NUREG-1826

APEX-AP1000 Confirmatory Testing

To Support 

AP1000 Design Certification 

(Non-Proprietary)

 U.S. Nuclear Regulatory Commission

Office of Nuclear Regulatory Research

Washington, DC 20555-0001
 

 NUREG-1826

APEX-AP1000 Confirmatory Testing

To Support

AP1000 Design Certification

(Non-Proprietary)

Manuscript Completed: June 2005

Date Published: August 2005

Prepared by

K.B. Welter and S.M. Bajorek

Advanced Reactors and Regulatory Effectiveness Branch

Division of Systems Analysis and Regulatory Effectiveness

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555-0001

 

Jose' Reyes, Jr., Brian Woods, John Groome, John Hopson,

Eric Young, John DeNoma, and Kent Abel

Department of Nuclear Engineering

Institute of Thermal-Hydraulics

Oregon State University

Corvallis, OR 97601

Gene S. Rhee, NRC Project Manager

Prepared for

Division of Systems Analysis and Regulatory Effectiveness

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555-0001
ii 

PROPRIETARY INFORMATION NOTICE 

This report does not contain any information that is proprietary to Westinghouse Electric 

Company. The U.S. Nuclear Regulatory Commission (NRC) has also prepared a proprietary 

version of this report for internal distribution. 

In order to conform to the Commission's regulations concerning the protection of proprietary 

information so submitted to the NRC, as set forth in Title 10, Section 2.790, of the Code of 

Federal Regulations (10 CFR 2.790), the proprietary information in the proprietary version 

of this document is contained within brackets and, where the proprietary information has been 

deleted in the non-proprietary version, only the brackets remain (the information within the brackets 

in the proprietary version having been deleted). The justification for claiming the information 

as proprietary is indicated in both versions by means of lower case letters (a) through (f), 

located as a superscript immediately following the brackets enclosing each bracketed item 

of proprietary information, or in the margin opposite such information. These lower case letters 

refer to the types of information Westinghouse customarily holds in confidence, as identified 

in Sections (4)(ii)(a) through (4)(ii)(f) of the affidavit accompanying the transmittal of these 

documents pursuant to 10 CFR 2.790(b)(1). 
iii 

ABSTRACT 

This report summarizes the confirmatory testing performed using the Advanced Plant Experiment 

(APEX) facility at Oregon State University (OSU). The APEX is a unique thermal-hydraulic 

integral system test facility, which is used to assess the performance of passive safety systems 

and was modified to accurately represent the design of the Westinghouse AP1000 advanced 

passive nuclear reactor. The U.S. Nuclear Regulatory Commission (NRC) sponsored eight 

beyond-design-basis accident (DBA) tests in the APEX facility, which were successfully completed 

from June 2003 through July 2004 and are discussed in detail in this report. Those eight 

beyond-DBA tests investigated scenarios with two or more simultaneous failures of the AP1000 

passive safety systems during large- and small-break loss-of-coolant accidents, including 

station blackout and cold shutdown conditions. The experiments run in APEX-AP1000 confirm 

significant liquid entrainment and carryover of water to the automatic depressurization system 

(ADS) during and after actuation of the fourth-stage (ADS4) valves. These processes are 

important as thermal-hydraulic codes used to analyze the AP1000 design must adequately 

predict or bound upper plenum and hot leg entrainment. In addition, the tests show that failure of 

ADS4 valves on the non-pressurizer side of the plant results in a greater delay in in-containment 

refueling water storage tank (IRWST) injection than failure of ADS4 valves on the pressurizer 

side of the plant; while failure of two out of four ADS4 valves on the non-pressurizer side of the 

plant produces low two-phase mixture levels in the core during certain simulated vessel injection 

line and cold leg breaks. 
iv 
v 

FOREWORD 

The U.S. Nuclear Regulatory Commission (NRC) has reviewed Westinghouse Electric Company’s 

request, dated March 28, 2002, for final design approval and standard design certification 

for the AP1000 advanced passive nuclear reactor. The AP1000 standard design is essentially 

an uprated version of Westinghouse’s AP600 design, which the NRC certified on December 16, 1999. 

Following receipt of an application for certification of a new nuclear power plant design, 

the NRC’s standard design certification process involves an acceptance review, a technical review, 

and a rulemaking to certify the design. The AP1000 design certification review involved 

technical reviewers from the NRC’s Offices of Nuclear Reactor Regulation, Nuclear Security 

and Incident Response, and Nuclear Regulatory Research. These reviewers carefully considered 

the AP1000 Design Control Document (DCD) and Probabilistic Risk Assessment in accordance with 

the applicable and technically relevant standards for reviewing design certification applications, 

as set forth in Title 10, Section 52.48, of the Code of Federal Regulations (10 CFR 52.48). 

In addition, the reviewers considered those standards that were modified by the exemptions 

identified in the final safety evaluation report (FSER) for the AP1000 design, which the NRC 

published as NUREG-1793 in September 2004. 

In support of the NRC’s design certification activities, this report documents the confirmatory 

integral testing conducted at Oregon State University’s Advanced Plant Experiment (APEX) facility. 

Specifically, the NRC sponsored 8 beyond-design-basis accident (DBA) tests run in APEX, 

which is a one-quarter height, one-half time scale, reduced-pressure, integral test facility 

modified to represent AP1000 conditions. Those 8 beyond-DBA tests investigated scenarios 

with two or more simultaneous failures of the AP1000 passive safety system during large- and 

small-break loss-of-coolant accidents, including station blackout and cold shutdown conditions. 

The objectives of these beyond-DBA tests were to confirm AP1000 safety margins and provide a 

database to assess NRC’s thermal-hydraulic computer codes. This report presents the results of 

these tests, along with a detailed scaling analysis and facility description. 

Review of the APEX-AP1000 test results conclusively shows that the core remains cooled and did 

not experience heatup for most of the beyond-DBA scenarios. These test results were factored 

into the final evaluation of the AP1000 design and, on the basis of their independent analysis 

described in the FSER, the NRC staff concluded that Westinghouse’s application for AP1000 

design certification meets the applicable content requirements of 10 CFR 52.47, as well as 

the review standards of 10 CFR 52.48. 

_______________________________ 

Carl J. Paperiello, Director 

Office of Nuclear Regulatory Research 
vi 
vii 

CONTENTS 

ABSTRACT ...................................................................................................................................i

FOREWORD.................................................................................................................................v

EXECUTIVE SUMMARY ...........................................................................................................xiii

ABBREVIATIONS ......................................................................................................................xv

1. INTRODUCTION............................................................................................................1-1

2. SCALING ANALYSIS RESULTS ..................................................................................2-1

3. FACILITY DESCRIPTION .............................................................................................3-1

3.1 Reactor Coolant System (RCS)..........................................................................3-2

3.2 Reactor Pressure Vessel (RPV).........................................................................3-3

3.3 Pressurizer (PZR).............................................................................................3-10

3.4 Steam Generators (SGs)..................................................................................3-12

3.5 Reactor Coolant Pump (RCP) ..........................................................................3-16

3.6 Primary Loop Piping .........................................................................................3-19

3.6.1 Hot Legs (HLs) .....................................................................................3-20

3.6.2 Cold Legs (CLs)....................................................................................3-21

3.6.3 Pressurizer Surge Line.........................................................................3-23

3.7 Passive Safety Systems...................................................................................3-23

3.7.1 Automatic Depressurization System (ADS) 1–4...................................3-23

3.7.2 In-Containment Refueling Water Storage Tank (IRWST).....................3-26

3.7.3 Core Makeup Tanks (CMT) ..................................................................3-29

3.7.4 Passive Residual Heat Remover (PRHR) Heat Exchanger (HX) .........3-32

3.7.5 Accumulators (ACC).............................................................................3-34

3.7.6 Containment Sump...............................................................................3-36

3.8 Instrumentation, Control, and Power Systems .................................................3-37

3.8.1 Instrumentation Description..................................................................3-37

3.8.2 Data Acquisition....................................................................................3-38

3.8.3 Control System and Control Logic........................................................3-40

3.8.4 Power System ......................................................................................3-42

3.8.5 Break and ADS Measurement System (BAMS) ...................................3-44

3.9 Balance of Plant ...............................................................................................3-47

3.9.1 Chemical and Volume Control System (CVS) ......................................3-47

3.9.2 Normal Residual Heat Removal System (RNS) ...................................3-47

3.9.3 Feedwater System................................................................................3-49

3.9.4 Water Purification System ....................................................................3-50

4. TEST MATRIX AND OVERVIEW ..................................................................................4-1

4.1 Typical Sequence of Events for an APEX-AP1000 SBLOCA.............................4-3

4.1.1 High-Pressure Blowdown Phase............................................................4-3

4.1.2 Natural Circulation Phase.......................................................................4-4

4.1.3 ADS Blowdown Phase............................................................................4-4

4.1.4 IRWST Injection Phase ..........................................................................4-4
viii 

4.1.5 Sump Recirculation Phase .....................................................................4-5

4.2 Test NRC-AP1000-01 ........................................................................................4-5

4.3 Test NRC-AP1000-02 ........................................................................................4-7

4.4 Test NRC-AP1000-03 ........................................................................................4-7

4.5 Test NRC-AP1000-04 ......................................................................................4-10

4.6 Test NRC-AP1000-05 ......................................................................................4-12

4.7 Test NRC-AP1000-06 ......................................................................................4-12

4.8 Test NRC-AP1000-10 ......................................................................................4-15

4.9 Test NRC-AP1000-11 ......................................................................................4-15

5. COMPARISONS AND DISCUSSION............................................................................5-1

5.1 Test DBA-02 versus Test NRC-AP1000-05 .......................................................5-1

5.2 Test NRC-AP1000-03 versus NRC-AP1000-05.................................................5-4

6. SUMMARY AND CONCLUSIONS ................................................................................6-1

7. REFERENCES...............................................................................................................7-1
ix 

APPENDICES 

A. Scaling Analysis .............................................................................................................. A-1

B. APEX Component Dimensions, Control Logic, and Setpoints.................................... B-1

C. Test Transient Plots and Sequence of Events.............................................................. C-1
x 

Figures 

1.1 APEX Facility Layout......................................................................................................1-2

2.1 Important Phenomena During Transition from ADS4 Blowdown to IRWST Injection. ...2-3

2.2 Upper Plenum Pool Entrainment Regions in AP1000 and APEX-AP1000. ...................2-3

3.1 Elevation View of Reactor Coolant System....................................................................3-2

3.2 Plan View of Reactor Coolant System ...........................................................................3-2

3.3 Reactor Cross-Section View ..........................................................................................3-4

3.4 View of Reactor Core from Above..................................................................................3-4

3.5 Photograph of Pressurizer in the APEX Facility...........................................................3-10

3.6 Steam Generator Being Delivered to the APEX Facility...............................................3-12

3.7 Steam Generator U-Tube Bundle ................................................................................3-13

3.8 Steam Generator Components ....................................................................................3-13

3.9 Image of Two Reactor Coolant Pumps from Bottom....................................................3-17

3.10 Reactor Coolant Pump Head Curve.............................................................................3-18

3.11 Primary Loop Piping in the Reactor Coolant System ...................................................3-19

3.12 Pressurizer Surge Line Geometry................................................................................3-23

3.13 Illustration of Automatic Depressurization System 1-3 Sparger...................................3-24

3.14 Photograph of Installed Sparger ..................................................................................3-25

3.15 Photograph of the IRWST ............................................................................................3-29

3.16 Photograph of a Core Makeup Tank ............................................................................3-30

3.17 View of the PRHR HX Located Inside the IRWST .......................................................3-33

3.18 Photograph of One of the APEX Accumulators ...........................................................3-35

3.19 Data Acquisition System Hardware Overview..............................................................3-39

3.20 Control System Interface..............................................................................................3-40

3.21 Photograph of the APEX Control Panel .......................................................................3-41

3.22 Photograph of the SCR Cabinet...................................................................................3-43

3.23 Main Power System Simple Layout .............................................................................3-43

3.24 BAMS Layout ...............................................................................................................3-44

3.25 RNS Pump Performance Curves .................................................................................3-48

3.26 Image of RNS, CVS, and MF Pumps...........................................................................3-49

4.1 Pressure History for a Typical 2-inch Cold Leg Break in APEX .....................................4-3

4.2 Test NRC-AP1000-01 Passive Safety System Performance.........................................4-6

4.3 Test NRC-AP1000-02 Passive Safety System Performance.........................................4-8
xi 

4.4 Test NRC-AP1000-03 Passive Safety System Performance.........................................4-9

4.5 Test NRC-AP1000-04 Passive Safety System Performance.......................................4-11

4.6 Test NRC-AP1000-05 Passive Safety System Performance.......................................4-13

4.7 Test NRC-AP1000-06 Passive Safety System Performance.......................................4-14

4.8 Test NRC-AP1000-10 Passive Safety System Performance.......................................4-17

4.9 Test NRC-AP1000-11 Passive Safety System Performance.......................................4-18

5.1 Reactor Vessel Liquid Level (DBA-02 and NRC-AP1000-05)........................................5-1

5.2 IRWST Injection Flow Rates (DBA-02 and NRC-AP1000-05). ......................................5-2

5.3 PZR Liquid Levels (DBA-02 and NRC-AP1000-05). ......................................................5-3

5.4 ADS4-2 Flow Qualities (DBA-02 and NRC-AP1000-05). ...............................................5-3

5.5 System Pressures (NRC-AP1000-03 and 05)................................................................5-5

5.6 Safety Injection Flow Rates (NRC-AP1000-03 and 05) .................................................5-5

5.7 ADS4 Flow Qualities (NRC-AP1000-03 and 05)............................................................5-6

5.8 RPV Collapsed Liquid Levels (NRC-AP1000-03 and 05) ..............................................5-6

5.9 RPV Two-Phase Mixture Levels (NRC-AP1000-03 and 05) ..........................................5-7

5.10 HL2 Collapsed Liquid Levels (NRC-AP1000-03 and 05) ...............................................5-7

5.11 HL1 Collapsed Liquid Levels (NRC-AP1000-03 and 05) ...............................................5-8

5.12 PZR Surge Line Liquid Levels (NRC-AP1000-03 and 05) .............................................5-8

5.13 PZR Collapsed Liquid Levels (NRC-AP1000-03 and 05)...............................................5-9

5.14 IRWST Injection Flow Rates (NRC-AP1000-03 and 05)................................................5-9

Tables 

2.1 Modifications to APEX Components To Simulate AP1000 Conditions. .........................2-1

3.1 Thermocouples in the Reactor Pressure Vessel............................................................3-5

3.2 Pressure Instrumentation in the Reactor Pressure Vessel.............................................3-8

3.3 Pressurizer Instrument Names and Locations .............................................................3-11

3.4 Steam Generator Instrument Names and Locations (Secondary Side) .......................3-14

3.5 Steam Generator Instrument Names and Locations (Primary Side)............................3-15

3.6 Reactor Coolant Pump Instrument Names and Locations ...........................................3-18

3.7 Hot Leg Instrument Names and Locations...................................................................3-20

3.8 Cold Leg Instrument Names and Locations .................................................................3-21

3.9 Automatic Depressurization System 1-4 Instrument Names and Locations ................3-25

3.10 IRWST LDP, LCT, FMM, DP, and PT Cells and Locations..........................................3-27
xii 

3.11 IRWST Thermocouples and Location ..........................................................................3-28

3.12 Core Makeup Tank Instrument Names and Locations.................................................3-30

3.13 PRHR Instrument Names and Locations .....................................................................3-34

3.14 Accumulator Instrument Names and Locations ...........................................................3-35

3.15 Sump Instrument Names and Locations ......................................................................3-36

3.16 Instrument Tag Locations.............................................................................................3-38

3.17 ADS4-1/4-2, and Break Separator Instrument Names and Locations..........................3-45

3.18 CVS Instrument Names and Locations ........................................................................3-47

3.19 RNS Instrument Names and Locations........................................................................3-48

3.20 FST, MFP, and SG Steam Vent Instrument Names and Locations .............................3-50

4.1 APEX-AP1000 Confirmatory Test Matrix .......................................................................4-2
xiii 

EXECUTIVE SUMMARY 

This report documents the confirmatory testing performed using the Advanced Plant Experiment 

(APEX) facility at Oregon State University (OSU). The APEX is a unique thermal-hydraulic 

integral system test facility, which is used to assess the performance of passive safety systems 

and was modified to accurately represent the designs of the Westinghouse AP600 and AP1000 

advanced passive nuclear reactors. The U.S. Nuclear Regulatory Commission (NRC) sponsored 

eight beyond-design-basis accident (DBA) tests of the AP1000 design, which were successfully 

completed in the APEX facility from June 2003 through July 2004. This program had the following 

objectives: 

• Conduct tests to assess the performance of the Westinghouse AP1000 passive safety systems 

for a select range of beyond-DBA scenarios. 

• Generate data to confirm analyses that Westinghouse submitted for AP1000 design 

certification. 

• Obtain experimental data to benchmark the NRC’s thermal-hydraulic computer codes. 

The NRC extensively used data from the APEX facility in the AP600 design certification process, 

drawing upon approximately 75 tests for use in code assessment and qualification of the AP600 

safety margin. To address performance specific to the AP1000 design (essentially an uprated 

version of the AP600 design), the APEX facility underwent significant modifications in 2002 

in order to more accurately represent the AP1000. That facility modification included an increase 

in the maximum core power; new pressurizer (PZR) and surge lines; larger core makeup tanks 

(CMTs); larger-diameter fourth-stage automatic depressurization system (ADS4) piping; 

and decreased line resistances for the CMTs, ADS4, and passive residual heat removal (PRHR) 

heat exchanger. In addition, Reyes performed a comprehensive scaling analysis (Appendix A) 

to guide the facility modifications to ensure that the APEX experiments preserved AP1000 conditions. 

This study investigated several beyond-DBA scenarios, including 1- and 2-inch cold leg breaks, 

double-ended direction vessel injection (DEDVI) line breaks, Mode 5 operation with loss of 

residual normal heat removal system (RNS) cooling, and station blackout. Based on experience 

from analyzing previous AP600 tests, one of the main purposes of the APEX-AP1000 test program 

was to investigate the sensitivity of the failure of the automatic depressurization system (ADS) 

on core coolability. 

Based on the experiments run in the APEX facility, we can draw the following conclusions: 

(1) The APEX-AP1000 tests confirm significant liquid entrainment and carryover of water to 

the ADS4 system during and after ADS4 valve actuation. High liquid carryover to the 

ADS4 may occur in the AP1000 reactor. These data are appropriate for assessment of 

thermal-hydraulic codes used to analyze the AP1000. 

(2) Test results from tests NRC-AP1000-03, 05, and DBA-02 show that failure of ADS4 

valves on the non-pressurizer side of the plant results in a greater delay in incontainment refueling water storage tank (IRWST) injection than failure of ADS4 valves 

on the pressurizer side of the plant. 
xiv 

(3) Failure of two out of four ADS4 valves on the non-pressurizer side of the plant results in 

core uncovery (based on the two-phase mixture level) during simulated DEDVI line and 

cold leg breaks. 

(4) Review of the APEX-AP1000 DBA test results and code calculations submitted by 

Westinghouse for design certification conclusively show that the core remains cooled, 

and heatup was not experienced for design-basis scenarios. In fact, core heatup was 

only experienced in select beyond-DBA tests (NRC-AP1000-05 and 06). 
xv 

ABBREVIATIONS 

ACC Accumulator 

ADS Automatic Depressurization System 

ADS4 Fourth-Stage Automatic Depressurization System 

AP600 Advanced Passive 600-MWe Plant (Westinghouse) 

AP1000 Advanced Passive 1,000-MWe Plant (Westinghouse) 

APEX Advanced Plant Experiment (OSU Facility) 

ASME American Society of Mechanical Engineers 

BAMS Break and ADS Measurement System 

CCFL Counter-Current Flow Limitation 

CFR Code of Federal Regulations 

CL Cold Leg 

CMT Core Makeup Tank 

CMT-PBL Core Makeup Tank Pressure Balance Line 

CRP Condensate Return Pump 

CRT Condensate Return Tank 

CVS Chemical and Volume System 

DAS Data Acquisition System 

DBA Design-Basis Accident 

DCD Design Control Document (AP1000) 

DEDVI Double-Ended Direction Vessel Injection 

DI Demineralization System 

DP Differential Pressure 

DVI Direct Vessel Injection 

FMM Magnetic Flowmeter 

FSER Final Safety Evaluation Report 

FST Feedwater Storage Tank 

FVM Vortex Flowmeter 

H2TS Hierarchical Two-Tiered Scaling 

HL Hot Leg 
xvi 

HPS Heated Phase Switch 

HX Heat Exchanger 

IRWST In-Containment Refueling Water Storage Tank 

KW Heater Power 

LCT Load Cell Transducer 

LDP Differential Pressure Level 

LOCA Loss-of-Coolant Accident 

LT Level Transducer 

MFP Main Feedwater Pump 

NI National Instruments 

NRC U.S. Nuclear Regulatory Commission 

NSSS Nuclear Steam Supply System 

OSU Oregon State University 

P&ID Piping and Instrumentation Drawings 

PBL Pressure Balance Line 

PC Process Controller 

PIRT Phenomena Identification and Ranking Table 

PLC Programmable Logic Controller 

PORV Power Operated Relief Valve 

PRHR Passive Residual Heat Removal 

PT Pressure Transducer 

PZR Pressurizer 

RCP Reactor Coolant Pumps 

RCS Reactor Coolant System 

RNS Residual Normal Heat Removal System 

ROSA Rig of Safety Assessment 

RPV Reactor Pressure Vessel 

RSO Return to Saturation Oscillation 

Rx Reactor Vessel 

SB Small Break 
xvii 

SBLOCA Small-Break Loss-of-Coolant Accident 

SCR Silicon Control Rectifier 

SG Steam Generator 

T/C Thermocouple 

TF Fluid Temperature 

TH Heater Rod Temperature 

TR Thermocouple Rod 

TW Wall Temperature 
 1-1 

1. INTRODUCTION 

This report documents the confirmatory testing performed using the Advanced Plant Experiment 

(APEX) facility at Oregon State University (OSU). The APEX is a unique thermal-hydraulic 

integral system test facility (Figure 1.1), which is used to assess the performance of passive 

safety systems and was modified to accurately represent the Westinghouse AP600 and AP1000 

advanced passive nuclear reactor designs. The U.S. Nuclear Regulatory Commission (NRC) 

sponsored this research through Contract #NRC-04-03-049, which encompassed eight beyonddesign-basis accident (DBA) tests of the AP1000 design. Testing began in June 2003, 

and the final test was completed in July 2004. This program had the following objectives: 

• Conduct tests to assess the performance of the Westinghouse AP1000 passive safety systems 

for a select range of beyond-DBA scenarios. 

