or must be made, pursuant to regulatory provisions or proprietary agreements which provide for maintenance of the information in confidence. Its initial designation as proprietary information, and the subsequent steps taken to prevent its unauthorized disclosure, are as set forth in paragraphs (6) and (7) following.
(6)      Initial approval of proprietary treatment of a document is made by the manager of the originating component, the person most likely to be acquainted with the value and sensitivity of the information in relation to industry knowledge. Access to such documents within GE is limited on a "need to know" basis.
(7)     The procedure for approval of external release of such a document typically requires review by the staff manager, project manager, principal scientist or other equivalent authority, by the manager of the cognizant marketing function (or his delegate), and by the Legal Operation, for technical content, competitive effect, and determination of the accuracy of the proprietary designation. Disclosures outside GE are limited to regulatory bodies, customers, and potential customers, and their agents, suppliers, and licensees, and others with a legitimate need for the information, and then only in accordance with appropriate regulatory provisions or proprietary agreements.
(8)     The information identified in paragraph (2), above, is classified as proprietary because it contains detailed bases, results, and conclusions regarding GE Methods supporting evaluations of the safety-significant changes necessary to demonstrate the regulatory acceptability for the expanded power/flow range of extended power uprate for a GE BWR, utilizing analytical models and methods, including computer codes, which GE has developed, obtained NRC approval of, and applied to perform evaluations of transient and accident events in the GE Boiling Water Reactor ("BWR"). The development and approval of these system, component, and thermal hydraulic models and computer codes was achieved at a significant cost to GE, on the order of several million dollars.
The development of the evaluation process along with the interpretation and application of the analytical results is derived from the extensive experience database that constitutes a major GE asset.

(9)     Public disclosure of the information sought to be withheld is likely to cause substantial harm to GE's competitive position and foreclose or reduce the availability of profit-making opportunities. The information is part of GE's comprehensive BWR safety and technology base, and its commercial value extends beyond the original development cost.
The value of the technology base goes beyond the extensive physical database and analytical methodology and includes development of the expertise to determine and apply the appropriate evaluation process. In addition, the technology base includes the value derived from providing analyses done with NRC-approved methods.
The research, development, engineering, analytical and NRC review costs comprise a substantial investment of time and money by GE.
The precise value of the expertise to devise an evaluation process and apply the correct analytical methodology is difficult to quantify, but it clearly is substantial.
GE's competitive advantage will be lost if its competitors are able to use the results of the GE experience to normalize or verify their own process or if they are able to claim an equivalent understanding by demonstrating that they can arrive at the same or similar conclusions.
The value of this information to GE would be lost if the information were disclosed to the public. Making such information available to competitors without their having been required to undertake a similar expenditure of resources would unfairly provide competitors with a windfall, and deprive GE of the opportunity to exercise its competitive advantage to seek an adequate return on its large investment in developing these very valuable analytical tools.
I declare under penalty of perjury that the foregoing affidavit and the matters stated therein are true and correct to the best of my knowledge, information, and belief.
Executed on this 3 0    day of July 2005.
Robert E. Gamble General Electric Company

BVY 05-072 Docket No. 50-271 Attachment 13 Vermont Yankee Nuclear Power Station Proposed Technical Specification Change No. 263 - Supplement No. 30 Extended Power Uprate Response to Request for Additional Information Commitments I Total number of pages in Attachment 13 l (excludina this cover sheet) is 1.       I

ENN                  NON-QUALITY RELATED               EN-LI-106       Revision I NUCLEAR                      ADMINISTRATIVE z-Entergy                MANAGEMENT MANUAL                    INFORMATION   USE                   Page       I of     I Licensee Identified Commitment Form This form identifies actions discussed in this letter for which Entergy Nuclear Operations, Inc. (Entergy) commits to perform. Any other actions discussed in this submittal are described for the NRC's information and are not commitments.
(BVY 05-072)
TYPE (Check one)
COT)MI ME                                                SCHEDULED E                COMPLETION DATE
:  c      a E.         (If Required) 0 Entergy has established administrative controls to assure performance of a once per operating cycle tabletop review                             X         Oct. 31, 2005 of the procedures that complete the actions to repower a VYNPS 4KV bus from the VHS. Pursuant to discussions with co-host REMVEC, a system-wide annual tabletop review will take place in October 2005. During this meeting Entergy will lead a tabletop review of all actions required to support the restoration of 4KV AC to VYNPS.
This review will review the interfaces with the operator of VHS and the regional grid operator to verify that roles and responsibilities and timelines are understood and that there have been no changes that would impact the assumption in the VYNPS SBO coping strategy. Entergy will also provide the participants with additional insights regarding offsite power issues for nuclear power stations including plant response to and consequences of a SBO.
Training on the changes to OP 2124 and OT 3122 is currently underway. Once the training cycle is complete                      X                  Oct. 15, 2005 (scheduled for September 1, 2005), the procedures will be revised and issued for use.
Various operating, surveillance and administrative procedures will be revised to incorporate a higher condensate storage tank inventory limit as either a precaution or an administrative limit by October 15, 2005.}}