• Generate data to confirm analyses that Westinghouse submitted for AP1000 design 

certification. 

• Obtain experimental data to benchmark the NRC’s thermal-hydraulic computer codes. 

The NRC extensively used data from the APEX facility in the AP600 design certification process, 

drawing upon approximately 75 tests for use in code assessment and qualification of AP600 

safety margin. To address performance specific to the AP1000 design (essentially an uprated 

version of the AP600 design), the APEX facility underwent significant modifications in 2002 

in order to more accurately represent the AP1000. The facility modification included an increase 

in the maximum core power; new pressurizer (PZR) and surge lines; larger core makeup tanks 

(CMTs); larger-diameter fourth-stage automatic depressurization (ADS4) system piping; and 

decreased line resistances for the CMTs, ADS4, and passive residual heat removal (PRHR) 

heat exchanger (HX). 

The study investigated several beyond-DBA scenarios, including 1- and 2-inch cold leg breaks, 

double-ended direction vessel injection (DEDVI) line breaks, Mode 5 operation with loss of 

residual normal heat removal system (RNS) cooling, and station blackout. Based on experience 

from analyzing previous AP600 tests, one of the main purposes of the APEX-AP1000 test program 

was to investigate the sensitivity of the failure of the automatic depressurization system (ADS) 

on core coolability. Therefore, for each accident scenario, one or more ADS valves were failed 

on demand or were unavailable for each test. 

Chapter 1 includes this introduction, while Chapter 2 describes the results of the comprehensive 

scaling analysis (Appendix A) conducted by Reyes to guide the modifications to upgrade the APEX 

facility from AP600 to AP1000 geometries and test conditions. In Chapter 3, the reader can find a 

detailed facility description with component and instrument information (component dimensions can be 

found in Appendix B). Chapter 4 summarizes the APEX-AP1000 test results, while Appendix C contains 

more detailed instrument plots and sequence of events for each test. Readers should note that the 

non-proprietary version of this report the majority of the information in Appendices A, B, and C 

has been removed. In addition, a supplement to the proprietary version of this report includes 

three compact disks containing APEX facility drawings, test reports, and the APEX database. 
 1-2 

Figure 1.1. APEX Facility Layout 
 2-1 

2. SCALING ANALYSIS RESULTS 

The APEX-AP1000 scaling analysis by Reyes (Appendix A) provided the basis for modifying 

core decay power; CMT and PZR volumes; the in-containment refueling water storage tank 

(IRWST) and containment flood-up elevations; ADS4, IRWST, PRHR, and CMT line resistances; 

PZR surge line diameter; ADS4, upper core plate, and upper support plate flow areas; and 

upper plenum structures. The majority of the APEX-AP1000 scaling analysis was based on results 

from the original AP600 scaling analysis because the AP1000 is geometrically similar to the AP600, 

and the thermal-hydraulic phenomena studied fell within the scope of the original AP600 test program. 

The purpose of the APEX-AP1000 confirmatory test program was to examine AP1000 passive 

safety system performance and provide high-quality benchmark data for the NRC’s thermalhydraulic computer codes used in reactor safety analysis. Previous testing in the APEX facility 

demonstrated that all of the thermal-hydraulic phenomena observed in large-scale AP600 

facilities (such as ROSA-AP600 or SPES) were also observed in APEX and, consequently, 

provide a broad database for code assessment. However, the most faithful representation of 

the AP600 phenomena was for the very-long-term cooling phase, which occurs at low system 

pressure. The APEX-AP1000 test program was modified to include full-pressure ADS4 

blowdown tests to investigate the transition period that leads to the onset of IRWST injection. 

The scaling analysis provided the basis for the plant modifications listed in Table 2.1. 

Table 2.1. Modifications to APEX Components To Simulate AP1000 Conditions 

Component Modification to APEX 

Reactor Power Increase core power by 67% 

PZR Increase PZR volume 

Reduce PZR Surge Line Diameter 

CMTs Increase CMT volumes by 25% 

Reduce line resistance to 64% of original value 

IRWST Increase IRWST liquid level 

ADS4 Increase ADS4 flow area by 76% 

Reduce line resistance to 28% of original value 

PRHR HX Increase PRHR flow capacity by 74% by reducing line 

resistance 

No change in surface area/tube number required for testing 

Containment Sump Flood-Up Elevation Change flood-up elevation in primary sump tank 
 2-2 

The following paragraphs briefly summarize the results of the scaling study. 

(1) The original AP600 scaling analysis, WCAP-14270 [Ref. 1], which was conducted in support 

of the existing APEX facility design, was found to be applicable to the APEX-AP1000 design 

for two reasons. First, the thermal-hydraulic phenomena to be studied fell within the scope 

of the original AP600 test program, with the exception of upper plenum behavior. Second, 

the AP1000 design is geometrically similar to the AP600 design. As a result, all tank volumes, 

flow areas, and line resistance scaling ratios found in WCAP-14270 were used 

in the present study. 

(2) An NRC PZR surge line study by Bessette and Di Marzo [Ref. 2] guided the modification 

of the APEX-AP1000 PZR surge line. The reduced line diameter is expected to produce 

a better simulation of the PZR draining behavior. 

(3) The reactor coolant system (RCS) depressurization scaling analysis represents a more 

comprehensive approach than that implemented in the original AP600 study. Nonetheless, 

the derived similarity criteria resulted in scaling ratios that were identical to those in 

the original study. Thus, it was determined that the characteristic time ratios for the AP1000 

reactor would be well-matched in APEX-AP1000. The greatest distortion (a factor of 2) 

was the result of pressure-scaled fluid properties. 

(4) A scaling analysis was performed to design the APEX-AP1000 facility to simulate 

the important phenomena that occur during the transition from ADS4 blowdown to the onset 

of IRWST injection, as shown in Figure 2.1. The key phenomena have been identified 

and described, and the scaling analysis indicated that full-pressure ADS4 blowdown tests 

could be performed using the modified APEX-AP1000 test facility. The maximum decay 

power, at the time of ADS4 valve opening, was calculated for the DEDVI break. The 

AP1000 core axial void fraction profile and core averaged void fraction were also shown to 

be preserved in the APEX-AP1000 facility. 

(5) A detailed pool entrainment scaling analysis was performed to design the upper plenum 

components in the APEX-AP1000 facility. The analysis indicated that well-scaled entrainment 

behavior is expected in the APEX facility to a depth of about 18.9 cm (0.55 in terms of 

dimensionless height) below the hot leg, as shown in Figure 2.2. Below this depth, the APEX 

facility is in the momentum-controlled region, whereas the AP1000 reactor would be in the 

deposition-controlled region at the equivalent scaled depth, as shown in Figure 2.2. It was 

also determined that the entrainment rates in the near-surface region are a factor of 150 

times greater than those generated in the momentum-controlled region, which are a factor 

of 4 to 6 times greater than those generated in the deposition-controlled region. Thus the 

APEX facility will conservatively simulate the entrainment behavior when the level is below 

the midpoint of the upper plenum. 
 2-3 

PADS-4

Sweepout of

Liquid Above

Hot Leg

Liquid Boil-off,

Flashing, and

Carryover

RPV MIXTURE LEVEL

PRESSURE

PIRWST

Top of Hot Leg

Top of Core

Two-Phase Fluid

ADS-4 Flow

High Quality Vapor

ADS-4 Flow

IRWST

Injection

Figure 2.1. Important Phenomena During Transition from ADS4 Blowdown to IRWST Injection 

0 0.2 0.4 0.6 0.8 1

Dimensionless Height (Bottom of Hot Leg to Top of Active Core)

APEX-AP1000 AP1000

NEAR 

SURFACE

REGION

MOMENTUM CONTROLLED

REGION

DEPOSITION CONTROLLED

REGION

NEAR 

SURFACE

REGION

MOMENTUM CONTROLLED

REGION

Bottom of 

Hot Leg

Top of Active Core

Figure 2.2. Upper Plenum Pool Entrainment Regions in AP1000 and APEX-AP1000
 3-1 

3. FACILITY DESCRIPTION 

This chapter provides a comprehensive description of the APEX-AP1000 test facility, including 

the reactor coolant system (RCS); passive safety systems; instrumentation, controls, and power 

systems; and balance-of-plant components. 

The OSU Department of Nuclear Engineering has modified its APEX facility to assess the performance 

of the AP1000 passive safety system. APEX is a unique, world-class, one-quarter height, 

one-half time scale, reduced-pressure, thermal-hydraulic, integral system test facility. In the past, 

the data obtained using the APEX facility formed an essential aspect of the AP600 design certification 

program. Modified to accurately represent the AP1000 reactor, the APEX facility is referred to 

as APEX-AP1000. Appendix B to this report presents information about the APEX-AP1000 

facility, such as piping and component dimensions, setpoints, and control logic. 

The APEX test facility, depicted in Figure 1.1, has been specifically designed and constructed 

to provide high-quality data for use in computer code benchmark calculations. Table 2.1 

summarizes the significant modifications required to adapt the facility for AP1000 testing, 

which included new reactor heater rods to increase the maximum facility power from 600 kW 

to 1000 kW. In addition, the PZR and CMTs were increased in size, passive core cooling system 

injection line resistance was reduced, and the ADS4 valves and associated piping were increased 

in size to provide larger RCS depressurization capacity. In addition, a formal scaling analysis 

(Appendix A) was performed to ensure that the facility accurately represents the AP1000 

reactor. The resulting APEX-AP1000 test facility includes the following systems: 

• Reactor Coolant System: This includes an electrically heated 48-rod bundle core, 

a reactor vessel with internals, two hot legs (HLs), four cold legs (CLs), two 133 U-tube 

steam generators (SGs), one PZR, and four reactor coolant pumps (RCPs). 

• Passive Safety Systems: This includes two CMTs, two accumulators (ACCs), a four-stage 

ADS, one PRHR HX, one IRWST, and portions of the lower containment compartments. 

• Balance of Plant: This includes a feedwater system, non-safety grade chemical volume 

control system (CVS), and an active residual heat removal system. The geometry 

of the interconnecting pipe routings was also duplicated. 

All of the RCS components are constructed of stainless steel and capable of consistent operation 

at 400 psia (2.76 MPa) while at saturation temperatures. All primary system components 

are insulated to minimize heat loss. 
 3-2 

3.1 Reactor Coolant System (RCS)

The APEX facility RCS is a complete model of the AP1000 nuclear steam supply system (NSSS). 

As such, the RCS includes the reactor pressure vessel (RPV), one PZR, two SGs, four RCPs, 

and associated primary loop piping. See Figures 3.1 and 3.2 for elevation and plan views 

of the RCS, respectively. 

Figure 3.1. Elevation View of Reactor Coolant System 

Figure 3.2. Plan View of Reactor Coolant System 
 3-3 

3.2 Reactor Pressure Vessel (RPV)

The RPV models the upper and lower internals of the AP1000 reactor vessel, core barrel, 

downcomer, and core. The maximum core power is approximately 1000 kW, which can be 

distributed in two radial power zones and can be programmed to simulate time-dependent 

decay power. The RPV includes connections for the two HLs, four CLs, and two direct vessel 

injection (DVI) lines. 

During normal operation, cold water enters through the four CLs into an annular downcomer region 

that is bounded by the inside surface of the reactor vessel shell and the outside surface 

of the core barrel. The cylindrical reactor vessel shell is fabricated from Stainless Steel (SS) 

304 plate. The core barrel extends from the upper head flange down to the lower core plate and 

is made of SS 304. 

The cold water in the downcomer flows into the lower plenum, a cylindrical fluid volume region 

bounded vertically by the lower core plate and the lower head plate. The lower head plate 

is bolted to the carbon-steel reactor vessel flange, which is welded to the cylindrical reactor 

vessel shell. The fluid changes direction in the lower plenum and travels upward through holes 

in the lower core plate into the core. The upper vessel head is a 2:1 elliptical head. 

The heated zone of the core consists of 48 heater rods and 5 fluid thermocouple (T/C) rods. 

Six of the heater rods have a T/C located at the disc end (top) of the heater. In addition to 

the heater rod T/Cs and individual fluid T/Cs in the RPV, there are five T/C rods that measure 

the axial fluid temperature distribution in the core. All of the heater rods penetrate the lower 

head plate, lower plenum, and lower core plate to create a heated zone between the lower 

core plate and the upper core plate (see Figure 3.3). Two spacer grids are provided for support 

of the heaters, one at the mid-plane of the heaters and the other near the core exit. The heater 

bundle is surrounded by a reflector/baffle that directs the fluid through the core (see Figure 3.4). 

The reflector is filled with PUROCastTM (ceramic) to minimize the heat capacity of the core region. 
 3-4 

Figure 3.3. Reactor Cross-Section View 

Figure 3.4. View of Reactor Core from Above 
 3-5 

The heated water leaves the top of the core and flows upward through the upper core plate 

into the upper plenum. The upper plenum is bounded by the upper core plate and lower support 

plate (see Figure 3.3). The upper core plate supports the upper plenum internals that simulate 

the control rod guide tubes and other upper plenum structure. The heated water leaves 

the upper plenum via the two HL nozzles. 

The volume above the upper support plate is the upper head region. It is connected to 

the upper plenum via 10 holes and a single hole in the upper support plate. The hole in the center 

of the upper support plate allows a capacitance level probe to penetrate to just above the upper 

core plate. The Drexelbrook Universal LiteTM capacitance probe provides a mixture level from the 

top of the upper core plate to the upper support plate. It is also connected to the downcomer via 10 

bypass holes located at the top of the downcomer. The Universal LiteTM instrument is a 

capacitance-to-current transducer. A change in level produces a change in capacitance 

causing a corresponding change in current (4-20 mA). It is called a two-wire transmitter 

because the same two wires that are used to power the unit also indicate the change in level. 

Table 3.1 lists the T/Cs located in the RPV, while Table 3.2 lists he pressure instrumentation 

in the RPV. 

Table 3.1. Thermocouples in the Reactor Pressure Vessel

Tag 

Name Description Tag Name Description 

TF-101 CL-3/Reactor Flange @ Top TF-163 Lower Rx Vessel Layer A-A @ 135 degrees 

TF-102 CL-4/Reactor Flange @ Top TF-164 Upper Rx Vessel Layer H-H @ 0 degrees 

TF-103 CL-3/Reactor Flange @ Bottom TF-165 Upper Rx Vessel Layer H-H @ 6.2 degrees 

TF-104 CL-4/Reactor Flange @ Bottom TF-166 Upper Rx Vessel Layer I-I @ 0 degrees 

TF-105 CL-1/Reactor Flange @ Top TF-167 Heater Rod B2-319 @ 40.13" 

TF-106 CL-2/Reactor Flange @ Top TF-168 Upper Rx Vessel Layer K-K @ 270 degrees 

TF-107 CL-1/Reactor Flange @ Bottom TF-169 Upper Rx Vessel Layer M-M @ 90 degrees 

TF-108 CL-2/Reactor Flange @ Bottom TF-170 Upper Rx Vessel Layer M-M @ 270 degrees 

TF-113 DVI-1/Reactor Flange @ Top TF-171 Top of Reactor Down to within 0.50" 

of Upper Support Plate 

TF-114 DVI-2/Reactor Flange @ Bottom TF-172 Lower Rx Vessel Layer AA-AA @ 0 degrees 

TF-115 DVI-1/Reactor Flange @ Bottom TF-173 Lower Rx Vessel Layer AA-AA @ 270 degrees 

TF-116 DVI-2/Reactor Flange @ Top TR-001-1 Core Thermocouple Rod D-001 @ 10.50" 

TF-118 Lower Rx Vessel Layer Y-Y 

@ 30 degrees 

TR-001-2 Core Thermocouple Rod D-001 @ 19.13" 

TF-120 Top of Reactor @ 8.5" 

& 350 degrees 

TR-001-3 Core Thermocouple Rod D-001 @ 25.13" 
 3-6 

Table 3.1. Thermocouples in the Reactor Pressure Vessel

Tag 

Name Description Tag Name Description 

TF-126 Lower Rx Vessel Layer A-A 

@ 225 degrees 

TR-001-4 Core Thermocouple Rod D-001 @ 31.13" 

TF-127 Lower Rx Vessel Layer A-A 

@ 315 degrees 

TR-001-5 Core Thermocouple Rod D-001 @ 37.13" 

TF-128 Lower Rx Vessel Layer C-C 

@ 0 degrees 

TR-001-6 Core Thermocouple Rod D-001 @ 43.13" 

TF-129 Lower Rx Vessel Layer C-C 

@ 32 degrees 

TR-303-1 Core Thermocouple Rod D-303 @ 10.51" 

TF-130 Lower Rx Vessel Layer G-G 

@ 0 degrees 

TR-303-2 Core Thermocouple Rod D-303 @ 19.13" 

TF-131 Lower Rx Vessel Layer G-G 

@ 11.3 degrees 

TR-303-3 Core Thermocouple Rod D-303 @ 25.13" 

TF-132 Upper Rx Vessel Layer F-F 

@ 0 degrees 

TR-303-4 Core Thermocouple Rod D-303 @ 31.13" 

TF-133 Upper Rx Vessel Layer F-F 

@ 8 degrees 

TR-303-5 Core Thermocouple Rod D-303 @ 37.13" 

TF-134 Upper Rx Vessel Layer E-E 

@ 0 degrees 

TR-303-6 Core Thermocouple Rod D-303 @ 43.13" 

TF-135 Upper Rx Vessel Layer E-E 

@ 6.2 degrees 

TR-308-1 Core Thermocouple Rod E-308 @ 22.13" 

TF-140 HL-2/Reactor Flange @ Top TR-308-2 Core Thermocouple Rod E-308 @ 34.13" 

TF-141 HL-1/Reactor Flange @ Top TR-308-3 Core Thermocouple Rod E-308 @ 46.13" 

TF-142 HL-2/Reactor Flange @ Bottom TR-308-4 Core Thermocouple Rod D-001 @ 49.13" 

TF-143 HL-1/Reactor Flange @ Bottom TR-308-5 Core Thermocouple Rod D-001 @ 51.13" 

TF-147 Upper Rx Vessel Layer I-I 

@ 180 degrees 

TR-308-6 Core Thermocouple Rod D-303 @ 49.13" 

TF-148 Upper Rx Vessel Layer I-I 

@ 188 degrees 

TR-313-1 Core Thermocouple Rod D-313 @ 10.50" 

TF-149 Upper Rx Vessel Layer H-H 

@ 180 degrees 

TR-313-2 Core Thermocouple Rod D-313 @ 19.13" 

TF-150 Upper Rx Vessel Layer H-H 

@ 186.2 degrees 

TR-313-3 Core Thermocouple Rod D-313 @ 25.13" 
 3-7 

Table 3.1. Thermocouples in the Reactor Pressure Vessel

Tag 

Name Description Tag Name Description 

TF-151 Upper Rx Vessel Layer E-E 

@ 186.2 degrees 

TR-313-4 Core Thermocouple Rod D-313 @ 31.13" 

TF-152 Upper Rx Vessel Layer E-E 

@ 180 degrees 

TR-313-5 Core Thermocouple Rod D-313 @ 37.13" 

TF-153 Upper Rx Vessel Layer F-F 

@ 180 degrees 

TR-313-6 Core Thermocouple Rod D-313 @ 43.13" 

TF-154 Upper Rx Vessel Layer F-F 

@ 188 degrees 

TR-318-1 Core Thermocouple Rod F-318 @ 28.13" 

TF-155 Lower Rx Vessel Layer G-G 

@ 180 degrees 

TR-318-2 Core Thermocouple Rod F-318 @ 40.13" 

TF-156 Lower Rx Vessel Layer G-G 

@ 191.3 degrees 

TR-318-3 Core Thermocouple Rod F-318 @ 51.86" 

TF-157 Lower Rx Vessel Layer C-C 

@ 212 degrees 

TR-318-4 Core Thermocouple Rod D-303 @ 51.13" 

TF-158 Lower Rx Vessel Layer C-C 

@ 180 degrees 

TR-318-5 Core Thermocouple Rod D-313 @ 49.13" 

TF-162 Lower Rx Vessel Layer A-A 

@ 45 degrees 

TR-318-6 Core Thermocouple Rod D-313 @ 51.13" 
 3-8 

Table 3.2. Pressure Instrumentation in the Reactor Pressure Vessel

Tag Name Description 

DP-111 DP Across Upper Core Plate 

DP-114 DP Across Upper Support Plate 

DP-121 DVI-1/Cold Leg 1 Differential Pressure 

DP-122 DP Between DVI-2 and CL-2 

DP-123 DP Between DVI-1 and CL-3 

DP-124 DP Between DVI-2 and CL-4 

DP-125 HL-1 Entrance Losses 

DP-126 HL-2 Entrance Losses 

DP-128 HL-2 Entrance Losses 

LDP-101 CL to Bypass Holes Uncompensated Water Level (270) 

LDP-102 CL to Bypass Holes Water Uncompensated Level (180) 

LDP-103 DVI to CL Uncompensated Water Level (270) 

LDP-104 DVI to CL Uncompensated Water Level (180) 

LDP-105 Upper Core Plate to DVI Uncompensated Water Level (270) 

LDP-106 Bottom of Core to Lower Core Plate Uncompensated Water Level (180) 

LDP-107 Bottom of Core to Lower Core Plate Uncompensated Water Level (270) 

LDP-108 Bottom of Core to Lower Core Plate Uncompensated Water Level (0) 

LDP-109 Lower Core Plate to Mid-Core Spacer Grid (0) Uncompensated Water Level 

LDP-110 Mid-Core Spacer Grid to Upper-Core Spacer Grid (0) Uncompensated Water Level 

LDP-112 Upper Core Plate to DVI Water Level (0) Uncompensated Water Level 

LDP-113 DVI to Bottom of Upper Support Plate (0) Uncompensated Water Level 

LDP-115 Upper Support Plate to Top of Reactor Uncompensated Water Level 

LDP-116 Bottom of Reactor to Bottom of Bypass Holes (270) Uncompensated Water Level 

LDP-117 Upper Core Spacer Grid to DVI Water Uncompensated Level (180) 

LDP-118 Lower Core Plate to Upper Core Plate (270) Uncompensated Water Level 

LDP-119 Lower Core Plate to Upper Core Support Grid (180) Uncompensated Water Level 

LDP-127 Reactor Wide Range Uncompensated Level 
 3-9 

Table 3.2. Pressure Instrumentation in the Reactor Pressure Vessel

Tag Name Description 

LDP-138 Upper Core Spacer Grid to Bottom of Upper Support Plate (180) Uncompensated Water 

Level 

LDP-139 Top of Lower Core Plate to Upper Core Spacer Grid Uncompensated Level 

LDP-140 Bottom of Reactor to Bottom of Flow Holes (180) Uncompensated Water Level 

LDP-141 Rx Vessel Upper Core Plate Uncompensated Level 

PT-101 CL-1 Pressure @ Reactor Flange 

PT-102 CL-2 Pressure @ Reactor Flange 

PT-103 CL-3 Pressure @ Reactor Flange 

PT-104 CL-4 Pressure @ Reactor Flange 

PT-107 Reactor Upper Head Pressure 

PT-108 Bottom of Reactor Pressure 

PT-109 DVI-1 Pressure @ Reactor Flange 

PT-110 DVI-2 Pressure @ Reactor Flange 

PT-111 Reactor Annular Pressure @ Flow Bypass Holes 

PT-112 Reactor Annular Pressure @ Bottom of Reactor 

PT-113 Reactor Pressure Below Mid-Core Spacer Grid 
 3-10 

3.3 Pressurizer (PZR)

A fully functional PZR, with internal heaters and a relief valve system capable of controlling 

the RCS pressure, has been included. The PZR is connected to HL-2 through the PZR surge line. 

The diameter of the PZR is not constant along its entire length. The larger diameter top 

was required for the APEX facility because there was not enough vertical space in the laboratory 

to accommodate the height of the PZR if the diameter remained constant (see Figure 3.5; 

note the notch in the beam to accommodate the taller PZR). The larger upper portion of the PZR 

is above the normal PZR water level, so only steam is present in the upper portion. The PZR 

in the APEX facility does not use a condensing spray for reduction of pressure; instead a vent 

is used to exhaust steam. A line from the first three stages of the ADS is connected to the top 

of the PZR for modeling the ADS depressurization of the primary system. The PZR has four heaters, 

and a T/C is installed in the disk end (top) of each heater. The PZR also has an air-operated valve 

to allow remote draining of the PZR. See Table 3.3 for a list of PZR-related instruments. 

Figure 3.5. Photograph of Pressurizer in the APEX Facility 
 3-11 

Table 3.3. Pressurizer Instrument Names and Locations 

LDP, HPS, DP, KW, and PT Cells Thermocouples 

Tag Name Description Tag 

Name Description 

DP-611 PZR Surge Line DP TF-601 PZR Surge Line @ PZR Inlet 

HPS-604-1 Lower PZR Surge Line Heat Transfer 

Coefficient 

TF-602 ADS1–3 Common Line @ PZR 

HPS-604-2 Lower PZR Surge Line Heater dT above 

Fluid Temperature 

TF-603 PZR Surge Line @ HL-2 

HPS-604-3 Lower PZR Surge Line Fluid Temperature TF-605 PZR Water Space Temperature 

HPS-606-1 ADS1–3 Common Inlet Heat Transfer 

Coefficient 

TF-608 PZR Fluid Temperature 

HPS-606-2 ADS1–3 Common Inlet Heater dT above 

Fluid Temperature 

TF-614 PZR Steam Vent Line 

Temperature 

HPS-606-3 ADS1–3 Common Inlet Fluid Temperature TH-601 PZR Heater Rod #1 

KW-601 PZR Heater Power TH-602 PZR Heater Rod #2 

LDP-601 PZR WR Uncompensated Water Level TH-603 PZR Heater Rod #3 

LDP-602 PZR Surge Line Uncompensated 

Water Level 

TH-604 PZR Heater Rod #4 

LDP-605 PZR Upper Surge Line Pipe 

Uncompensated Water Level 

LDP-606 PZR Surge Line Pipe Level @ PZR Inlet 

Uncompensated Water Level 

LDP-607 PZR Middle Surge Line Pipe 

Uncompensated Water Level 

LDP-608 PZR Lower Surge Line Pipe 

Uncompensated Water Level 

LDP-609 PZR Surge Line Pipe 

Uncompensated Water Level @ HL-2 

PT-602 PZR NR Pressure 

PT-603 PZR NR Pressure 

PT-604 PZR Pressure 
 3-12 

3.4 Steam Generators (SGs)

Two SGs, one on each loop, have been included in the APEX facility. Each SG is instrumented 

and has a tube and shell made to simulate a Westinghouse Delta-75 Steam Generator 

(see Figure 2-8). The SG lower channel head includes connections for two RCPs and a single HL. 

SG2 contains a line from the PRHR HX located inside the IRWST. The T/Cs are attached to 

the outer surface and are embedded into the walls of selected SG tubes. Each SG contains 

133 U-tubes. See Figures 3.6, 3.7, and 3.8 for images of the SG. 

Figure 3.6. Steam Generator Being Delivered to the APEX Facility 
 3-13 

Figure 3.7. Steam Generator U-Tube Bundle 

Figure 3.8. Steam Generator Components 
 3-14 

A power-operated relief valve (PORV), and associated piping, is installed on both SGs 

and is connected to the exhaust header. An air-operated ball valve provides remote actuation. 

A flow orifice is installed in each line to limit steam flow. Each line includes a T/C, a pressure 

transmitter, and a Vortex flow meter. To provide flow measurement within the desired range, 

the PORV line was enlarged and a Vortex flow meter was installed. The PORV line discharges 

into a common steam line, which is then routed into the steam exhaust stack. Table 3.4 lists 

the instruments associated with the secondary side of the SGs, while Table 3.5 lists the 

instruments associated with the primary side of the SGs. 

Table 3.4. Steam Generator Instrument Names and Locations (Secondary Side)

LDP and PT Cells Thermocouples 

Tag 

Name Description 

Tag 

Name Description 

LDP-301 SG1 WR 

Uncompensated Water Level TF-301 SG1 Steam Temperature 

LDP-302 SG2 WR 

Uncompensated Water Level TF-305 SG1 Downcomer HL Side Temperature 

LDP-303 SG1 NR 

Uncompensated Water Level TF-306 SG2 Downcomer HL Side Temperature 

LDP-304 SG2 NR 

Uncompensated Water Level TF-307 SG1 Downcomer CL Side Temperature 

PT-301 SG1 Pressure TF-308 SG2 Downcomer CL Side Temperature 

PT-302 SG2 Pressure TF-310 SG2 Steam Temperature 

TF-311 SG1 Feed Header Temperature 

TF-312 SG2 Feed Header Temperature 
 3-15 

Table 3.5. Steam Generator Instrument Names and Locations (Primary Side)

LDP, DP, and PT Cells Thermocouples 

Tag 

Name Description Tag Name Description 

DP-211 SG-1 Short Tube Entrance Losses TF-207 SG1 Short Tube @ Middle Outlet 

Side Temperature 

DP-212 SG2 Long Tube Exit Losses TF-208 SG2 Short Tube @ Middle Outlet 

Side Temperature 

DP-213 SG1 Long Tube Exit Losses TF-209 SG1 Short Tube @ Middle Inlet Side 

Temperature 

DP-214 SG2 Long Tube Entrance Losses TF-210 SG2 Short Tube @ Middle Inlet Side 

Temperature 

LDP-207 SG1 to HL1 Elbow Plenum 

Uncompensated Water Level 

TF-211 SG1 Long Tube @ Middle Outlet 

Temperature 

LDP-208 SG2 to HL2 Elbow Plenum 

Uncompensated Water Level 

TF-212 SG2 Long Tube @ Middle Outlet 

Temperature 

LDP-209 SG1 HL Plenum 

Uncompensated Water Level 

TF-213 SG1 Long Tube @ Middle Inlet 

Temperature 

LDP-210 SG2 CL4 Plenum 

Uncompensated Water Level 

TF-214 SG2 Long Tube @ Middle Inlet 

Temperature 

LDP-211 SG1 CL3 Plenum 

Uncompensated Water Level 

TF-215 SG1 Short Tube @ Top Temperature 

LDP-212 SG2 CL2 Plenum 

Uncompensated Water Level 

TF-216 SG2 Short Tube @ Top Temperature 

LDP-213 SG1 CL1 Plenum 

Uncompensated Water Level 

TF-217 SG1 Long Tube @ Top Temperature 

LDP-214 SG2 HL Plenum 

Uncompensated Water Level 

TF-218 SG2 Long Tube @ Top Temperature 

LDP-215 SG1 Long Tube HL 

Uncompensated Water Level 

TW-201 SG1 Short Tube Bottom Outlet 

LDP-216 SG2 Short Tube HL 

Uncompensated Water Level 

TW-202 SG2 Short Tube Bottom Outlet 

LDP-217 SG1 Short Tube HL 

Uncompensated Water Level 

TW-203 SG1 Short Tube Bottom Inlet 
 3-16 

Table 3.5. Steam Generator Instrument Names and Locations (Primary Side)

LDP, DP, and PT Cells Thermocouples 

Tag 

Name Description Tag Name Description 

LDP-218 SG2 Long Tube HL 

Uncompensated Water Level 

TW-204 SG2 Short Tube Bottom Inlet 

LDP-219 SG1 Long Tube CL 

Uncompensated Water Level 

TW-205 SG1 Long Tube Bottom Outlet 

LDP-220 SG2 Short Tube CL 

Uncompensated Water Level 

TW-206 SG2 Long Tube Bottom Outlet 

LDP-221 SG1 Short Tube CL 

Uncompensated Water Level 

TW-208 SG2 Long Tube Bottom Inlet 

LDP-222 SG2 Long Tube CL 

Uncompensated Water Level 

TW-209 SG1 Short Tube Top Outlet 

PT-201 SG1 Long Tube Pressure (Top) TW-210 SG2 Short Tube Top Outlet 

PT-204 SG2 Long Tube Pressure (Top) 

3.5 Reactor Coolant Pump (RCP)

Four variable-speed RCPs have been included as part of the RCS. The RCPs simulate 

the AP1000 canned motor pumps and are attached to the lower channel head of each SG. 

Each pump outlet is connected to a single CL. The pumps are fabricated from SS 304 

and can be programmed to simulate RCP coastdown. Figure 3.9 shows two RCPs mounted 

on SG2, while Figure 3.10 shows the pump head curve of the RCPs. See Table 3.6 for a list 

of the instrumentation related to the RCPs. 
 3-17 

Figure 3.9. Image of Two Reactor Coolant Pumps from Bottom 
 3-18 

Figure 3.10. Reactor Coolant Pump Head Curve 

Table 3.6. Reactor Coolant Pump Instrument Names and Locations

Differential Pressure Cells Thermocouples 

Tag Name Description Tag Name Description 

DP-202 RCP-2 DP TF-201 CL-1 @ RCP-1 Inlet Temperature 

DP-203 RCP-1 DP TF-202 CL-2 @ RCP-2 Inlet Temperature 

DP-205 RCP-3 DP TF-203 CL-3 @ RCP-3 Inlet Temperature 

DP-206 RCP-4 DP TF-204 CL-4 @ RCP-4 Inlet Temperature 
 3-19 

3.6 Primary Loop Piping

The primary loop piping models two primary loops, each consisting of a single HL and two CLs. 

Figure 3.11 shows the PZR side of the RCS. 

Figure 3.11. Primary Loop Piping in the Reactor Coolant System 

Break spool pieces are installed on the HLs, CLs, DVI lines, and CMT pressure balance line 

(PBL) to simulate several pipe break scenarios. The APEX facility is able to simulate 

the following primary loop breaks: 

• Top break on CL-3 

• Bottom break on CL-3 

• Bottom break on CL-4 

• Bottom break on HL-2 

• Single-ended DVI break 

• Single-ended CMT balance line break 

• Double-ended DVI break 

• Double-ended CMT balance line break 

Each of the primary loop breaks is initiated by pneumatic valves connected to the aforementioned 

locations. The break flow passes through an orifice plate, which is intended to set the size 

of the break, and may be changed out at anytime to simulate different break flow areas. 

The discharge from the break vents proceeds to the break and ADS measurement system 

(BAMS) to separate and measure the liquid and vapor components of the break flow rate. 
 3-20 

3.6.1 Hot Legs (HLs) 

In the primary side, each loop consists of a single pipe, which in turn, feeds the inverted U-tube SG 

mounted vertically on the bottom of each SG channel head. The HLs have connections 

for the fourth-stage ADS (ADS4). Refer to Table 3.7 for a list of instruments pertaining to the HL. 

Table 3.7. Hot Leg Instrument Names and Locations

Tag Name Description 

DP-209 HL2 DP 

DP-210 HL2 DP 

DP-216 HL Break DP 

DP-217 HL1 to CL1 DP @ SG1 

DP-218 HL2 to CL2 DP @ SG2 

DP-219 HL1 to CL3 DP @ SG1 

DP-220 HL2 to CL4 DP @ SG2 

DP-221 HL1 to CL1 DP @ RX 

DP-222 HL2 to CL2 DP @ RX 

DP-223 HL1 to CL3 DP @ RX 

DP-224 HL2 to CL4 DP @ RX 

HPS-205-1 HL1 Heat Transfer Coefficient 

HPS-205-2 HL1 Heater dT above fluid temperature 

HPS-205-3 HL1 Fluid temperature 

HPS-206-1 HL2 Heat Transfer Coefficient 

HPS-206-2 HL2 Heater dT above fluid temperature 

HPS-206-3 HL2 Fluid Temperature 

LDP-205 HL1 Uncompensated Water Level 

LDP-206 HL2 Uncompensated Water Level 

LDP-207 SG1 to HL1 Elbow Plenum Uncompensated Water Level 

LDP-208 SG2 to HL2 Elbow Plenum Uncompensated Water Level 

LDP-209 SG1 HL Plenum Uncompensated Water Level 

LDP-214 SG2 HL Plenum Uncompensated Water Level 

PT-202 HL2 Pressure @ SG2 Flange 

PT-205 HL1 Pressure @ SG1 Flange 

PT-206 HL Break Pressure @ Break Valve 
 3-21 

3.6.2 Cold Legs (CLs) 

In the primary side, each loop consists of two pipes, which in turn, feed the reactor vessel. 

The CL interconnects the reactor vessel and one of the two RCPs attached at the bottom of the SG. 

CL1 and CL3 have connections for the CMT-PBLs connected to the tops of CMT2 and CMT1, 

respectively. See Table 3.8 for a list of CL-related instrumentation. 

Table 3.8. Cold Leg Instrument Names and Locations

DP, LDP, HPS, FMM, and PT Cells Thermocouples 

Tag Name Description Tag Name Description 

DP-201 CL1 DP TF-201 CL1 @ RCP1 Inlet Temperature 

DP-204 CL2 DP TF-202 CL2 @ RCP2 Inlet Temperature 

DP-207 CL3 DP TF-203 CL3 @ RCP3 Inlet Temperature 

DP-208 CL4 DP TF-204 CL4 @ RCP4 Inlet Temperature 

DP-217 HL1 to CL1 DP @ SG1 TF-221 CL3 T/C Rod @ 3.25 in. (8.26 cm) 

DP-218 HL2 to CL2 DP @ SG2 TF-222 CL4 T/C Rod @ 3.25 in. (8.26 cm) 

DP-219 HL1 to CL3 DP @ SG1 TF-223 CL3 T/C Rod @ 2.50 in. (6.35 cm) 

DP-220 HL2 to CL4 DP @ SG2 TF-224 CL4 T/C Rod @ 2.50 in. (6.35 cm) 

DP-221 HL1 to CL1 DP @ RX TF-225 CL3 T/C Rod @ 1.75 in. (4.45 cm) 

DP-222 HL2 to CL2 DP @ RX TF-226 CL4 T/C Rod @ 1.75 in. (4.45 cm) 

DP-223 HL1 to CL3 DP @ RX TF-227 CL3 T/C Rod @ 1.25 in. (3.18 cm) 

DP-224 HL2 to CL4 DP @ RX TF-228 CL4 T/C Rod @ 1.25 in. (3.18 cm) 

FMM-201 CL1 Loop Flow TF-229 CL3 T/C Rod @ 0.75 in. (1.91 cm) 

FMM-202 CL2 Loop Flow TF-230 CL4 T/C Rod @ 0.75 in. (1.91 cm) 

FMM-203 CL3 Loop Flow TF-231 CL3 T/C Rod @ 0.25 in. (0.64 cm) 

FMM-204 CL4 Loop Flow TF-232 CL4 T/C Rod @ 0.25 in. (0.64 cm) 

HPS-201-1 CL1 Heat Transfer Coefficient TF-251-1 CL1 Loop Seal Upper Temperature 

HPS-201-2 CL1 Heater dT Above Fluid 

Temperature 

TF-251-2 CL1 Loop Seal Middle Temperature 

HPS-201-3 CL1 Fluid Temperature TF-251-3 CL1 Loop Seal Lower Temperature 

HPS-202-1 CL2 Heat Transfer Coefficient TF-252-1 CL2 Loop Seal Upper Temperature 

HPS-202-2 CL2 Heater dT Above Fluid 

Temperature 

TF-252-2 CL2 Loop Seal Middle Temperature 

HPS-202-3 CL2 Fluid Temperature TF-252-3 CL2 Loop Seal Lower Temperature 

HPS-203-1 CL3 Heat Transfer Coefficient TF-253-1 CL3 Loop Seal Upper Temperature 

HPS-203-2 CL3 Heater dT Above Fluid 

Temperature 

TF-253-2 CL3 Loop Seal Middle Temperature 
 3-22 

Table 3.8. Cold Leg Instrument Names and Locations

DP, LDP, HPS, FMM, and PT Cells Thermocouples 

Tag Name Description Tag Name Description 

HPS-203-3 CL3 Fluid Temperature TF-253-3 CL3 Loop Seal Lower Temperature 

HPS-204-1 CL4 Heat Transfer Coefficient TF-254-1 CL4 Loop Seal Upper Temperature 

HPS-204-2 CL4 Heater dT Above Fluid 

Temperature 

TF-254-2 CL4 Loop Seal Middle Temperature 

HPS-204-3 CL4 Fluid Temperature TF-254-3 CL4 Loop Seal Lower Temperature 

LDP-201 CL1 Water Uncompensated Level TF-255 CL1 Safety Injection Nozzle 

Temperature 

LDP-202 CL2 Water Uncompensated Level TF-256 CL2 Safety Injection Nozzle 

Temperature 

LDP-203 CL3 Uncompensated Water Level TF-257 CL3 Safety Injection Nozzle 

Temperature 

LDP-204 CL4 Uncompensated Water Level TF-258 CL4 Safety Injection Nozzle 

Temperature 

LDP-210 SG2 CL4 Plenum 

Uncompensated Water Level 

LDP-211 SG1 CL3 Plenum 

Uncompensated Water Level 

LDP-212 SG2 CL2 Plenum 

Uncompensated Water Level 

LDP-213 SG1 CL1 Plenum 

Uncompensated Water Level 

LDP-219 SG1 Long Tube CL 

Uncompensated Water Level 

LDP-220 SG2 Short Tube CL 

Uncompensated Water Level 

LDP-221 SG1 Short Tube CL 

Uncompensated Water Level 

LDP-222 SG2 Long Tube CL 

Uncompensated Water Level 

PT-203 CL Break Pressure @ Break Valve 
 3-23 

3.6.3 Pressurizer Surge Line 

The PZR surge line connects the bottom of the PZR to the top of HL2, near the SG end of the HL. 

The surge line enables continuous pressure adjustments between the RCS and the PZR. 

The PZR surge line spirals up from the HL to the PZR so that its geometry properly matches 

that of the AP1000 reactor (see Figure 3.12). 

Figure 3.12. Pressurizer Surge Line Geometry (Top View) 

3.7 Passive Safety Systems

3.7.1 Automatic Depressurization System (ADS) 1–4 

The ADS1–3 valves in the APEX facility model two trains in the AP1000 reactor. Each valve 

is located after a flow orifice that models the scaled choked flow area of the AP1000. The 

ADS1–3 valve train discharges to the ADS1–3 separator, where the two-phase flow is separated 

and measured. All portions of the flow are recombined before dumping into the ADS1–3 

sparger. 

The ADS1–3 valves exit at the top of the PZR through a single section of pipe. The pipe then 

splits in three, and is reduced prior to the ADS1, 2, and 3 flow nozzles. After the valves, 

the piping is recombined into a common pipe, which is connected to the ADS1–3 separator. 

The fluid is separated into its steam and liquid components inside the ADS1–3 moisture separator. 

The steam flow and liquid flow are measured and then recombined and dumped though a sparger 

into the IRWST. The sparger consists of four separate legs (see Figure 3.13). The ADS1 valve 

is opened on “S” signal, “S” signal plus a time delay, or CMT low level, depending on the selected 

logic. The ADS2 and ADS3 valves open based on a time delay after actuation of ADS1. 
 3-24 

Figure 3.13. Illustration of Automatic Depressurization System 1–3 Sparger 

Figure 3.14 shows the ADS1–3 sparger installed within the IRWST. Refer to Table 3.9 for a list 

of instrumentation pertaining to the ADS1–4 system. 
 3-25 

Figure 3.14. Photograph of Installed Sparger 

Table 3.9. Automatic Depressurization System 1–4 Instrument Names and Locations

LDP, FDP, HPS, FMM, and PT Cells Thermocouples 

Tag Name Description Tag Name Description 

FDP-604 ADS2 Flow Differential Pressure TF-602 ADS1–3 Common Line 

@ PZR Temperature 

FDP-605 ADS1 Flow Differential Pressure TF-609 ADS4–1 Discharge Temperature 

FDP-606 ADS3 Flow Differential Pressure TF-610 ADS4–2 Discharge Temperature 

FMM-601 ADS1–3 Loop Seal Flow TF-615 ADS1–3 Common Line 

from PZR Temperature 

FMM-602 ADS4–2 Loop Seal Flow TF-616 ADS1–3 Separator Loop Seal 

Temperature 

FMM-603 ADS4–1 Loop Seal Flow TF-617 ADS1–3 Separator Steam Outlet 

Temperature 

LDP-603 ADS4A Vertical Exit Line 

Uncompensated Water Level 

TF-618 ADS4–2 Loop Seal Temperature 

LDP-604 ADS4B Vertical Exit Line 

Uncompensated Water Level 

TF-619 ADS4–1 Loop Seal Temperature 
 3-26 

Table 3.9. Automatic Depressurization System 1–4 Instrument Names and Locations

LDP, FDP, HPS, FMM, and PT Cells Thermocouples 

Tag Name Description Tag Name Description 

LDP-610 ADS1–3 Separator 

Uncompensated Water Level 

TF-620 ADS4–2 Inlet From HL2 

Temperature 

LDP-611 ADS4–1 Separator 

Uncompensated Water Level 

TF-621 ADS4–1 Inlet From HL1 

Temperature 

LDP-612 ADS4–2 Separator 

Uncompensated Water Level 

TF-622 ADS4–2 Separator Steam Outlet 

Temperature 

PT-605 ADS1–3 Separator Pressure TF-623 ADS4–1 Separator Steam Outlet 

Temperature 

PT-606 IRWST Sparger Line Pressure TW-601 ADS1–3 Separator Wall 

Temperature 

PT-610 ADS4–2 Separator Pressure TW-602 ADS4–2 Separator Wall 

Temperature 

PT-611 ADS4–1 Separator Pressure TW-603 ADS4–1 Separator Wall 

Temperature 

HPS-606-1 ADS1–3 Common Inlet Heat Transfer 

Coefficient 

 

HPS-606-2 ADS1–3 Common Inlet Heater dT 

Above Fluid Temperature 

 

HPS-606-3 ADS1–3 Common Inlet Fluid 

Temperature 

 

Two ADS4 valves are connected to the top of HL1 and HL2, respectively. Both ADS4 lines 

are similar, with the exception that ADS4–2 (PZR side) provides a tee connection for the PRHR 

HX system. Both ADS4 lines include valve flow area and line resistance. The flow from ADS4 

is directed to a an orifice to match the separator, where the steam and liquid flows are 

separated and measured individually. The steam flow is routed through one of two lines 

depending on the expected flow rate and is then vented from the facility. The liquid flow is 

directed through a loop seal and into the primary sump. 

3.7.2 In-Containment Refueling Water Storage Tank (IRWST) 

The APEX IRWST is constructed of SS 304 and has four major connections, including 

the ADS1–3 sparger, DVI1 and 2, the primary sump, and the PRHR connections. Table 3.10 

lists the pressure, weight, and flow rate instruments for the IRWST, while Table 3.11 lists 

the temperature measurement instrumentation. The temperature of the fluid in the IRWST 

is measured at several heights to determine stratification. 
 3-27 

Table 3.10. IRWST LDP, LCT, FMM, DP, and PT Cells and Locations

Tag Name Description 

DP-701 IRWST-1 Injection Differential Pressure 

DP-702 IRWST-2 Injection Differential Pressure 

FMM-701 IRWST-1 Injection Flow 

FMM-702 IRWST-2 Injection Flow 

FMM-703 IRWST Overflow 

LDP-701 IRWST Uncompensated Water Level 

PT-701 IRWST Pressure 

LCT-701 IRWST Weight 
 3-28 

Table 3.11. IRWST Thermocouples and Location

Tag Name Description Tag Name Description 

TF-701 IRWST/PRHR T/C Rod @ Bottom 

Temperature 

TF-713 IRWST Discharge to DVI-01 

@ IRWST Temperature 

TF-702 IRWST/PRHR T/C Rod @ 7.98 in. 

(20.27 cm) 

TF-714 IRWST Discharge to DVI-02 

@ IRWST Temperature 

TF-703 IRWST/PRHR T/C Rod @ 15.97 in. 

(40.56 cm) 

TF-715 IRWST Sparger T/C Rod @ 8.97 in. 

(22.78 cm) 

TF-704 IRWST/PRHR T/C Rod @ 25.85 in. 

(65.67 cm) 

TF-716 IRWST Sparger Nozzle Temperature 

TF-705 IRWST/PRHR T/C Rod @ 35.73 in. 

(90.75 cm) 

TF-717 IRWST Sparger T/C Rod @ 66.34 in. 

(168.50 cm) 

TF-706 IRWST/PRHR T/C Rod @ 45.61 in. 

(115.85 cm) 

TF-718 IRWST Sparger T/C Rod @ 98.45 in. 

(250.06 cm) 

TF-707 IRWST/PRHR T/C Rod @ 55.49 in. 

(140.94 cm) 

TF-719 IRWST Sparger Outlet Temperature 

TF-708 IRWST/PRHR T/C Rod @ 65.36 in. 

(166.01 cm) 

TF-720 IRWST/DVI2 Injection Line 

Temperature 

TF-709 IRWST/PRHR T/C Rod @ 75.24 in. 

(191.11 cm) 

TF-721 IRWST/DVI1 Injection Line 

Temperature 

TF-710 IRWST/PRHR T/C Rod @ 86.36 in. 

(219.35 cm) 

TF-722 IRWST Steam Exhaust Line 

Temperature 

TF-711 IRWST/PRHR T/C Rod @ 97.47 in. 

(247.57 cm) 

TF-723 IRWST/Primary Sump Overflow 

Temperature 

TF-712 IRWST/PRHR T/C Rod @ 108.59 in. 

(275.82 cm) 

The two IRWST injection lines that connect into each DVI line start at the bottom of the IRWST, 

which in turn, connects to the SS 304 pipe connected to the DVI line. The IRWST injection lines 

are similar, with the exception that IRWST-1 provides a connection to the RNS pump suction. 

The IRWST injection line resistance is fine-tuned by an orifice plate installed in each line. 

The IRWST isolation valves (RCS-711/712) are opened on decreasing RCS pressure. See 

Figure 3.15 for an image of the IRWST. 
 3-29 

Figure 3.15. Photograph of the IRWST 

3.7.3 Core Makeup Tanks (CMT) 

The CMTs are constructed of SS 304, and each CMT is connected to a CL and a DVI line 

by SS 304 pipe. The CMT injection line isolation valves (RCS-529/530) open on “S” signal 

for the station blackout subroutine or “S” signal plus a time delay for all other routines. The line 

resistance of the CMT injection lines is fine-tuned through the use of an orifice. The isolation 

valves in the CMT balance line (RCS-535/536) have no automatic control and remain open 

during facility testing. See Figure 3.16 for an image of a CMT. Table 3.12 lists all CMT-related 

instruments and piping. 
 3-30 

Figure 3.16. Photograph of a Core Makeup Tank 

Table 3.12. Core Makeup Tank Instrument Names and Locations

LDP, FMM, HPS, DP, and PT Cells Thermocouples 

Tag Name 

Description Tag 

Name 

Description 

DP-501 CMT1 Injection Differential 

Pressure 

TF-501 CMT1 Long T/C Rod @ 0.30 in. (0.76 cm) 

Temperature 

DP-502 CMT2 Injection DP TF-502 CMT2 Injection Line Temperature 

DP-503 CMT1 Balance Line DP TF-503 CMT1 @ 1⁄2 Lower Head Height 

DP-504 CMT2 Balance Line DP TF-504 CMT2 Long T/C Rod @ 0.30 in. (0.76 cm) 

Temperature 

FMM-501 CMT1 Injection Flow TF-505 CMT1 @ 20% Volume-Height 

FMM-502 CMT2 CL Balance Line Flow TF-506 CMT2 @ 1⁄2 Lower Head Height 

FMM-503 CMT1 CL Balance Line Flow TF-507 CMT1 Long T/C Rod @ 20.87 in. (53.01 cm) 

Temperature 
 3-31 

Table 3.12. Core Makeup Tank Instrument Names and Locations

LDP, FMM, HPS, DP, and PT Cells Thermocouples 

Tag Name 

Description Tag 

Name 

Description 

FMM-504 CMT2 Injection Flow TF-508 CMT2 @ 20% Volume-Height 

LDP-501 CMT1 NR Uncompensated 

Water Level (Bottom) 

TF-509 CMT1 Long T/C Rod @ 36.89 in. 

(93.70 cm) Temperature 

LDP-502 CMT2 WR Uncompensated 

Water Level 

TF-510 CMT2 Long T/C Rod @ 20.87 in. 

(53.01 cm) Temperature 

LDP-503 CMT1 NR Uncompensated 

Water Level (Middle) 

TF-511 CMT1 @ 50% Volume-Height 

LDP-504 CMT2 NR Uncompensated 

Water Level (Bottom) 

TF-512 CMT2 Long T/C Rod @ 36.89 in. 

(93.70 cm) Temperature 

LDP-505 CMT1 NR Uncompensated 

Water Level (Top) 

TF-513 CMT1 Long T/C Rod @ 40.59 in. 

(103.10 cm) Temperature 

LDP-506 CMT2 NR Uncompensated 

Water Level (Middle) 

TF-514 CMT2 @ 50% Volume-Height 

LDP-507 CMT1 WR Uncompensated 

Water Level 

TF-515 CMT1 Long T/C Rod @ 43.41 in. 

(110.26 cm) Temperature 

LDP-508 CMT2 NR Uncompensated 

Water Level (Top) 

TF-516 CMT2 Long T/C Rod @ 40.59 in. 

(103.10 cm) Temperature 

LDP-509 CL3 to CMT1 Balance Line 

Uncompensated Water Level 

TF-517 CMT1 @ 75% Volume-Height minus 3.7 in. 

(9.40 cm) 

LDP-510 CL1 to CMT2 Balance Line 

Uncompensated Water Level 

TF-518 CMT2 Long T/C Rod @ 43.41 in. 

(110.26 cm) Temperature 

PT-501 CMT1 Pressure TF-519 CMT1 Long T/C Rod @ 46.23 in. 

(117.42 cm) Temperature 

PT-502 CMT-02 Pressure TF-520 CMT2 @ 75% Volume-Height minus 3.7 in. 

(9.40 cm) 

HPS-509-1 CMT-01 CL Balance Line Heat 

Transfer Coefficient 

TF-521 CMT1 @ 75% Volume-Height 

HPS-509-2 CMT-01 CL Balance Line Heater 

dT above fluid temperature 

TF-522 CMT2 Long T/C Rod @ 46.23 in. 

(117.42 cm) Temperature 

HPS-509-3 CMT-01 CL Balance Line Fluid 

temperature 

TF-523 CMT1 Long T/C Rod @ 49.05 in. 

(124.59 cm) Temperature 

HPS-512-1 CMT-02 CL Balance Line Heat 

Transfer Coefficient 

TF-524 CMT 

HPS-512-2 CMT-02 CL Balance Line Heater 

dT above fluid temperature 

TF-525 CMT1 @ 1/2 Upper-Head Height 

HPS-512-3 CMT-02 CL Balance Line Fluid 

temperature 

TF-526 CMT2 SPARGER 2\3 TEMP 
 3-32 

Table 3.12. Core Makeup Tank Instrument Names and Locations

LDP, FMM, HPS, DP, and PT Cells Thermocouples 

Tag Name 

Description Tag 

Name 

Description 

TF-530 CMT2 Long T/C Rod @ 51.87 in. 

(131.75 cm) Temperature 

TF-527 CMT1 Long T/C Rod @ 51.87 in. 

(131.75 cm) Temperature 

TF-531 CMT1 Balance Line @ CMT Inlet 

Temperature 

TF-528 CMT 23% Head Temp 

TF-532 CMT2 Long T/C Rod @ 56.61 in. 

(143.79 cm) Temperature 

 

TF-533 CMT1/CL Balance Line @ CL-3 

Temperature 

TF-542 CMT2 @ 60% Volume-Height 

TF-535 CMT1 Injection Line 

Temperature 

TF-543 CMT1 @ 75% Volume-Height 

TF-536 CMT2/CL Balance Line @ CL-1 

Temperature 

TF-544 CMT2 @ 75% Volume-Height 

TF-537 CMT1 @ 20% Volume-Height TF-546 CMT2 Balance Line @ CMT Inlet 

Temperature 

TF-538 CMT2 @ 20% Volume-Height TF-547 CMT1 Long T/C Rod @ 54.24 in. 

(137.77 cm) Temperature 

TF-539 CMT1 @ 50% Volume-Height TF-548 CMT2 Long T/C Rod @ 54.24 in. 

(137.77 cm) Temperature 

TF-540 CMT2 @ 50% Volume-Height TF-549 CMT1 Discharge Line Temperature 

TF-541 CMT1 @ 60% Volume-Height TF-550 CMT2 Discharge Line Temperature 

 TF-551 CMT1 Short T/C Rod (225 degrees) 5.5 in. 

(13.97 cm) from top 

3.7.4 Passive Residual Heat Remover (PRHR) Heat Exchanger (HX) 

The PRHR HX consists of a C-shaped heat exchanger, located inside the IRWST, and two 

connections to the primary loop. One inlet connection is between the HL2 ADS4 tee 

and the upper PRHR HX manifold on the IRWST. The outlet connection is between the lower 

PRHR HX manifold on the IRWST and the outlet plenum of SG2. The PRHR flow is actuated by 

an isolation valve (RCS-804) that opens on “S” signal, actuation of the ADS1 valve, or CMT 

injection valve opening, depending on the selected logic subroutine. The PRHR HX is connected 

to the upper and lower heat exchanger manifolds on the IRWST. It is made up of 88 individual tubes. 

See Figure 3.17 for an image of the PRHR HX. Table 3.13 contains a list of all instruments 

related to the PRHR. 
 3-33 

Figure 3.17. View of the PRHR HX Located Inside the IRWST 
 3-34 

Table 3.13. PRHR Instrument Names and Locations

LDP, FMM, and HPS Thermocouples 

Tag Name Description Tag Name Description 

FMM-802 PRHR Inlet Flow TF-803 PRHR HX Inlet Temperature 

FMM-804 PRHR Outlet Flow TF-804 PRHR HX Outlet Temperature 

LDP-801 PRHR HX Inlet Head 

Uncompensated Water Level 

TF-805 PRHR HX Long Tube Outlet at Bend 

Temperature 

LDP-802 PRHR HX WR 

Uncompensated Water Level 

TF-806 PRHR HX Short Tube Outlet at Bend 

Temperature 

HPS-801-1 PRHR HX Inlet Heat Transfer 

Coefficient 

TF-808 PRHR HX Short Tube Outlet at Center 

Temperature 

HPS-801-2 PRHR HX Inlet Heater dT 

Above Fluid Temperature 

TF-809 PRHR HX Long Tube at Center 

Temperature 

HPS-801-3 PRHR HX Inlet Fluid Temperature TF-810 PRHR HX Short Tube Inlet at Bend 

Temperature 

TF-811 PRHR HX Long Tube Inlet at Bend 

Temperature 

TF-812 PRHR HX Outlet Head Temperature 

TW-801 PRHR HX Long Tube Outlet 

TW-802 PRHR HX Short Tube Outlet 

TW-803 PRHR HX Long Tube Lower Mid-piece 

TW-804 PRHR HX Short Tube Lower Mid-piece

TW-805 PRHR HX Short Tube Upper Mid-piece

TW-806 PRHR HX Long Tube Upper Mid-piece 

TW-807 PRHR HX Short Tube Inlet 

TW-808 PRHR HX Long Tube Inlet 

3.7.5 Accumulators (ACC) 

The two ACC tanks are constructed of a long piece of pipe with a pipe cap and a thick plate. 

The plate is cut to fit in one end of the pipe, and a pipe cap type end is used to seal the other end. 

Table 3.14 lists the instruments associated with the accumulators. 
 3-35 

Table 3.14. Accumulator Instrument Names and Locations

LDP, FMM, DP, and PT Cells Thermocouples 

Tag Name Description Tag Name Description 

DP-401 ACC1 Injection DP TF-401 ACC1 Outlet Temperature 

DP-402 ACC2 Injection DP TF-402 ACC2 Outlet Temperature 

FMM-401 ACC1 Injection Flow TF-403 ACC1 N2 Top Temperature 

FMM-402 ACC2 Injection Flow TF-404 ACC2 N2 Top Temperature 

LDP-401 ACC1 Uncompensated Water Level TF-405 ACC1 Injection Line Temperature 

LDP-402 ACC2 Uncompensated Water Level TF-406 ACC2 Injection Line Temperature 

PT-401 ACC1 Pressure 

PT-402 ACC2 Pressure 

Each ACC is connected to a DVI line, and the fluid flow initiates from an ACC to its respective 

DVI line by means of a check valve (RCS-423/424), which opens when plant pressure becomes 

less than the pressure in the ACC. See Figure 3.18 for an image of an ACC. 

Figure 3.18. Photograph of One of the APEX Accumulators 
 3-36 

3.7.6 Containment Sump 

The primary containment sump tank is constructed of SS 304 plate, and the cylindrical section 

is capped with elliptical ends. The primary sump tank is connected to the ADS4 separators, 

the secondary sump, the break separator, DVI1 and 2, and the steam exhaust header. 

The sump collects all liquid flow from the break and ADS4 lines. Two containment sump 

recirculation lines are connected to the bottom of the primary sump and provide a recirculation path 

to the RPV via the DVI lines. Each line consists of an isolation valve with a parallel line 

with check valves. The isolation valves (CSS-909/910) open on decreasing IRWST level. 

A secondary sump is also provided to model containment sump volumes that are not available 

for long-term circulation. A curb overflow is simulated between the primary and secondary sumps 

by a weir plate. Table 3.15 lists the instruments associated with the primary and secondary sumps. 

Table 3.15. Sump Instrument Names and Locations

LDP, FMM, FVM, LCT, and PT Cells Thermocouples 

Tag 

Name Description Tag Name Description 

FMM-901 Primary Sump-1 Recirculation 

Injection Flow 

TF-901 Primary Sump Inlet from Fill Line 

Temperature 

FMM-902 Primary Sump-2 Recirculation 

Injection Flow 

TF-902 Secondary Sump Bottom Temperature 

FMM-905 Break Separator Loop Seal Flow TF-903 Primary Sump Bottom Temperature 

FVM-903 Primary Sump Steam Exhaust 

Flow 

TF-904 Primary Sump/DVI2 Injection Line 

Temperature 

LDP-901 Primary Sump 

Uncompensated Water Level 

TF-905 Primary Sump @ Secondary Sump 

Crossover Level Temperature 

LDP-902 Secondary Sump 

Uncompensated Water Level 

TF-906 Primary Sump Exhaust Temperature 

LDP-903 CRT Uncompensated Water Level TF-907 Primary Sump @ Top Temperature 

LDP-905 Break Separator 

Uncompensated Water Level 

TF-909 Primary Sump/DVI1 Injection Line 

Temperature 

PT-901 Primary Sump Pressure TF-910 CRP Discharge to Primary Sump Temp 

LCT-901 Primary Sump Weight TF-911 CRP Discharge to IRWST Temperature 

LCT-902 Secondary Sump Weight TF-912 Break Separator Loop Seal 

Temperature 

TF-913 Break Separator Steam Outlet 

Temperature 

TF-914 CRT Outlet Temperature 
 3-37 

3.8 Instrumentation, Control, and Power Systems

This section describes the instrumentation, control, and power systems of the APEX test facility, 

as well as the data acquisition system, instrument devices, and associated control logic. 

3.8.1 Instrumentation Description 

The APEX test facility is equipped with approximately 600 instrument channels to capture 

the transient behavior of the safety systems. The APEX facility includes the following types 

of instrumentation channels: 

• Thermocouples (TF/TR/TH/TW) are used to measure fluid, heater, and wall temperatures. 

Premium thermocouples with controlled purity thermocouple wire were incorporated. 

• Magnetic flow meters (FMMs) are incorporated to measure single-phase liquid flows around 

the plant. 

• Vortex shedding flow meters (FVMs) are used to measure single-phase vapor flow rates. 

• Pressure transducers (PT) are used to measure the static pressures in various tanks 

and pipes within the facility. 

• Differential pressure (LDP, DP) transducers are used to measure liquid levels in various 

tanks and pipes. The differential pressure cells are also used to measure pressure drops. 

• Heated phase switches (HPSs) are used to determine the fluid phase at various points 

inside system piping. Each HPS measures fluid temperature, change in temperature (∆T) 

between the fluid and the heater, and a relative heat transfer coefficient. 

• Load cell transducers (LCTs) are incorporated to measure the liquid mass inside the IRWST, 

primary sump, and secondary sump. 

• Heater power (KW) to the core is measured with four power meters. The power applied 

to the PZR heaters is measured by an additional power meter. 

• A level transducer (LT) capacitance probe is used to measure collapsed liquid level 

in the reactor vessel. 
 3-38 

Instrumentation channels are subdivided into 10 different groups based on a particular location, 

using the numbering system described in Table 3.16. 

Table 3.16. Instrument Tag Locations

Channel Name Location 

xx-000s Feed, Main Steam 

xx-100s Reactor Pressure Vessel 

xx-200s Primary Loop 

xx-300s SG 

xx-400s ACC 

xx-500s CMTs 

xx-600s PZR, ADS1–3 

xx-700s IRWST 

xx-800s PRHR, CVS, RNS 

xx-900s Sump, ADS4-1, ADS4-2, and Break System 

3.8.2 Data Acquisition 

The data acquisition system (DAS) writes the data into a single large database, as shown in Figure 

3.19. Using a single database (rather than individual test files) offers the following advantages: 

• improved management and maintenance 

• enables retrieval to be independent of storage (i.e., the data are stored every second, 

but the database can be queried at another resolution, say every 5 seconds), which 

provides a very powerful and flexible retrieval capability 

• easier comparisons between multiple tests 

The DAS is manufactured by National Instruments (NI) and uses commercial off-the-shelf software. 

The signal type may be a 4–20 mA current signal, 1–5 VDC or 0–10 VDC voltage signal, or the 

millivolt signal originating from thermocouples. The instruments are connected to several terminal 

boards. Each terminal board can accept any combination of J-, K-, and T-type thermocouples and 

up to 32 channels of inputs with their own cold junction compensation. These terminal boards are 

also used for other types of signal inputs. Each terminal board is connected to one analog input 

module, which rests within a chassis that contains 12 modules. The DAS use three chassis to 

acquire the data. The signals from the chassis are then transferred to a computer that is 

equipped with a multifunction I/O board. Basically, the multifunction I/O boards switch their 

inputs between 622 channels at high speed to sample all input channels. The data for APEX 

are typically sampled at a rate of 1 Hertz per channel. 
 3-39 32 Signals per board

 Chassis

12 Modules

Chassis

Chassis

Computer with multifunction I/O board

622 Signals Total

DAS Cabinet

Figure 3.19. Data Acquisition System Hardware Overview 

All of the data from input channels are acquired by NI-DAQ hardware/software and then 

converted into engineering units. The real-time data are sampled and stored in an NI Citadel 

database at one time per second per channel. Citadel is a proprietary database used for 

process data. OSU developed an export program to query the Citadel database and format 

the data file in ASCII text and NRC Data Bank binary file formats. 

The NI DIAdem 8.1 trending software package is used to plot data for generating reports. 

DIAdem is capable of manipulating large data sets of more than a billion data points in up to 

65,000 columns. 
 3-40 

3.8.3 Control System and Control Logic 

The APEX test facility control logic system includes various field process transmitters, operator 

switches, an OMRON programmable logic controller (PLC), a set of Fischer & Porter process 

controllers (PCs), and a supervisory host computer, as shown in Figure 3.20. 

PLC

Digital

I/O Digital I/O

RS-232

Supervisory host computer

Analog

Signals

Binary

Signals

RS-232 Control Panel

To Hardware Logic Control Unit User Interface

Process

Controllers

Figure 3.20. Control System Interface 
 3-41 

The control system functions in three basic parts: 

(1) The PLC performs all binary logic functions for safety, sequencing, and operational control. 

(2) The PCs perform all dynamic (analog) process controls for smooth operation of variable 

control devices. 

(3) The supervisory system provides a graphical computer interface between the operator 

and the facility. 

The PLC consists of a central processor for program execution, a power supply, two local racks 

of digital I/O, and several remote I/O racks located throughout the facility. All inputs are sensed 

as 24-VDC signals. All outputs are driven by a 24-VDC supply to energize pneumatic valves, 

pump motor starters, control panel lights and alarms, and other components. Figure 3.21 is 

photograph of the APEX control panel. 

Figure 3.21. Photograph of the APEX Control Panel 

The process control system consists of seven dedicated PCs with a total of 40 analog inputs 

and 20 analog outputs. Analog inputs are used to monitor the process, either directly or as 

calculated variables. They include tank level, system pressure, pump flow rate, and applied 

heater power. The controller performs a series of calculations to determine the likely state 

of the process and adjusts the analog output of the controlling device to maintain the process 

within defined parameters. The controlling devices include valves, power silicon control 

rectifiers (SCRs), and variable-speed pump controllers. The DAS and control computer 

are connected by a high-speed local area network. 
 3-42 

There are six test logic routines implemented in APEX logic: 

(1) Station Blackout Routine 

(2) Small-Break Loss-of-Coolant Accident (LOCA) Routine 

(3) Large-Break LOCA Routine 

(4) Inadvertent ADS Initiation Routine 

(5) Full-Pressure ADS Routine 

(6) Natural Circulation Routine 

3.8.4 Power System 

The main purpose of the power system is to provide up to 1 megawatt of power to the reactor 

and PZR heater rods. Power comes from the local power grid at a voltage of 20,800 volts 

and is stepped down to a distribution voltage of 480 volts, three phase. From the 480-volt 

distribution power, one line (three-phase, max current ≈100A) is used for the pumps in the 

APEX facility (RCP1 to RCP4, MFP, CVSP, RNSP, CRP, and seal WTR fan). Another line 

(three-phase, max current ≈200A) provides the power for trace heaters. The three remaining 

lines are for the two Rx heater rod groups (three-phase, 500 kW max per line) and the PZR heater 

(three-phase, 20 kW max). Each of the heater power lines has the following components: 

• 1 breaker to manually turn the power on/off 

• 1 fuse for circuit protection 

• 1 contactor to remotely turn the power on/off 

• 1 phase angle-fired SCR to adjust the power level (SCR1 and SCR2 for Rx heater groups 1 

and 2, SCR3 for the PZR heater) 

During the test, the contactor controls the heater power (on/off), and the SCR signals are used 

to adjust the heater power level. The reactor heater rod SCRs, in conjunction with the heater 

controller, allow the APEX facility to simulate decay power by continually adjusting the reactor 

power level setpoint to follow the desired decay curve. See Figure 3.22 for a photograph of the 

SCR cabinet, and Figure 3.23 for a simplified sketch of the main power distribution system. 
 3-43 

Figure 3.22. Photograph of the SCR Cabinet 

SCR-3

SCR-2 SCR-1

20800V

3φ AC

Transformer

480V

3φ AC

≈1MW

Motors

Breaker Cabinet

Contactors

SCR Cabinet

To PZR heater

To RX Heater #2

To RX Heater #1

Figure 3.23. Main Power System Simple Layout 
 3-44 

3.8.5 Break and ADS Measurement System (BAMS) 

The BAMS, depicted in Figure 3.24, is used to measure two-phase volumetric flow rates from 

the break(s) and four stages of the ADS. All of the steam lines from the break separator, 

ADS4 separators, IRWST, and primary sump feed into the BAMS header and eventually vent 

via the exhaust header. All of the steam lines and moisture separators are heated with strip 

heaters to prevent condensation. Steam volumetric flow rate, pressure, and temperature 

are measured by vortex flow meters, pressure cells, and thermocouples respectively. Table 

3.17 lists all instruments related to the ADS4-1/4-2 and break separator. 

The BAMS system includes the following components: 

• break separator 

• two ADS4 separators 

• ADS1–3 separator 

 IRWST

Break

Separator

SUMP

ADS 4-1 ADS 4-2

ADS 1-3

Breaks

BAMS Header EXHAUST

Liquid line Steam line Two phase line

Separator

Separator

Separator Sump

Injection

Figure 3.24. BAMS Layout 
 3-45 

Table 3.17. ADS4-1/4-2, and Break Separator Instrument Names and Locations

LDP, FVM, FMM, DP, and PT Cells Thermocouples 

Tag 

Name Description Tag Name Description 

DP-905 Break Separator Entrance DP TF-908 Break Separator Inlet Temperature 

FMM-905 Break Separator Loop Seal Flow TF-910 CRP Discharge to Primary Sump 

Temperature 

FVM-901 BAMS HDR 6 in. Line Steam Flow TF-911 CRP Discharge to IRWST Temperature 

FVM-902 BAMS HDR 10 in. Line Steam Flow TF-912 Break Separator Loop Seal Temperature 

FVM-904 Break Separator 3 in. Line Steam Flow TF-913 Break Separator Steam Outlet 

Temperature 

FVM-905 Break Separator 6 in. Line Steam Flow TF-914 CRT Outlet Temperature 

FVM-906 Break Separator 8 in. Line Steam Flow TF-915 Break Separator 6 in. Steam Line 

Temperature 

LDP-903 CRT Uncompensated Water Level TF-916 BAMS Header 10 in. Steam Line 

Temperature 

LDP-905 Break Separator 

Uncompensated Water Level 

TF-917 BAMS Header Temperature 

PT-902 BAMS Header Pressure TF-918 Break Separator 8 in. Steam Line 

Temperature 

PT-905 Break Separator Pressure TW-905 Break Separator Wall Temperature 

The APEX test facility can model several different break locations: 

• CL3 bottom break (TS-205 actuates) 

• CL3 top break (TS-201 & T-205 both actuates) 

• single-ended DVI break (TS-202 actuates, pressure equalization line open during warm-up, 

closed prior to break) 

• single-ended CMT balance line break (TS-203 actuates, pressure equalization line open 

during warm-up, closed prior to break) 

• HL2 bottom break (TS-204 actuates) 

• CL4 bottom break (TS-206 actuates) 

• double-ended CMT balance line break (TS-203 and TS-202 actuates) 

• DEDVI break (TS-203 and TS-202 actuates) 

The two-phase flow that exits the break is piped into the break separator. The primary purpose 

of the break separator is to separate the steam and liquid phase of two-phase flow for the 

convenience of measurement. Water/steam mixture enters through the middle and passes 

through a cyclone separator. The separated steam exits the top of the separator, while the liquid 

is collected and allowed to drain out of the bottom of the separator through a loop seal and into 

the primary sump. The wall of the separator is heated by a long strip heater to prevent steam 

condensation inside the separator. 
 3-46 

The break separator steam outlet line is divided from a large pipe into three smaller parallel paths, 

which are recombined before the exhaust. The steam flow is routed through one of these three paths, 

depending on the expected flow rate. Each of these paths is controlled by an air-operated logic 

ball valve and measured by a vortex flow meter to control the separator liquid level within 

the desired range to contain the steam, while the liquid goes to the correct outlet. If the steam line 

size is too small for a given flow rate, the increased pressure will push the water out of the loop 

seal, and there will be a two-phase mixture in the liquid line (which cannot be measured). By 

contrast, if the selected steam line is too large, it is difficult to obtain an accurate measure of the 

steam flow rate because it is too small. The ADS4 separator steam line has a similar design, 

but uses only two paths. The primary and secondary sumps are used to simulate 

the containment. 

The two ADS4 separators are similar in design to the break separator. The steam exiting the 

separators can be directed through one of two different sized pipes, depending on the expected 

flow rate. The reasoning behind the use of different size steam lines is the same as for the flow 

from the break separator. The liquid phase from each ADS4 separator is collected and directed 

through a loop seal and into the primary sump. 

The ADS1–3 separator is a large cylinder. The exiting two-phase mixture flow from ADS1–3 

valves is directed into the ADS1–3 separator through an inlet. The liquid flow exits the bottom 

of the separator, while the steam flow exits the separator via a different pipe located at the 

opposite side of the inlet. The steam volumetric flow rate is measured by a vortex flow meter 

(FVM-601). Liquid phase is directed out of the separator via a pipe at the bottom of the separator 

and measured by a magnetic flow meter. The steam and liquid phases are finally recombined 

and injected into the IRWST via a sparger. 

In the prototype, the steam phase that exits the break condenses on the containment wall 

and drains back to the IRWST or sump. In the test facility, the steam phase is vented to 

the atmosphere. To compensate for the mass loss attributable to steam exhaust to the atmosphere, 

the containment condensate return system is available to inject fluid into the IRWST or sump 

at the same mass flow rate of the vented steam if desired. The condensate return system is fed 

by the feedwater storage tank (FST) into the condensate return tank (CRT) in which the water 

is preheated. The CRT contains heaters to preheat the water prior to injection into the IRWST 

or the primary sump. The outlet water temperature is measured by a thermocouple, 

and the measurement is feedback to the CRT heaters to maintain the proper outlet temperature. 

The liquid is pumped from the CRT by means of the condensate return pump (CRP) through either 

a recirculation line with the line resistance controlled by an orifice plate, or the IRWST or primary sump 

via motor-operated valves (CCS-928/927). 
 3-47 

3.9 Balance of Plant

The APEX-AP1000 facility incorporated several of the AP1000 non-safety systems, including 

the CVS, RNS, SG, feedwater system, main steam system, and water purification system. 

These non-safety systems were designed and constructed to meet the necessary functional 

requirements for testing. The pipe routings and relative locations within the APEX facility 

are not necessarily prototypic of the AP1000 design. 

3.9.1 Chemical and Volume Control System (CVS) 

The CVS is used to maintain proper water chemistry in the primary coolant for a full-sized plant. 

The CVS pump is a Goulds 3-stage centrifugal pump. See Table 3.18 for a list of instruments 

related to the CVS. In the APEX facility, the CVS pump is used for the following five main purposes: 

(1) maintenance of fluid during startup 

(2) modeling AP1000 operation if desired 

(3) normal supply of makeup water to the APEX facility 

(4) hydrostatic testing 

(5) backup pump for the main feedwater pump (MFP) 

Table 3.18. CVS Instrument Names and Locations

FMM, PT Cells, and Thermocouples 

Tag Name Description 

FMM-801 CVS Discharge Flow 

PT-801 CVS Discharge Pressure 

TF-801 CVS Discharge Temperature 

3.9.2 Normal Residual Heat Removal System (RNS) 

The RNS pump draws water from the IRWST or primary sump, injects it into the DVI line, 

and provides makeup water under low-pressure conditions. The RNS pump is used to fill 

the CMTs and ACCs before performing a test. See Table 3.19 for a list of instruments related to 

the RNS. 
 3-48 

Table 3.19. RNS Instrument Names and Locations

FMM and PT Cells Thermocouples 

Tag Name Description Tag Name Description 

FMM-803 RNS to DVI2 Flow TF-802 RNSP Discharge Temperature 

FMM-805 RNS Discharge Flow TF-813 RNSP Discharge to DVI1 Temperature 

PT-802 RNS Discharge Pressure TF-814 RNSP Discharge to DVI2 Temperature 

Figure 3.25 depicts the performance curve for the RNS pump [CR4-100N (5.0 HP (3.73 kW))]. 

Figure 3.25. RNS Pump Performance Curves 
 3-49 

3.9.3 Feedwater System 

The purpose of the feedwater system is to provide water to the SG. The feedwater system 

consists of the demineralization system (DI), FST, and MFP, as shown in Figure 3.26. The FST 

stores water from the DI. Table 3.20 lists all instruments related to the FST, MFP, and SG 

steam vent. 

For the main feed flow path, the MFP takes water from the FST and pumps it into both SGs. 

The MFP is a Goulds 3-stage centrifugal pump. The pump curve is at the end of this section. 

There is a magnetic flow meter (FMM-001 or FMM-002) and a motor-driven valve (MF-011 or 

MF-012) in each feed line (respectively) to monitor and control the feedwater flow. 

Figure 3.26. Image of RNS, CVS, and MF Pumps 

When the steam exits the SG, there are two possible paths for it to follow: 

(1) One path is controlled by a normally open air-operated ball valve (MS-001 for SG1 and MS-002 

for SG2) and monitored by a vortex flow meter (FVM-001 for SG1 and FVM-002 for SG2). 

(2) The other path is controlled by a normally closed air-operated ball valve (MS-009 for SG1 

and MS-010 for SG2) and monitored by a vortex flow meter (FVM-009 for SG1 and FVM-010 

for SG2). 

Both of these lines will eventually exit the building through the steam exhaust. The difference 

between the two lines is that the first goes through the turbine simulation path with a motor-operated 

logic ball valve (MS-008) and monitored by a vortex flow meter (FVM-003), while the other is used 

for startup of the facility. 
 3-50 

Table 3.20. FST, MFP, and SG Steam Vent Instrument Names and Locations

LDP, FMM, FVM, and PT Cells Thermocouples 

Tag Name Description Tag Name Description 

FMM-001 SG1 Feed Flow TF-005 Lab Ambient Temperature @ Ground Level 

FMM-002 CL2 Loop Flow TF-006 Lab Ambient Temperature @ 2nd Level 

FVM-001 SG1 Main Steam Flow TF-007 Lab Ambient Temperature @ 3rd Level 

FVM-002 SG2 Main Steam Flow TF-009 SG1 PORV Line Temperature 

FVM-003 Main Steam Combined Flow TF-010 SG2 PORV Line Temperature 

FVM-009 SG1 PORV Steam Flow TF-301 TF-301/SG1 (via Signal Conditioner) 

FVM-010 SG2 PORV Steam Flow TF-310 TF-310/SG2 (via Signal Conditioner) 

LDP-001 FST Level TF-311 SG1 Feed Header Temperature 

PT-001 MFP Discharge Pressure TF-312 SG2 Feed Header Temperature 

PT-002 MS Header Pressure 

PT-003 Lab Barometric Pressure 

PT-009 SG1 PORV Line Pressure 

PT-010 SG2 PORV Line Pressure 

PT-301 SG1 Pressure 

PT-302 SG2 Pressure 

3.9.4 Water Purification System 

The water for the feedwater system comes from potable water. The DI purifies the incoming 

potable water for use in the secondary loop. The water first passes through two parallel resin traps 

(resin traps 1 and 2). It is then demineralized by four demineralization tanks, which are arranged 

into two parallel groups (demineralization banks 1 and 2). The water then passes through 

two additional resin traps (resin traps 3 and 4). The flow rate is then measured by a vortex flow 

meter (FVM-004), and an injection line then allows chemicals to be injected from a chemical 

addition tank. The chemicals that are added to the water include ammonium hydroxide 

(NH4OH) and NALCO 1254. Ammonium hydroxide is used to reduce the pH in the system 

and to minimize corrosion of the components. NALCO 1254 is an additive that prevents buildup 

of sludge within the system. After all treatments, the feedwater is sent to the FST. 
 4-1 

4. TEST MATRIX AND OVERVIEW 

Eight NRC-sponsored beyond-DBA tests (listed in Table 4.1) were run at the APEX-AP1000 

facility to confirm safety margins to support AP1000 design certification. Based on experience 

from previous AP600 design certification activities, these tests focused on DEDVI line breaks 

with various failures of the ADS. 

Of the 8 beyond-DBA tests, three (NRC-AP1000-01, 03, and 05) considered DEDVI line breaks 

with failure of the ADS1, 2, and 3 valves, and ADS4 valves on the PZR and non-PZR sides 

of the plant. Two CL break tests (NRC-AP1000-04 and 06) were run to examine the sensitivity 

of ADS4 valve failure with respect to a smaller size breaks than a DEDVI line break. A 

Probabilstic Risk Assessment verification test (NRC-AP1000-10) and station blackout test 

(NRC-AP1000-11) were run to complete the confirmatory test matrix. Finally, a Mode 5 

operation test (NRC-AP1000-02) was run with the ACCs and CMTs unavailable. Appendix C 

presents detailed time transient plots, sequence of events, and facility configurations for each 

test. These beyond-DBA tests were selected, based on previous experience with the AP600 

design certification, as the scenarios which would potentially be the most challangeing to the 

AP1000 passive safety systems. 

In addition, three upper plenum steady-state entrainment tests (NRC-AP1000-07, 08, and 09) 

were run to provide data for thermal-hydraulic code assessment, but those tests are not 

discussed in this report. Please note that all transient plots presented in this section are plotted 

against dimensionless time. 
 4-2 

Table 4.1. APEX-AP1000 Confirmatory Test Matrix 

Test Number Description Date 

Completed

NRC-AP1000-01 DEDVI Line Break with Failure of ADS1, 2, and 3. 

Both ADS4 trains fully functional. 

6/5/03 

NRC-AP1000-02 Mode 5 Operation with Loss of RNS Cooling. 

No break. ADS1, 2, and 3 open. SG secondary side assumed drained. 

CMTs and ACCs unavailable or failed (one valve on PZR side available). 

9/30/04 

NRC-AP1000-03 DEDVI Line Break with Failure of 2 of 4 ADS4 Valves on PZR side. 

ADS1, 2, and 3 valves fully functional. 

6/11/03 

NRC-AP1000-04 CL #4, 1" Break with Failure of 1 of 4 ADS4 Valves on non-PZR side. 

ADS1, 2, and 3 valves fully functional. Containment sump degraded 

(CMT#1, ACC #1, sump recirculation line #1 and IRST line #1 unavailable). 

10/9/03 

NRC-AP1000-05 DEDVI Line Break with Failure of 2 of 4 ADS4 Valves on non-PZR side. 

ADS1, 2, and 3 valves fully functional. 

6/10/03 

NRC-AP1000-06 CL #4, 2” Break with Failure of 2 of 4 ADS4 Valves on non-PZR side 

and PRHR. 

ADS1, 2, and 3 valves fully functional. 

8/15/03 

NRC-AP1000-07* Upper Plenum Entrainment Test with Upper Internals Installed. 

Core inventory allowed to boil-off from top of hot legs to top of core. 

Three different power levels were run. 

8/24/04 

NRC-AP1000-08* Upper Plenum Entrainment Test with Modified Power Stepping. 

Repeat of NRC-AP1000-07 with additional power steps. 

9/7/04 

NRC-AP1000-09* Upper Plenum Entrainment Test with Upper Internals Removed. 

Core inventory allowed to boil-off from top of hot legs to top of core. 

Three different power levels were run. 

9/16/04 

NRC-AP1000-10 PRA Verification Test. 

Bottom of CL #4, 2” break with failure of 1 of 4 ADS4 valves 

on non-PZR side and failure of ADS1, 2, and 3 valves. 

4/29/04 

NRC-AP1000-11 Station Blackout. 

No break. Failure of 1 of 4 ADS4 valves on non-PZR side. 

7/29/04 

*Not discussed in this report 
 4-3 

4.1 Typical Sequence of Events for an APEX-AP1000 SBLOCA

This section presents the typical sequence of events for a small break (SB) LOCA scenario 

in the APEX-AP1000 test facility [Ref. 3]. The AP1000 safety logic (Appendix B) was closely 

modeled in the APEX test facility. Each SBLOCA test was initiated from steady-state operating 

conditions and started by opening a valve located on the primary loop piping. For SBLOCAs, 

the bottom of CL4 was typically chosen as the break location. Fluid flow from the break 

was directed to the break separator, where the vapor and liquid flows were measured separately. 

The vapor flow was then directed to the BAMS header, and the liquid flow to the primary sump. 

The SBLOCA scenario can be divided into five different phases, as shown in Figure 4.1 [Ref. 3]: 

(1) High-Pressure Blowdown 

(2) Natural Circulation 

(3) ADS Blowdown 

(4) IRWST Injection 

(5) Sump Recirculation 

Figure 4.1 Pressure History for a Typical 2-inch Cold Leg Break in APEX [Ref. 3] 

4.1.1 High-Pressure Blowdown Phase 

The break in the primary loop caused a rapid initial blowdown of the system. Depending on the 

break size, this period was relatively short. As the PZR level decreased, it caused the issuance 

of an “S signal,” which generated the following actions: 

• PRHR isolation valves opened 

• CMT isolation valves opened 

• feedwater pump tripped 

• steam generator steam flow isolated 

• reactor coolant pumps tripped 

• reactor power in decay mode to simulate a reactor scram 

As the system blowdown continued, saturated conditions were reached in the HLs. The onset 

of saturated conditions in the HLs concluded the high-pressure blowdown phase. 

High-Pressure 

Blowdown 

Phase 

Natural 

Circulation 

Phase 

ADS Blowdown 

Phase 

IRWST 

Injection 

Phase 

Sump 

Recirculation 

Phase 

Time

Pressure 
 4-4 

4.1.2 Natural Circulation Phase 

When the RCPs were tripped, the primary loop flow transitioned from forced to natural 

circulation. This natural circulation flow continued until voiding occurred in the primary side of 

the SG tubes. Operation of the PRHR and CMTs created additional paths for buoyancy-driven 

core heat removal. Water from the HL/ADS4 line connection rose to the PRHR inlet header and 

traveled through the PRHR HX. The cold water leaving the PRHR HX flowed to the CLs. In a 

similar manner, each CMT provided cooling to the core by recirculating its volume. Relatively 

hot water from the CLs rose into the CMT-PBL to the CMT’s upper head. The hot water then 

displaced the cold water leaving the CMTs into the DVI injection line and finally into the RPV. 

As the facility depressurized, steam collected or formed as a result of flashing in the CMT’s 

upper head. This caused the CMT to drain and ceased the recirculation mode. In some cases, 

the CMTs partially refilled as a result of steam condensation. Eventually, the liquid level in the CMT 

reached the ADS1 actuation setpoint. The natural circulation phase of the transient concluded 

upon ADS1 actuation. 

4.1.3 ADS Blowdown Phase 

Both trains of the AP1000 ADS1–3 lines located on the top of the PZR were modeled in APEX. 

The flow nozzles used in the ADS lines were scaled to simulate choked flow, assuming pressure 

similitude subsequent to ADS3 actuation. The control logic opened ADS1, ADS2, and ADS3 in 

sequence. The ACCs typically began to inject sometime during ADS1–3 blowdown. The large 

flows caused by ACC injection typically delayed CMT draining. After ACC injection flow rates 

became low enough, the CMTs continued to drain. Eventually, the CMT liquid level dropped to 

the ADS4 actuation setpoint. APEX simulated both trains of the AP1000 ADS4 lines. After ADS4 

actuation, the primary system depressurized to atmospheric conditions. This permitted the IRWST 

check valves to open in response to the hydrostatic head of the IRWST liquid. The onset of 

IRWST injection concluded the ADS blowdown phase. 

4.1.4 IRWST Injection Phase 

Upon IRWST injection, the liquid in the reactor core became subcooled. This condition persisted 

until the level in the IRWST was relatively low. At that time, the reduced IRWST injection flow rate 

was no longer sufficient to keep the core fluid subcooled. When saturated conditions were reached 

at the core exit, oscillations were observed in the IRWST injection flow, system pressure, RPV level, 

and ADS4 flow. These oscillations are an artifact of the APEX primary to secondary sump 

connect, were labeled “Return to Saturation Oscillations” (RSOs) in the APEX-AP600 

confirmatory testing program, and became the subject of a separate parametric test. In the 

APEX-AP1000 and APEX-AP600 tests, these oscillations did not impact core cooling because 

the RPV liquid level always remained near the top of the HL. When the liquid level in the IRWST 

reached the sump recirculation setpoint, the sump recirculation valves opened. The oscillations 

were terminated at this time, and the IRWST injection phase concluded. 
 4-5 

4.1.5 Sump Recirculation Phase 

The sump recirculation phase marked the beginning of long-term cooling. Liquid from 

the primary sump flowed through the sump recirculation lines into the DVI lines and then into the 

RPV. The boiling in the core provided the density differences required to draw the water from the 

sump to the RPV. Steam generated in the core was vented through the ADS4 lines. Sump 

recirculation usually continued for 30 minutes until terminated by the operator, thus concluding 

the test. In some cases, the sump recirculation phase continued longer when a degraded sump 

was simulated. 

4.2 Test NRC-AP1000-01

Test NRC-AP1000-01 simulated a DEDVI line break with a complete failure of ADS1, 2, and 3. 

This test was designed as a beyond-DBA scenario because of the multiple valve failures. 

All four ADS4 valves were available, and this test investigated the ability of the ADS4 system 

to accomplish system depressurization to allow for IRWST injection. The PRHR HX was inactive 

during this test. 

The APEX-AP1000 facility models both trains of the AP1000 ADS1–3 lines located at the top 

of the PZR after a nozzle that models the scaled choked flow area of the AP1000. The ADS1–3 

valve train typically discharges to the ADS1–3 separator, where the two-phase flow is separated 

and measured. 

Figure 4.2 illustrates the facility configuration during this test and contains non-dimensional plots 

of RPV-wide collapsed liquid level, passive safety system injection flow rates (sides 1 and 2), 

and core heater rod temperatures. The DEDVI Line 1 break opened at zero seconds 

and created a rapid depressurization, which caused the CMTs to begin injection immediately. 

However, half of this makeup flow exited the DVI Line 1 break. As the facility depressurized, steam 

collected or formed as a result of flashing in the CMT’s upper head. This caused the CMT to 

drain, terminating the natural recirculation phase of the transient. The ACCs typically begin 

injecting during ADS1–3 actuations, but since ADS1–3 were failed shut, ACC2 began injecting 

around time 1. When the liquid level in the CMTs was low enough, ADS4-1 opened and helped 

to depressurize the plant even further. After the ADS4 valves opened, CMT2 and ACC2 injection 

flow rates increased significantly. The reactor vessel collapsed liquid level remained 

approximately constant during ACC and CMT injection (between time 2.7 and 8) as the plant 

was depressurized using the ADS. 

At around time 7, the plant reached atmospheric conditions. This permitted the IRWST check 

valves to open in response to the hydrostatic head of the IRWST liquid, thus ending the ADS 

blowdown phase of the transient. It should be noted that there was a noticeable overlap between 

when CMT2 injection flow ended and IRWST injection began, which provided adequate core 

cooling during this time period. As IRWST injection flow increased, the vessel collapsed liquid 

level began to rise around time 10. Later in the test (around time 14), CMT1 experienced short 

periods of refill as a result of condensation-induced pressure drops, which is an artifact of the 

one-quarter height scaling. For the remainder of the test, the HL water level remained high, and 

the core remained covered as a result of adequate injection flow from the IRWST. The test was 

ended before sump recirculation or long-term cooling was initiated. No heater rod hightemperature excursion was experienced, and the core remained adequately cooled during the 

entire test. 
 4-6 

0 10 20 30 40 50

Top of Core

Bottom of Core

0 5 10 15 20 25

0 5 10 15 20 25 0 5 10 15 20 25

Figure 4.2 Test NRC-AP1000-01 Passive Safety System Performance. 

RPV Liquid Level 

Dimensionless Time 

Collapsed Liquid Level 

Core Temp. and System Pressure

Dimensionless Time 

Temperature and Pressure 

Side 2 Injection Flows 

Dimensionless Time 

Volumetric Flow Rate 

Side 1 Injection Flows 

Dimensionless Time 

Volumetric Flow Rate 

CMT 2 

ACC 2 

CMT 1 

ACC 1 

IRWST 

ADS4-1 

Opened 

DVI 1 

SG 2 SG 1 

DVI 2 

IRWST 

CMT 2 CMT 1 

ACC 2 ACC 1 

ADS 1, 2, 3 

PZR 

RX 

CL 2,4 HL 2 CL 1,3 HL 1 

Sparger 

ADS4-2 ADS4-1

Failed Closed: 

Break To CL 1 To CL 3 

Side 2 Side 1 

DEDVI Line Break

ADS4-2 

Opened 

TEMP 

PRESS 
 4-7 

4.3 Test NRC-AP1000-02

Test NRC-AP1000-02 investigated passive safety system performance during Mode 5 

(cold shutdown) operation with a loss of the RNS. The objectives of this test were to determine 

system behavior during a loss of shutdown cooling. Specifically, if the RCS would pressurize 

sufficiently to prevent IRWST injection, and if counter-current flow limitation (CCFL) prevent the 

PZR from draining. There was no break in this test; ADS3 was opened at zero seconds 

(simulated two trains of ADS1, 2, and 3 in AP1000), the SG secondary side was assumed 

drained, the CMTs and ACCs were unavailable, and one ADS4 valve was failed closed. Power 

was increased to about 120 kW at the beginning of the test. 

Figure 4.3 illustrates the facility configuration during this test, and contains non-dimensional plots 

of RPV-wide collapsed liquid level, IRWST injection flow rates, core heater rod temperatures, 

and PZR wide-range collapsed liquid level. The test was begun at zero seconds when the 

reactor power was at 120 kW to simulate a loss of RNS cooling. As water boiled off from the 

core and exited through the open ADS3 line, the RPV collapsed liquid level dropped to below 

the top of the core at around time 5. The ADS4-1 valve (non-PZR HL) failed to open on demand 

at around time 9.7 and decreased the RPV pressure to allow for IRWST injection. By time 10.2, the 

RPV collapsed liquid level rose above the top of the core and remained high until around time 20. 

The PZR completely drained just after IRWST injection began at around time 10. At about time 

30, the sump recirculation valves were opened on low IRWST level, thus ending the IRWST 

injection phase, and the long-term cooling began. The core remained adequately cooled during 

the entire test. 

4.4 Test NRC-AP1000-03

This test was a DEDVI line break with failure of two of the four ADS4 valves on the PZR side 

of the plant. The purpose of this test was to investigate the sensitivity of ADS4 valve failure 

on core coolability. Figure 4.4 shows the plant configuration for this test and plots of safety 

injection flow rates (sides 1 and 2), along with reactor vessel collapsed liquid level, 

and core temperature. 

The break in DVI1 was opened at time zero, and the CMTs began injecting immediately as the 

primary system rapidly depressurized. As the CMTs emptied, the ADS1–3 valves began to 

actuate, beginning the ADS blowdown phase of the transient. Shortly after, ACC2 began injecting. 

The ADS4-1 valves opened just after time 1.7 to reduce the plant pressure to allow for IRWST 

injection, thus ending the ADS blowdown phase. IRWST injection began around time 14, while 

CMT2 injection ended around time 10. Between time 10 and 14, no makeup water was provided 

to the core. 

As water was injected into the core from the IRWST, the core level increased. Eventually, 

the IRWST finished draining, and the plant successfully entered sump recirculation for long-term 

cooling around time 75. During the test, core temperature spiked a few times for very brief periods, 

but enough water was provided from the passive safety systems to keep the core adequately 

cooled throughout the entire test. 
 4-8 

 

0 10 20 30 40 50

Top of Core

Bottom of Core

0 10 20 30 40 50

0 10 20 30 40 50 0 10 20 30 40 50

Figure 4.3 Test NRC-AP1000-02 Passive Safety System Performance. 

RPV Liquid Level 

Dimensionless Time 

Collapsed Liquid Level 

Core Temperature 

Dimensionless Time 

Temperature 

IRWST Injection Flows 

Dimensionless Time 

Volumetric Flow Rate 

PZR Liquid Level 

Dimensionless Time 

Collapsed Liquid Level 

ADS4-1 

Opened 

DVI 1 

SG 2 SG 1 

DVI 2 

IRWST 

CMT 2 CMT 1 

ACC 2 ACC 1 

ADS 1, 2, 3 

PZR 

RX 

CL 2,4 HL 2 CL 1,3 HL 1 

Sparger 

ADS4-2 ADS4-1

Always Open: 

Unavailable 

To CL 1 To CL 3 

Side 2 Side 1 

Mode 5 Loss of RNS 

Cooling

Unavailable 

Failed 

Open 
 4-9 

0 20 40 60 80 100

Top of Core

Bottom of Core

0 20 40 60 80 100

0 10 20 30 40 0 2 4 6 8 10

Figure 4.4 Test NRC-AP1000-03 Passive Safety System Performance. 

DVI 1 

SG 2 SG 1 

DVI 2 

IRWST 

CMT 2 CMT 1 

ACC 2 ACC 1 

ADS 1, 2, 3 

PZR 

RX 

CL 2,4 HL 2 CL 1,3 HL 1 

Sparger 

ADS4-2 ADS4-1 Failed 

Closed Break To CL 1 To CL 3 

Side 2 Side 1 

RX Vessel Liquid Level 

Dimensionless Time 

Collapsed Liquid Level 

Core Temperature 

Dimensionless Time 

Temperature 

Side 2 Injection Flows 

Dimensionless Time 

Volumetric Flow Rate 

Side 1 Injection Flows 

Dimensionless Time 

Volumetric Flow Rate 

CMT 2 

ACC 2 

IRWST 

CMT 1 

ACC 1 

ADS4-1 

Opened 

ADS 1, 2, 3 

Opened 

DEDVI Line Break
 4-10 

4.5 Test NRC-AP1000-04

Test NRC-AP1000-04 was a beyond-DBA test simulating a 1-inch CL break with one ADS4 valve 

failed shut (50% line flow) on the non-PZR HL. In addition, CMT1, ACC1, sump recirculation 

line 1, and IRWST injection line 1 were unavailable. The intent of this test was to investigate 

the sensitivity of ADS4 valve failure on core coolability with loss of several safety injection paths, 

and to determine the impact of a simulated degraded sump on long-term cooling. 

Figure 4.5 illustrates the facility configuration for this test, and contains plots of RPV-wide 

collapsed liquid level, passive safety injection flow rates (side 2), and core heater rod 

temperature and power. The break was initiated at zero seconds and CMT2 began injecting 

immediately as the natural circulation phase of the transient was entered. During the initial ADS 

blowdown phase, the reactor vessel collapsed water level dropped to the middle of the core. 

ACC2 began injecting around time 4 during ADS1–3 blowdown. The RPV collapsed liquid level 

continued to drop until IRWST injection began around time 7, thus ending the ADS blowdown 

phase of the transient. IRWST injection began before CMT2 was completely emptied, ensuring 

that makeup water was provided to the core at all times. At around time 20, the RPV liquid level 

began to decrease as IRWST injection flow rate decreased. 

At about time 60, the primary sump valve was opened, which transitioned the plant from IRWST 

injection to long-term cooling. Primary sump overflow began, and water was recirculated from 

the sump to the IRWST, through the DVI2 line, and into the reactor vessel as the long-term 

cooling phase was entered. During sump recirculation, a drain at the bottom of the sump was 

opened to simulate a degraded sump. The presence of a degraded sump did not affect core 

coolability, and the core remained adequately cooled throughout the test. The core power was 

shown in Figure 4.5 to illustrate the standard decay power curve used in many APEX LOCA 

tests. 
 4-11 

 

0 20 40 60 80 100

Top of Core

Bottom of Core

0 20 40 60 80 100

0 20 40 60 80 100

Figure 4.5 Test NRC-AP1000-04 Passive Safety System Performance. 

RX Vessel Liquid Level 

Dimensionless Time 

Collapsed Liquid Level 

Core Temperature and Power 

Dimensionless Time 

Temperature and Power 

Side 2 Injection Flows 

Volumetric Flow Rate 

Dimensionless Time 

CMT 2 

ACC 2 

TEMP 

IRWST 

POWER 

DVI 1 

SG 2 SG1 

DVI 2 

IRWST 

CMT 2 CMT 1 

ACC 2 ACC 1 

ADS 1, 2, 3 

PZR 

RX 

CL 2,4 HL 2 CL 1,3 HL 1 

Sparger 

ADS4-2 ADS4-1

Failed 

Closed Break 

To CL 1 To CL 3 

Side 2 Side 1 

1-inch CL 4 Break with 

Deegraded Sump
 4-12 

4.6 Test NRC-AP1000-05

This test was a DEDVI line break with failure of two of the four ADS4 valves on the non-PZR 

side of the plant. This test was the counterpart to test NRC-AP1000-03, where two of the four 

ADS4 valves failed to open on the PZR side of the plant. 

Figure 4.6 shows the plant configuration for this test and plots of injection flow rates (sides 1 and 2), 

along with reactor vessel level and core temperature. At time zero, the DVI Line 1 break 

was opened and the plant began to depressurize; the CMTs began injecting immediately. 

The plant went through the same natural circulation and ADS blowdown phases as in test 

NRC-AP1000-03. At around time 30 in test NRC-AP1000-05, the vessel inventory began to decrease 

because no injection flow was provided to the vessel. Water exiting the break and ADS4-2 line 

caused the vessel inventory to continue to decrease, even as the two-phase mixture level 

dropped below the bottom of the HL at around time 32. 

Upper plenum entrainment into the HL and out the ADS4-2 line contined to deplete the rector 

vessel inventory from time 30 to 40, which lead to core uncovery before the IRWST injection 

could begin. At time 40, there was an increase in core tempertature and the heater rods were 

turned off to avoid damage. Consequently, failing closed the two ADS4 valves on the PZR side 

of the plant resulted in core heatup. The PZR was unable to drain during this test because of 

the presence of liquid entrainment out of ADS4-2 in HL2 or because of the presence of CCFL in 

the surge line from flow out the ADS1-3 valves. Section 5.2 of this report provides a comparison 

between tests NRC-AP1000-03 and 05, which explains why the core was uncovered in test NRCAP1000-05 and not in test NRC-AP1000-03. 

4.7 Test NRC-AP1000-06

This test was a 2-inch break off the bottom of CL4 with a failure of two of the four ADS4 valves 

on the non-PZR side of the plant. It was apparent from NRC-AP1000-05 that this type of ADS4 

failure can lead to core uncovery, so the purpose of this test was to determine whether moving 

the break from DVI1 to CL4 resulted in similar core uncovery. Figure 4.7 shows the plant 

configuration for this test and plots passive safety system injection flow rate (side 2), along with 

reactor vessel level, core temperature, primary system pressure, and PZR collapsed liquid level. 

As in previous tests, the break was opened at time zero. CMT1 and 2 began to inject immediately, 

but since the break was relatively small, the ADS setpoint (based on CMT low level) was not 

reached until around time 15 when the ADS1 and ADS2 valves actuated. After ADS1, 2, and 3 

actuation, the total injection flow increased significantly until falling to zero by time 37, thus 

ending the ADS blowdown phase of the trainsient. The ADS blowdown phase in this test was 

significantly longer than the ADS blowdown phase in DVI line breaks, as would be expected 

from a smaller CL break. 

IRWST injection had not started by the time the CMTs emptied. Without any additional safety 

injection, the reactor vessel water level continued to drop as a result of boiling in the core and 

upper plenum entrainment. Eventually, without IRWST injection, the core temperature rose as a 

result of boil-off of primary liquid, and a temperature excursion occurred at around time 50, similar to 

test NRC-AP1000-05. The heater rods were shut off to avoid damage, thus ending the test. 
 4-13 

0 10 20 30 40 50

Top of Core

Bottom of Core

0 10 20 30 40 50

0 10 20 30 40 50 0 10 20 30 40 50

Figure 4.6 Test NRC-AP1000-05 Passive Safety System Performance. 

RPVLiquid Level 

Dimensionless Time 

Collapsed Liquid Level 

Core Temp. and Pressure 

Dimensionless Time 

Temperature and Pressure 

Side 2 Injection Flows 

Dimensionless Time 

Volumetric Flow Rate 

PZR Liquid Level 

Dimensionless Time 

Collapsed Liquid Level 

CMT 2 

ACC 2 

IRWST 

DVI 1 

SG 2 SG 1 

DVI 2 

IRWST 

CMT 2 CMT 1 

ACC 2 ACC 1 

ADS 1, 2, 3 

PZR 

RX 

CL 2,4 HL 2 CL 1,3 HL 1 

Sparger 

ADS4-2 ADS4-1 Failed 

Break Closed To CL1 To CL 3 

Side 2 Side 1 

ADS4-2 

Opened 

DEDVI Line Break

PRESS 

TEMP 
 4-14 

0 10 20 30 40 50

Top of Core

Bottom of Core

0 10 20 30 40 50

0 10 20 30 40 50 0 10 20 30 40 50

Figure 4.7 Test NRC-AP1000-06 Passive Safety System Performance. 

DVI 1 

SG 2 SG 1 

DVI 2 

IRWST 

CMT 2 CMT 1 

ACC 2 ACC 1 

ADS 1, 2, 3 

PZR 

RX 

CL 2,4 HL 2 CL 1,3 HL 1 

Sparger 

ADS4-2 ADS4-1 Failed 

Closed 

Break 

To CL 1 To CL 3 

Side 2 Side 1 

RPV Liquid Level 

Dimensionless Time 

Collapsed Liquid Level 

Core Temp and Pressure 

Dimensionless Time 

Temperature and Pressure 

Side 2 Injection Flows 

Dimensionless Time 

Volumetric Flow Rate 

PZR Liquid Level 

Dimensionless Time 

Collapsed Liquid Level 

CMT 2 

ACC 2 

TEMP 

PRESS 

2-inch CL 4 Break
 4-15 

4.8 Test NRC-AP1000-10

The objective of this test was to validate PRA assumptions with regard to multiple failures of the 

ADS. This test simulated a 2-inch cold leg break off the bottom of CL4. The test was initiated 

from normal operating temperature and pressure. For this test, ADS1–3 and one of the four 

ADS4 valves failed to open on the non-PZR side of the plant. The steam generator PORVs 

were allowed to cycle to help remove decay heat during the initial blowdown stage of the 

transient. The test was terminated after 30 minutes of steady-state long-term cooling. 

Figure 4.8 shows the facility configuration for this test, along with figures showing the passive 

safety injection flow rate; RPV collapsed liquid level, and ADS4-1 and 2 flow qualities. For this 

test, the side 1 and side 2 injection timing and flow rates were very similar; hence, only side 2 is 

shown in Figure 4.8. The break was initiated at zero seconds, and the CMTs immediately 

began to drain. The ACCs began to inject shortly after ADS4 valve actuation at around time 3. 

This caused the plant to further depressurize and increased ACC injection flow rates considerably. 

Between time 6 and 9, there was a noticeable oscillation of IRWST injection flow rate, core 

liquid level, and ADS4-2 line flow qualitiy. By time 9, the RPV pressure was very close to 

atmospheric and the HLs were almost completely flooded (not shown). This oscillatory behavior 

is consistent with separate effects experiments run in the Air-Water Test Loop for Advanced 

Thermal-hydraulic Studies (ATLATS) [Ref. 4], which investigated the intermittent nature of liquid 

entrainment from the HLs into a vertical ADS4 line. At around time 10, only liquid exited the 

ADS4 valves as water from the IRWST was injected at a high rate to the core. 

The RPV liquid inventory reached a maximum around time 15, and then began to slowly decrease as 

the IRWST injection flow rate decreased. Oscillations were experienced again from time 26 to 

35. These RSOs are consistent with previous AP600 tests, and are described in Section 4.1.4 

of this report. The plant entered long-cooling at time 43, when the sump recirculation lines were 

opened, thus ending the IRWST injection phase of the transient. No high-temperature 

excursion was experienced and the core remained covered and adequately cooled throughout 

the test. 

4.9 Test NRC-AP1000-11

This test simulated a station blackout with a failure of one of the four ADS4 valves on the non-PZR 

side of the plant. Previous tests (NRC-AP1000-05 and 06) showed that failure of ADS4 valves 

on the non-PZR side of the plant are the most challenging to the AP1000 passive safety systems. 

The APEX-AP1000 control logic (Appendix B) was modified to simulate a station blackout 

(loss of all AC power). The core operated at a reduced power (600 kW) and with a modified 

decay power curve to preserve the integrated core energy. The SG PORVs were set to cycle 

between two relatively high pressures. When the SG water level decreased to a specified level 

as a result of PORV cycling, an “S“ signal was sent to open the CMT injection valves and the PRHR 

HX outlet valve. The ADS was actuated at around time 68 based on the control logic. 

Figure 4.9 shows the plant configuration for this test and plots of passive safety system injection 

flow rates (side 2), along with reactor vessel collapsed liquid level, PZR pressure and temperature, 

and core temperature. For the first several hours of the station blackout test, the PORVs cycled 

on and off. The “S“ signal was generated around time 68, and the ADS1 valve actuated. The 
 4-16 

plant immediately began to depressurize, and the ADS control logic depressurized the plant to 

allow passive safety system injection. The injection flow decreased very close to zero around 

time 70, but only for a short while until IRWST injection began. No temperature excursion was 

experienced, and the core remained covered and adequately cooled throughout the test, even 

with one of the four ADS4 valves failing to open on the non-PZR side. 
 4-17 

 

0 10 20 30 40 50

Top of Core

Bottom of Core

0 10 20 30 40 50

0 10 20 30 40 50

 

Figure 4.8 Test NRC-AP1000-10 Passive Safety System Performance. 

RX Vessel Liquid Level 

Dimensionless Time 

Collapsed Liquid Level 

ADS4-2 Flow Quality 

Dimensionless Time 

Flow Quality 

DVI 1 

SG 2 SG 1 

DVI 2 

IRWST 

CMT 2 CMT 1 

ACC 2 ACC 1 

ADS 1, 2, 3 

PZR 

RX 

CL 2,4 HL 2 CL 1,3 HL 1 

Sparger 

ADS4-2 ADS4-1

Failed 

Closed 

Break 

To CL1 To CL 3 

Side 2 Side 1 

2-inch CL 4 Break

Dimensionless Time 

IRWST 

ACC 2 

CMT 2 

Volumetric Flow Rate 

Side 2 Injection Flows 
 4-18 

0 20 40 60 80

Top of Core

Bottom of Core

0 20 40 60 80

65 67 69 71 73 75 0 20 40 60 80

Figure 4.9 Test NRC-AP1000-11 Passive Safety System Performance.

RPVLiquid Level 

Dimensionless Time 

Collapsed Liquid Level 

PZR Pressure and Level 

Dimensionless Time 

Pressure and Liquid Level 

PRESS 

LEVEL 

Side 2 Injection Flows 

Dimensionless Time 

Volumetric Flow Rate 

Core Temperature 

Dimensionless Time 

Temperature 

CMT2 

ACC 2 

IRWS 

CMT 2 

DVI1 

SG 2 SG 1 

DVI 2 

IRWST 

CMT 2 CMT 1 

ACC 2 ACC 1 

ADS 1, 2, 3 

PZR 

RX 

CL 2,4 HL 2 CL 1,3 HL 1 

Sparger 

ADS4-2 ADS4-1

Failed 

Open 

To CL1 To CL 3 

Side 2 Side 1 

Station Blackout
 5-1 

5. COMPARISONS AND DISCUSSION 

This section provides comparisons between design-basis test DBA-02 and beyond-DBA test 

NRC-AP1000-05, and between tests NRC-AP1000-03 and NRC-AP1000-05. The comparisons 

illustrate the fact that under select beyond-DBA scenarios, core uncovery was experienced 

in APEX-AP1000, while the similar design-basis test did not lead to core uncovery. In addition, 

comparison of beyond-DBA tests NRC-AP1000-03 and NRC-AP1000-05 helps to explain why 

the core uncovered when two of the four ADS4 valves were failed on the non-PZR side of the plant, 

compared to the same ADS4 failure on the PZR side of the plant. 

5.1 Test DBA-02 versus Test NRC-AP1000-05

It is interesting to compare design-basis test DBA-02 [Ref. 5] to beyond-DBA test NRC-AP1000-05. 

Both tests simulated a DEDVI line break with ADS4 valve failure on the non-PZR side of the plant. 

The only difference between the two tests was that in DBA-02, only one of the four ADS4 valves 

were failed closed, while in NRC-AP1000-05, two of the four ADS4 valves were failed closed. 

Figure 5.1 shows a comparison of reactor vessel collapsed level for DBA-02 and NRC-AP1000-05. 

As previously shown in Figure 4.6, the core uncovered in NRC-AP1000-05 as a result of 

the increased rate at which water was entrained out of the upper plenum, into the HL, and out 

ADS4-2. In DBA-02, IRWST injection was significantly delayed as shown in Figure 5.2. 

In NRC-AP1000-05, the reactor vessel liquid level dropped fast enough and IRWST injection 

never began, since the heater rods were shut off around time 32 to avoid damage to the facility. 

It is also important to note that there existed a significant amount of time in DBA-02 (between 

time 23 and 28) when the reactor vessel did not receive any injection flow as shown in Figure 

5.2. During this period, core cooling was maintained by boil-off of the existing vessel inventory. 

0 10 20 30 40

Figure 5.1 Reactor Vessel Liquid Level (DBA-02 and NRC-AP1000-05). 

Dimensionless Time

Collapsed Liquid Level 

DBA-02 

NRC-05

Top of Core

Bottom of Core
 5-2 

20 25 30 35

Figure 5.2 IRWST Injection Flow Rates (DBA-02 and NRC-AP1000-05). 

Figure 5.3 compares the PZR water level for both tests. At around time 10, the PZR was able to 

drain as the plant depressurized in DBA-02, providing additional makeup water to the core, but 

in NRC-AP1000-05, the PZR wasn’t able to drain until the heater rods were turned off at around 

time 30. The PZR was unable to drain in NRC-AP1000-05 because of liquid entrainment in HL2 

flowing out the ADS4-2 line. This phenomenon is discussed in more detail in Section 5.2. 

Examination of the ADS4-2 flow quality (Figure 5.4) for both tests shows significant liquid carryover 

from the reactor vessel and entrainment from the HL out the ADS4-2 line. Even after the twophase mixture dropped below the bottom of the HL in NRC-AP1000-05 (time 25 in Figure 5.1) 

the ADS4-2 quality was still relatively low (~0.6). The primary mechanism through which water 

left the reactor vessel at low vessel liquid levels (liquid level below the bottom of the HLs) was upper 

plenum entrainment. Droplets from the two-phase reactor vessel mixture level were entrained 

by the steam, passed through the upper internals (with some droplets de-entraining), and 

entered the hot leg. From there, some droplets struck the HL pipe walls and created an annular 

flow of liquid out the ADS4 line or were de-entrained into the HL liquid level and entrained out 

the vertical ADS4 line by a mechanism of liquid entrainment in horizontal pipes with vertical-up 

branches [Ref. 4]. In DBA-02, there was still one operational ADS4 valve on the non-PZR side 

of the plant, providing another outlet for the steam. In NRC-AP1000-05, all of the steam exited 

the ADS4-2 line and the DEDVI break. Hence, a higher flow quality was observed in NRCAP1000-05 than in DBA-02. 

Dimensionless Time

Volumetric Flow Rate 

DBA-02 

NRC-05 

No Injection 

Flow
 5-3 

0 10 20 30 40

Figure 5.3 PZR Liquid Levels (DBA-02 and NRC-AP1000-05). 

0

0.2

0.4

0.6

0.8

1

5 15 25 35

Figure 5.4 ADS4-2 Flow Qualities (DBA-02 and NRC-AP1000-05). 

NRC-05

DBA-02

DBA-02

NRC-05

Dimensionless Time

Collapsed Liquid Level 

Dimensionless Time

Flow Quality 
 5-4 

5.2 Test NRC-AP1000-03 versus NRC-AP1000-05

As previously shown, core uncovery was experienced in NRC-AP1000-05, but not in 

NRC-AP1000-03. The only difference between these two DVI line break tests was the location 

of the ADS4 valve failures. In NRC-AP1000-03 (no core uncovery), two of the four ADS4 valves 

were failed closed on the PZR side of the plant, while in NRC-AP1000-05 (core uncovery), 

two of the four ADS4 valves were failed closed on the non-PZR side of the plant. As the plant 

went through the ADS blowdown phase of the transient, water was entrained through the ADS4-2 

line in NRC-AP1000-05 and the ADS4-1 line in NRC-AP1000-03. Both tests had the exact same 

timing (i.e., control logic), except that the heater rods were turned off in NRC-AP1000-05 

at around time 25 to avoid damage to the facility. 

The pressure transient was almost exactly the same for both tests, as shown in Figure 5.5. 

There was a slight difference between tests NRC-AP1000-03 and NRC-AP1000-05 in system 

pressure at around time 10. For both tests, there was no injection flow to the core between time 

17 and 25, as shown in Figure 5.6. During this time of no injection flow, the quality out of the 

open ADS4 valves for both tests was relatively low, ~0.6 for test NRC-AP1000-05 and ~0.4 for 

test NRC-AP1000-03, as shown in Figure 5.7. 

As shown in Figure 5.8, the RPV collapsed liquid level began to decrease in both tests at around 

the same time (time 18). However, the RPV liquid level decreased faster in test NRC-AP1000-

05 than in test NRC-AP1000-03. At around time 21, the two-phase mixture level in the vessel for 

both tests dropped below the bottom of the HL, as shown in Figure 5.9. Figure 5.10 shows an 

oscillation in the HL2 liquid level in both tests at about the same time the RPV two-phase 

mixture level dropped below the bottom of the HL. Examination of the HL1 liquid level (Figure 5.11) 

does not reveal any noticeable oscillations between time 21 and 25. In test NRC-AP1000-05, 

the liquid in HL2 was entrained out the ADS4-2 line, while no liquid was entrained out the ADS4-

2 line in test NRC-AP1000-03, since it was failed closed in that test. Hence, the source of the 

initial oscillations in HL2 was not attributable to the presence of HL entrainment. Examination of 

the PZR surge line liquid level (Figure 5.12) reveals that the PZR began to slowly drain in 

surges during the HL2 oscillations. Therefore, the oscillations in hot leg level were attributable 

to the PZR draining. In addition, in test NRC-AP1000-03, the oscillations in HL2 slowly 

diminished over time, while they were maintained in test NRC-AP1000-05 by the presence of 

liquid entrainment out of ADS4-2. Figure 5.13 shows the PZR liquid level for both tests. 

In summary, compared to test NRC-AP1000-03, the RPV liquid level dropped faster in 

test NRC-AP1000-05 because of the presence of the PZR, which enhanced liquid entrained 

out the open ADS4 line in test NRC-AP1000-05. As shown in Figure 5.14, IRWST injection 

was significantly delayed in NRC-AP1000-03, but never began in NRC-AP1000-05. 

(Test NRC-AP1000-05 was terminated at time 25, which caused IRWST injection to begin.) Lack of 

injection water from the IRWST caused the core to uncover in test NRC-AP1000-05. 

In similar full-scale AP600 tests run in ROSA [Ref. 3], core uncovery was not experienced. 

This is because the APEX facility has a reduced IRWST head, since it was scaled to one-quarter 

height. Hence, the delay in IRWST injection is a known conservatism in the APEX test facility. 
 5-5 

0 5 10 15 20 25 30

Dimensionless Time

Pressure

NRC-05

NRC-03

Figure 5.5 System Pressures (NRC-AP1000-03 and 05) 

0 5 10 15 20 25 30

Dimensionless Time

Volumetric Flow Rate

NRC-05

NRC-03

Figure 5.6 Safety Injection Flow Rates (NRC-AP1000-03 and 05) 
 5-6 

0

0.2

0.4

0.6

0.8

1

0 5 10 15 20 25 30

Dimensionless Time

Flow Quality

NRC-03 (ADS 4-1)

NRC-05 (ADS 4-2)

Figure 5.7 ADS4 Flow Qualities (NRC-AP1000-03 and 05) 

0 5 10 15 20 25 30

Dimensionless Time

Collapsed Liquid Level

NRC-05

NRC-03

Top of Core

Bottom of Core

Figure 5.8 RPV Collapsed Liquid Levels (NRC-AP1000-03 and 05) 
 5-7 

15 17 19 21 23 25

Dimensionless Time

Two-Phase Mixture Level

NRC-05

NRC-03

Bottom of Hot Leg

Figure 5.9 RPV Two-Phase Mixture Levels (NRC-AP1000-03 and 05) 

15 17 19 21 23 25

Dimensionless Time

Collapsed Liquid Level

NRC-03 (HL 2)

ADS 4-2 Entrainment NRC-05 (HL 2)

(NRC-05)

Figure 5.10 HL2 Collapsed Liquid Levels (NRC-AP1000-03 and 05) 
 5-8 

15 17 19 21 23 25

Dimensionless Time

Collapsed Liquid Level

NRC-03 (HL 1)

NRC-05 (HL 1)

ADS 4-1 Entrainment

(NRC-03)

Figure 5.11 HL1 Collapsed Liquid Levels (NRC-AP1000-03 and 05) 

15 17 19 21 23 25

Dimensionless TIme

Volumetric Flow Rate

NRC-05

NRC-03

Figure 5.12 PZR Surge Line Liquid Levels (NRC-AP1000-03 and 05) 
 5-9 

0 5 10 15 20 25 30

Dimensionless Time

Collapsed Liquid Level

NRC-05

NRC-03

Figure 5.13 PZR Collapsed Liquid Levels (NRC-AP1000-03 and 05) 

0 5 10 15 20 25 30

Dimensionless Time

Volumetric Flow Rate

NRC-05

NRC-03

Figure 5.14 IRWST Injection Flow Rates (NRC-AP1000-03 and 05)
 6-1 

6. SUMMARY AND CONCLUSIONS 

The APEX facility is a unique thermal-hydraulic integral system test facility used to assess 

the performance of passive safety systems in the Westinghouse AP600 and AP1000 designs. 

To address performance specific to the AP1000 design, the APEX facility underwent significant 

modifications in 2002 in order to more accurately represent the AP1000 reactor. Those facility 

modifications included an increase in the maximum core power; new PZR and surge lines; 

larger CMTs; larger-diameter ADS4 system piping; and decreased line resistances for the CMTs, 

ADS4, and PRHR HX. A comprehensive scaling analysis, performed by Reyes and presented 

in Appendix A, guided the facility modifications to ensure that AP1000 conditions were preserved in 

the APEX experiments. 

The accident scenarios investigated in the APEX-AP1000 facility included 1- and 2-inch CL breaks, 

DEDVI line breaks, Mode 5 operation with loss of RNS cooling, and station blackout. Based on 

experience from analyzing previous AP600 tests, one of the main purposes of the APEX-AP1000 

test program was to investigate the sensitivity of ADS failure on core coolability. 

Based on the experiments run in APEX, the following conclusions can be made: 

(1) The APEX-AP1000 tests confirm significant liquid entrainment and carryover of water 

to the ADS4 system during and after ADS4 valve actuation. 

(2) Test results from NRC-AP1000-03, 05, and DBA-02 show that failure of ADS4 valves 

on the non-PZR side of the plant results in a greater delay in IRWST injection 

than failure of ADS4 valves on the PZR side of the plant. 

(3) Failure of two out of four ADS4 valves on the non-PZR side of the plant results in 

core uncovery (based on the two-phase mixture level) during simulated DEDVI line 

and cold leg breaks. 

(4) Review of the APEX-AP1000 test results conclusively show that the core remains cooled 

and heatup was not experienced for most of the beyond-DBA tests. 
 7-1 

7. REFERENCES 

1. 

2. Bessette, D., and di Marzo, M., “Transition from Depressurization to Long-Term Cooling 

in AP600 Scaled Integral Test Facilities,” Nuclear Engineering and Design, Vol. 188, pp. 331–344, 

Elsevier, 1999. 

3. Reyes, J.N., et al., “Final Report of NRC AP600 Research Conducted at Oregon State University,” 

NUREG/CR-6641, U.S. Nuclear Regulatory Commission, Washington, DC, 1999. 

4. Welter, K.B., et al., “Experimental Investigation and Theoretical Modeling of Liquid Entrainment 

in a Horizontal Tee with a Vertical Branch,” International Journal of Multiphase Flow, Vol. 30, 

No. 12, pp. 1451–1484, December 2004. 

5. 

6. 

7. U.S. Nuclear Regulatory Commission, “APEX Testing in Support of AP1000,” Memorandum from 

Jack Rosenthal (RES) to Jared Wermiel (NRR), Washington, DC, 2004. 

8. 

 

9. Zuber, N., “A Hierarchical, Two-Tiered Scaling Analysis: An Integrated Structure 

and Scaling Methodology for Severe Accident Technical Issue Resolution,” Appendix D 

to NUREG/CR-5809, U.S. Nuclear Regulatory Commission, Washington, DC, November 1991. 

10. 

11. American Nuclear Society, ”American National Standard for Decay Heat Power in Light-Water 

Reactors,” ANSI/ANS-5.1.24, La Grange, IL, 1979. 

12. Lay, J.F., Statistical Mechanics and Thermodynamics of Matter, Harper and Row Publishers, 

New York, pp. 453–456, 1990. 

13. Moody, F.J., Introduction to Unsteady Thermofluid Mechanics, A Wiley-Interscience Publication, 

New York, p. 60, 1990. 

14. Henry, F.E., and Fauske, H.K., “The Two-Phase Critical Flow of One-Component Mixtures 

in Nozzles, Orifices, and Short Tubes,” Journal of Heat Transfer, ASMAE Transactions, 

Vol. 93, No. 2, pp. 179–187, 1971. 

a,b,c 

a,b,c 

a,b,c 

a,b,c 

a,b,c 
 7-2 

15. Pimentel, D., “Two-Phase Fluid Break Flow Measurements and Scaling in the Advanced 

Plant Experiment (APEX),” Master’s Thesis, Oregon State University, May 1999. 

16. Kataoka, I., and Ishii, M., “Mechanistic Modeling and Correlations for Pool-Entrainment Phenomenon,” 

NUREG/CR-3304, U.S. Nuclear Regulatory Commission, Washington, DC, April 1983. 

17. Dallman, J.C. and Kirchner, W.L., “De-Entrainment Phenomena on Vertical Tubes in Droplet 

Cross Flow,” NUREG/CR-1421, U.S. Nuclear Regulatory Commission, April 1980. 

18. Chen, C.Y., “Filtration of Aerosols by Fibrous Media,” Chem. Rev. 55, p. 595, 1955. 

19. 

20. 

 

21. 

22. Bluman, G.W., and Cole, J.D., “Similarity Methods for Differential Equations,” Applied Mathematical 

Sciences 13, Springer-Verlag, New York, 1974. 

a,b,c 

a,b,c 

a,b,c 
 

APPENDIX A. 

... 
A-1 

APPENDIX A. SCALING ANALYSIS 

The following scaling analysis was performed by Reyes and was used to guide the APEXAP1000 test facility modifications. The basis for the majority of this work is the original AP600 

Scaling Analysis, WCAP-14270, which was performed to support the design of the existing 

Advanced Plant Experiment (APEX) facility at Oregon State University (OSU) [Ref. 1]. The 

analyses presented herein provide the scaling criteria that need to be implemented to model the 

AP1000 design changes in the APEX test facility. This analysis, in conjunction with the AP1000 

design information, has been used to identify and guide the design modifications needed to 

make the APEX facility suitable for AP1000 testing. 

A.1 Introduction

As part of the AP600 design certification, the APEX test facility was constructed at OSU. 

The APEX facility is a low-pressure, one-quarter height, integral systems test facility that can 

model the long-term core cooling phenomena associated with an advanced passive plant 

using passive safety systems. Extensive scaling studies were performed so that the AP600 

thermal-hydraulic phenomena of interest would be accurately modeled in the APEX facility. 

The APEX facility operates at one-half time scale and at a reduced pressure and temperature 

scale. Using the scaled parameters, a test in the APEX facility occurs twice as fast as a 

transient in the actual design. A total of 28 integral system tests were successfully performed for 

Westinghouse and the U.S. Department of Energy in support of the AP600 design certification. In 

addition, a total of 47 confirmatory tests were successfully performed for the NRC. 

A.1.1 Comparison of the AP600 and AP1000 Design Parameters 

The most significant difference between AP600 and AP100 is the increased core power. The 

reactor vessel volume remains unchanged. The SG tube surface area, PZR volume, and RCP 

flow have all increased to accommodate the increased core power. Table A.2 compares the 

changes to the passive safety system design. In particular, the volume of the CMTs and the 

IRWST has been increased; the flow capacity for the IRWST, ADS4, and PRHR has been 

increased by reducing the resistance in each line; and the PRHR surface area has also 

increased. The Westinghouse Plant Parameters document [Ref. 8] provides the details of the 

AP1000 design. Table A.1 compares the key design parameters of the AP1000 design with 

those of the AP600 reactor. 
A-2 

Table A.1 Important NSSS Design Changes [Ref. 8] 

Parameter AP600 AP1000 

Reactor Core Power, MWt 1933 3400 

Hot Leg Temperature, °F 600 610 

Number of Fuel Assemblies 145 157 

Active Fuel Length, ft 12 14 

Average Linear Power, kW/ft 4.10 5.707 

Average Heat Flux, Btu/(hr-ft2

) 142,869 198,933 

Steam Generator Heat Transfer Area per unit, ft2

 75,180 123,538 

Reactor Coolant Pump Flow, gpm 51,000 78,750 

Pressurizer Volume, ft3

 1,600 2,100 
A-3 

Table A.2 Changes to Passive Safety System Design Features

Component AP600 AP1000 Comment 

IRWST 

Volume, ft3 70,798 75,300 

Water Level, ft 28.0 28.79 

Line Resistance, % 100% 32% 

Design Flow Rate, % 100% 184% 

The IRWST level has been increased in the 

AP1000 by using more accurate level 

instruments. This permits a high operating 

level. 

CORE MAKEUP TANKS 

Number 2 2 

Volume, ft3 2000 2512 

Line Resistance, % 100% 64% 

Design Flow Rate, % 100% 124% 

The CMT volume and flow rate have been 

increased to provide additional safety injection 

flow. CMT elevations are maintained at the 

AP600 level. The duration of CMT injection 

is maintained similar to AP600. 

ACCUMULATORS 

Number 2 2 

Volume, ft3 2000 2000 

Pressure, psig 700 700 

The accumulators are the same as AP600. 

Accumulator sizing is based on LBLOCA 

performance and is determined largely on 

reactor vessels volume. 

AUTOMATIC DEPRESSURIZATION STAGES 1-3 

Location Top PZR Top PZR 

Configuration 6 paths 6 paths 

Vent Area, % 100% 100% 

The first three stages of ADS are the same as 

AP600. Their sizing basis is to reduce 

pressure to permit adequate injection from 

the accumulators and to permit transition to 

4th stage ADS. 

AUTOMATIC DEPRESSURIZATION STAGE 4 

Location Hot Leg Hot Leg 

Configuration 4 paths 4 paths 

Line size 10-inch 10-inch 

Nominal Vent Area, % 100% 176% 

Line Resistance 100% 28% 

Capacity 100% 189% 

The ADS4th stage vent area is increased 

more than the ratio of the core power. The 4th

stage ADS venting is the most important 

design feature to allow for stable 

IRWST/sump injection during long-term core 

cooling. 
A-4 

Table A.2 Changes to Passive Safety System Design Features

Component AP600 AP1000 Comment 

PASSIVE RHR HEAT EXCHANGER 

Type C-Tube C-Tube 

Surface Area, % 100% 122% 

Design Flow Rate, % 100% 174% 

Design Heat Trans., % 100% 172% 

The AP1000 PRHR HX retains the AP600 

configuration. The heat transfer surface area 

is increased by extending the length of the 

heat exchanger. The inlet and outlet piping 

has been increased resulting in higher flow 

rates. 

CONTAINMENT 

Diameter, ft 130 130 

Overall Height, ft 189.83 215.33 

Design Pressure, psig 45 59 

Net Free Volume, ft3 173 x106

 2.07 x106

The AP1000 containment volume and design 

pressure are increased to accommodate 

higher mass and energy releases. 

PASSIVE CONTAINMENT COOLING SYSTEM 

Water Storage Tank 

Volume (Top of Overflow), 

gallons

580,000 580,000 The PCS water storage tank was increased to 

accommodate higher flow rates. The PCS 

flow rates have been increased based on the 

increased in core power. 

A.1.2 Modifications to the APEX Test Facility 

Based on a review of the design changes listed in Tables A.1 and A.2, and their detailed description 

in the AP1000 Plant Parameters document [Ref. 8], the set of modifications listed in Table A.3 

were made to APEX to best simulate AP1000 thermal-hydraulic behavior. 
A-5 

Table A.3 Proposed Modifications to the APEX Test Facility 

Component Modification to APEX 

Reactor Power Increase core power by 67%. (Maximum of 1 MW) 

Pressurizer Increase PZR volume. 

Reduce PZR surge line diameter. 

Steam Generator Heat Transfer Area No change required for testing. 

Reactor Coolant Pump Flow No change required for testing. 

Core Makeup Tanks Increase CMT volumes by 25%. 

Reduce line resistance to 64% of original value. 

Accumulators No changes to original design. 

IRWST Increase IRWST liquid level. 

Automatic Depressurization Stages 1-3 No changes to original design. 

Automatic Depressurization Stage 4 Increase ADS4 flow area by 76%. 

Reduce line resistance to 28% of original value. 

Passive Residual Heat Removal (PRHR) 

Heat Exchanger (HX) 

Increase PRHR flow capacity by 74% by reducing line 

resistance. No change in surface area/tube number 

required for testing. 

Containment Sump Flood-Up Elevation Change flood-up elevation in primary sump tank. 

Passive Containment Cooling System Not part of APEX testing program. 

A.1.3 Rationale for APEX Modifications 

This section presents the rationale for the proposed modifications. An increase in the reactor 

core power is needed to preserve the proper power-to-volume scaling ratio. This is a key feature 

of the original scaling analysis and is essential to simulating decay heat behavior, loop natural 

circulation and primary system depressurization rates. The APEX power increase is limited to 

1 MW without having to replace the power distribution system. This power represents ~2.8% 

decay power for the AP1000 on a scaled basis. An increase in PZR volume is essential 

because it is a major source of liquid to the core during primary system blowdowns. The NRC 

identified a need for the reduction in PZR surge line diameter at the conclusion of the AP600 

research program [Ref. 2]. This modification is needed to ensure properly scaled PZR draining 

behavior during primary system blowdowns, and properly scaled surge line pressure drop 

behavior during ADS1-3 operation. 

Increases to the APEX SG tube volume and heat transfer area were not deemed necessary 

for several reasons. First, their tube volume is small compared to the remainder of the primary 

system. Second, they are oversized for the decay power operations involved in testing. That is, 

they have much more surface area than needed to remove core heat. One impact of not 

increasing the SG surface area is an increase in the core fluid temperature rise, but this is well 
A-6 

within the design constraints of the existing facility. The increased number of tubes, however, 

significantly reduces the primary loop pressure drop because of the increased number of 

parallel flow paths in the SG. As a result, the APEX facility is not able to match the scaled 

primary loop natural circulation flow rates that arise early in SBLOCA transients. However, 

all of the passive safety system and sump recirculation natural circulation flows can be properly 

scaled. 

An increase in RCP flow capacity was not necessary because the pumps are tripped at the start 

of each transient. Their primary contribution during testing is their resistance, which will be 

preserved. 

The CMT volumes will be increased because they are an essential part of the passive safety 

system response. ADS actuation is dependent on the CMT liquid volume. The CMT line resistance 

shall also be reduced to simulate the scaled CMT draining rates. CMT volume and flow rate 

are critical to the transition from ADS4 blowdown to IRWST injection. Full-pressure ADS4 

blowdown transition tests are being considered. 

The IRWST injection flow rate has a significant impact on core cooling. Therefore, the IRWST 

lines will be modified to increase their flow capacity to match the AP1000 design on a scaled basis. 

The onset of IRWST injection is dependent on the liquid level height in the tank. The APEX 

IRWST minimum liquid volume will be preserved to match the design change. However, 

the scaled liquid height and IRWST liquid volume cannot be exactly preserved simultaneously. 

The containment sump flood-up curb height will be adjusted accordingly. 

The fourth-stage valves of the automatic depressurization system (ADS4) ensure that the 

primary system pressure will be reduced below the IRWST liquid level head so that injection 

can begin. The increased flow area is needed to properly simulate the AP1000 ADS4 operation. 

In addition, the resistance of each ADS4 line will be decreased accordingly to properly match 

the pressure drop behavior under non-choked flow conditions. 

Increases to the PRHR heat exchanger tube volume and surface area were not deemed 

necessary for the experiment. As with the SG, the PRHR tube volume is small compared to 

the primary loop volume. Furthermore, it is oversized for the decay power operations involved 

in testing. That is, they have much more surface area than needed to remove core heat. 

The impact of not increasing the PRHR will be an increase in the core fluid temperature rise. 

This is well within the design constraints of the existing facility. Properly scaled natural 

circulation flow rates are ensured by adjusting the PRHR loop resistance. 

The ADS1-3 valves and the accumulators remain the same as the original AP600 design. 

Therefore, these components will not be modified for the AP1000 testing program. 
A-7 

A.1.4 Applicability of the AP600 Scaling Analysis 

The original scaling analysis performed in support of the design of the APEX test facility for 

AP600 certification is documented in WCAP-14270. The comprehensive report represents 

the first successful application of the hierarchical two-tiered scaling (H2TS) methodology 

for the design and construction of an experimental test facility. Westinghouse issued the report 

in January 1995 [Ref. 1]. The objective of the scaling study was to obtain the physical dimensions 

of a test facility that would simulate the flow and heat transfer behavior of importance to the AP600 

passive safety system operation. The report includes all of the scaling criteria that were needed 

to define the geometry and operating conditions of the reduced-scale APEX test facility. 

Figure A.1 shows the general scaling methodology used to design the APEX test facility for the 

AP600 test program. First, the general objective of the test program was stated. The general 

objective was to assess the passive safety system operation for loss-of-coolant-accidents 

(LOCAs). A small-break loss-of-coolant accident (SBLOCA) phenomena identification and 

ranking table (PIRT) was then developed to identify the important thermal-hydraulic phenomena 

that should be preserved in the test facility. The next step was to perform a scaling analysis 

for each of the AP600 operational modes during the LOCA. The four key modes of operation 

identified were natural circulation, system depressurization, venting/draining/injection, and 

recirculation. The analyses resulted in a set of dimensionless groups and similarity criteria that 

had to be preserved in the test facility. These similarity criteria were derived from the governing 

equations of mass, momentum, and energy conservation for the system, subsystems, and 

components for each of the four operating modes. The similarity criteria were used to obtain 

the scaling ratios needed to design the test facility geometry and operating conditions. 

As shown in Figure A.1, the scaling ratios were developed for a specific set of experiment 

objectives, which produce a specific set of system responses (i.e., operation modes and 

thermal-hydraulic phenomena) within a specific plant geometry. A review of the scaling ratios 

reveals that they are also directly applicable to the design of the APEX-AP1000 test facility 

for the following reasons. 

First, the experiment objectives of the APEX-AP1000 test program are a subset of the AP600 

experiment objectives. Therefore, the same operation modes and thermal-hydraulic 

phenomena will be produced. 

Westinghouse recently issued an AP1000 PIRT and Scaling Assessment, which reviews the 

original scaling analysis effort and assesses the applicability of the original AP600 test program 

to the AP1000 design [Ref. 6]. Westinghouse concluded that the AP1000 PIRT was the same 

as the AP600 PIRT, with the exception of the potential higher importance of HL entrainment 

during the post-ADS4 actuation phases of the SBLOCA. Therefore, entrainment scaling of 

the post-ADS4 actuation phase was of particular interest to the present study. 
A-8 

Specify Experiment

Objectives

 SBLOCA PIRT

(3) Perform Scaling Analysis

(2)

(1)

Evaluate Key T/H PIRT

Processes to Prioritize System

Design Specification

APEX Test Facility Design

Specifications and Q/A Critical

Attributes

(8)

(7)

Natural

Circulation

Scaling

Analysis

System

Depressurization

Scaling Analysis

Venting,

Draining and

Injection

Scaling Analysis

Develop Π

Groups and

Similarity

Criteria

Develop Π

Groups and

Similarity

Criteria

Significant

Scaling

Distortions in

Important P

Groups?

Significant

Scaling

Distortions in

Important P

Groups?

System Design

Specifications

System Design

Specifications (6)

(5)

(4)

No No

Yes Yes

Recirculation

Scaling Analysis

Develop Π

Groups and

Similarity

Criteria

Significant

Scaling

Distortions in

Important P

Groups?

System Design

Specifications

No

Yes

Develop Π

Groups and

Similarity

Criteria

Significant

Scaling

Distortions in

Important P

Groups?

System Design

Specifications

No

Yes

Figure A.1 General Scaling Methodology for the APEX Test Facility [Ref. 1] 

Second, the AP1000 design is geometrically similar to the AP600 design. This means that there 

is a one-to-one correspondence of primary loop components, arranged in the same sequence, 

and all of the important AP1000 geometric scales, (i.e., cross-sectional flow areas, volumes, 

lengths) can be related to the AP600 design by constant scaling factors. 

In conclusion, the general scaling methodology remains the same. Therefore, the scaling ratios 

obtained in the original scaling analysis, with a few exceptions, can be directly applied to the 

design of an APEX-AP1000 test facility. The following section describes the scope of the 

scaling assessment for the APEX-AP1000 test facility. 
A-9 

A.1.5 Scope of the Current Scaling Assessment 

The following analyses have been included in the OSU APEX-AP1000 Scaling Report to guide 

the design of the APEX test facility modifications presented in Table A.3: 

• Core Decay Power Scaling Analysis 

• IRWST, CMT, and PZR Liquid Volume and Height Scaling Analysis 

• IRWST, CMT, PRHR, and Sump Recirculation Line Scaling 

• Pressurizer Surge Line Scaling Analysis 

• RCS Depressurization Scaling Analysis 

• ADS4 Blowdown Scaling Analysis 

• Upper Plenum Pool Entrainment Scaling Analysis 

The core decay power analysis was needed to develop the revised decay power algorithm for 

the tests. The original power scaling ratios were used in the decay power scaling analysis. 

The IRWST, CMT, and PZR volume scaling analyses were straightforward and implemented 

the original volume scaling ratios. Similarly, the original line resistance scaling ratios were used 

to scale the IRWST, CMT, PRHR, and sump recirculation line pressure drops. The PZR surge 

line scaling analysis represents a change to the original scaling analysis. That is, it is intended 

to provide a better representation of the PZR draining process. The PZR surge line scaling 

analysis resulted in a new diameter scaling ratio for the PZR surge line. The ADS4 blowdown 

scaling analysis represents an improved depressurization scaling analysis technique based on 

the energy equation. Of particular interest to this study, is the transition from ADS4 blowdown 

to IRWST injection. The upper plenum pool entrainment scaling analysis is intended to provide 

a better simulation of the upper plenum entrainment and de-entrainment processes. This 

analysis also includes an upper core plate flow area scaling analysis to better simulate countercurrent flooding in the upper core plate. 
A-10 

A.2 Hierarchical Two-Tiered Scaling (H2TS) Methodology

The H2TS methodology was used to develop the similarity criteria necessary to scale the 

APEX-AP1000 systems and processes of importance to an SBLOCA transient. The H2TS 

methodology was developed by the NRC and is fully described in Appendix D to NUREG/CR-5809 

[Ref. 9]. This is the same methodology that was used to develop the similarity criteria for the 

original APEX facility. 

Figure A.2 is taken from NUREG/CR-5809. It presents the four basic elements of the H2TS 

analysis methodology. The first element consists of subdividing the plant into a hierarchy of 

systems. Each system was subdivided into interacting subsystems, which were further 

subdivided into interacting modules, which were further subdivided into interacting constituents 

(materials), which were further subdivided into interacting phases (liquid, vapor, or solid). Each 

phase could be characterized by one or more geometrical configurations, and each geometrical 

configuration could be described by three field equations (mass, energy, and momentum 

conservation equations). Each field equation could be characterized by several processes. 

After identifying and subdividing the system of interest, the next step was to identify the scaling 

level at which the similarity criteria should be developed. This was determined by examining 

the phenomena being considered. For example, if the phenomenon being considered involved 

mass, momentum, or energy transport between materials such as water and solid particles, 

the scaling analysis was performed at the constituent level. If the phenomenon of interest 

involved mass, momentum, or energy transport between vapor and liquid, the scaling analysis 

was performed at the phase level. Therefore, identifying the scaling level depended on 

the phenomenon being addressed. 

Thermal-hydraulic phenomena involving integral RCS interactions, such as primary system 

depressurization or loop natural circulation, were examined at the “system” level. Thermalhydraulic phenomena, such as SG heat transfer, were examined at the “subsystem” level. 

Specific interactions between the steam-liquid mixture and the stainless steel structure were 

examined at the “constituent” level. 

Chapter 3 presents the scaling analyses that were performed for the different APEX-AP1000 

test facility modifications. It identifies the thermal-hydraulic phenomena of interest, the system 

level (i.e., control volume) at which the analysis was performed, the geometrical configuration, 

the applicable balance equations and the processes important to the thermal-hydraulic 

phenomena of interest. 

The H2TS methodology required performing a “top-down” (system) scaling analysis, which 

examined the synergistic effects on the system caused by complex interactions between the 

constituents which are deemed important by the PIRT. Its purpose was to use the conservation 

equations at a given scaling level to obtain characteristic time ratios and similarity criteria. 

It also identified the important processes to be addressed in the bottom-up scaling analysis. 

The H2TS methodology also required performing a “bottom-up” (process) scaling analysis, 

which developed similarity criteria for specific processes, such as flow pattern transitions and 

flow-dependent heat transfer. The focus of the bottom-up scaling analysis was to develop 

similarity criteria to scale individual processes of importance to system behavior as identified by 

the PIRT. 
A-11 

A.2.1 Time Ratios 

The basic objective of the H2TS methodology was to develop sets of characteristic time ratios 

for the transfer processes of interest. This was done by writing the control volume balance 

equations for each constituent “k” as follows: 

= [ ] Q (j A ) + S dt

dV

k k kn kn k

k k ∆ ψ ± Σ ψ (A.1) 

where 

∆[ ψ ] [ ψ ] [ ψ ] k k k k in Qk k out Q = Q - (A.2) 

In equation (A.1) the ψk term represents the conserved property; ψk = ρ, ρu or ρε (mass, 

momentum, or energy per unit volume); Vk is the control volume; Qk is the volumetric flow rate; 

jkn is the flux of property ψk transferred from constituent “k” to “n” across the transfer area Akn. 

Hence, ∆[Qkψk] represents the usual mass, momentum, or energy convection terms, and ΣjknAkn

represents transport process terms such as condensation, and Sk represents the distributed 

sources, such as decay power or body forces acting internal to the control volume. Equation 

(A.1) can be put in dimensionless form by specifying the following dimensionless groups in 

terms of the constant initial and boundary conditions: 

S

S ,S = A

A , A = j

j , j = Q

Q , = , Q = V

V V =

k,0

+ k

k

kn,0

+ kn

kn

kn,0

+ kn

kn

k,0

+ k

k

k,0

+ k

k

k,0

+ k

k

ψ

ψ

ψ (A.3) 

Substituting these groups into equation (A.1) yields: 

= Q [Q ] ( ) j A j A + S S dt

dV

V + k,0 k +

kn +

kn,0 kn,0 kn

+

k

+

k,0 k,0 k

+

k +

k

k,0 k,0 ψ ∆ ψ ± Σ ψ

ψ (A.4) 

Dividing both sides of this equation by Qk,0 ψk,0 yields: 

= [Q ] j A + S dt

dV + sk k +

kn + kn kn

+

k

+

k

+

k +

k

k ∆ ψ ± ΣΠ Π ψ

τ (A.5) 

where the residence time of constituent “k” is 

Q

V =

k,0

k,0

τk (A.6) 
A-12 

Figure A.2 Flow Diagram for the Hierarchical, Two-Tiered Scaling Analysis [Ref. 9] 
A-13 

and the characteristic time ratio for a transfer process between constituents “k” and “n” is given by: 

ψ

Π

k,0 k,0

kn,0 kn,0

kn Q

j A = (A.7) 

The characteristic time ratio for the distributed source term within the control volume is given by: 

ψ

Π

k,0 k,0

k,0

sk Q

S = (A.8) 

Because each transfer process has a characteristic time ratio, it is possible to rank the importance 

of each process by comparing the time ratios. If a specific transfer process is to have the same 

effect in the prototype and the model, the characteristic time ratios must be preserved. 

A.2.2 Process Ranking Using Characteristic Time Ratios 

Let us define M[ (Pi,j), (Pi+1,j),...(PNi,Nj)] as the set of time ratios that characterize all of the 

individual processes that occur during the evolution of a transient. The subscripts i, j, Ni,Nj

identify the specific process, the hierarchical level, the total number of specific processes, 

and the total number of hierarchical levels, respectively. 

Because of differences in geometrical scale and fluid properties, it is impossible to exactly 

duplicate the “time ratio set” for the full-scale prototype, Mp, in a reduced-scale model. That is, 

exact similitude for all processes cannot be preserved; therefore: 

Mp ≠ Mm (A.9) 

The subscript, p, refers to the full-scale prototype and the subscript, m, refers to the reducedscale model. 

It is possible to design a reduced-scale test facility that preserves the similitude of a subset 

of time ratios T [Πi,j], that characterize the processes of greatest importance to the transient. 

This optimizes the model design to investigate the important processes, while distorting 

the less-important processes. 

To determine which processes govern the overall evolution of a transient, numerical estimates 

of the characteristic time ratios for the prototype and the model must be obtained for each 

hierarchical level of interest. Physically, each characteristic time ratio, Πi, is composed of 

a specific frequency, ωi, which is an attribute of the specific process, and the residence time 

constant, τcv, for the control volume. That is: 

Πi = ωi τcv (A.10) 

The specific frequency defines the mass, momentum, or energy transfer rate for a particular 

process. The residence time defines the total time available for the transfer process to occur 

within the control volume. A numerical value of: 

Πi < 1 (A.11) 
A-14 

means that only a small amount of the conserved property would be transferred in the limited 

time available for the specific process to evolve. As a result, the specific process would not be 

important to the overall transient. A numerical value of: 

Πi ≥ 1 (A.12) 

means that the specific process evolves at a high enough rate to permit significant amounts of 

the conserved property to be transferred during the time period, τcv. Such processes would be 

important to the overall transient behavior. 

A.2.3 Scaling Criteria Development 

The scaling analysis results in a set of characteristic time ratios (dimensionless Π groups) and 

similarity criteria for each mode of operation. Because it is impossible to identically satisfy all of 

the similarity criteria simultaneously, the set only includes those criteria that must be satisfied 

in order to scale the most important phenomena identified by the PIRT. 

Scaling criteria were developed by requiring that the characteristic time ratios for a subset of 

specific processes in the prototype (usually those of greatest importance) are matched in the 

model at each hierarchical level. That is, 

{ Π } { Πi} i m p T = T (A.13) 

These criteria were satisfied by adjusting the physical geometry, fluid properties, and operating 

conditions of the model, thereby optimizing the model design for the specific process of interest. 

A.2.4 Evaluation of Scale Distortion 

The scaling criteria were evaluated to determine if the scale model geometry, boundary 

conditions, or operating conditions would introduce significant scaling distortions. Distortions 

were also evaluated relative to other modes of operation. 

The effect of a distortion in the model for a specific process can be quantified as follows: 

[ ] [ ]

[ ] Π

Π Π

i p

i p i m - DF = (A.14) 

The distortion factor, DF, physically represents the fractional difference in the amount of 

conserved property transferred through the evolution of a specific process in the prototype to 

the amount of conserved property transferred through the same process in the model during 

their respective residence times. A distortion factor of zero would indicate that the model ideally 

simulates the specific process. A distortion factor of +0.05 would indicate that the specific 

process in the model transfers 5 percent less of the conserved property (on a scaled basis) than 

the same process in the prototype. The distortion factor can also be written as: 

[ω ] [τ ] i R cv R DF =1- (A.15) 
A-15 

or 

DF =1-[II ]i R (A.16) 

The degree to which a specific transfer process could impact a particular transient can be 

determined by comparing the maximum characteristic time ratio for each of the transfer 

processes that arise during the transient. 

Upon satisfying the important scaling criteria, the component geometries and operating 

conditions were specified for the APEX-AP1000 test facility. 
A-16 

A.3 Scaling Analysis for the APEX-AP1000 Modifications

This chapter presents the scaling analyses performed to obtain the component geometries and 

system operating conditions for the APEX-AP1000 test facility. Tables A.4 and A.5 present the 

scaling ratios that were derived in WCAP-14270 and used for the design of the APEX-AP1000 

test facility modifications. 

Table A.4 APEX-AP1000 Primary Loop Scaling Ratios from WCAP-14270 [Ref. 1]

Parameter Desired Scaling Ratio 

CORE 

Length Ratio 

Flow Area Ratio 

Fluid Velocity Ratio 

Power Ratio 

Fluid Residence Time Ratio 

Mass Flow Ratio 

Volume Ratio 

Power/Volume Ratio 

HOT AND COLD LEGS 

Length Ratio 

Diameter Ratio 

Flow Area Ratio 

Volume Ratio 

Fluid Velocity Ratio 

Fluid Residence Time Ratio 

Mass Flow Ratio 

PRESSURIZER 

Volume Ratio 

Liquid Height Ratio 

a,b,c 
A-17 

Table A.5 APEX-AP1000 Passive Safety System Scaling Ratios 

from WCAP-14270 

Parameter Desired Scaling Ratio 

CORE MAKEUP TANK 

Length Ratio 

Flow Area Ratio 

Draining Mass Flow Rate Ratio 

Volume Ratio 

ACCUMULATOR 

Volume Ratio 

Mass Flow Ratio 

IRWST 

Length Ratio 

Flow Area Ratio 

Draining Mass Flow Rate Ratio 

Volume Ratio 

LOWER CONTAINMENT SUMPS 

Length Ratio 

Flow Area Ratio 

Draining Mass Flow Rate Ratio 

Volume Ratio 

The horizontal HL and CL scaling ratios were established based on preserving flow regime 

transitions in the legs. 

The 1:2 time scaling requirement has been imposed on all of the system components and 

operations. The scaling ratios presented in Tables A.4 and A.5 have been applied in the 

following sections to obtain the revised APEX component geometries as needed to match the 

new AP1000 design. 

a,b,c 
A-18 

A.3.1 Core Decay Power and Flow Area Scaling Analysis 

One of the key changes to the AP1000 is the increase in core thermal power from 1,933 MW 

thermal to 3,400 MW thermal. Three modifications were made to the APEX test facility as a 

result of the power change: 

• A higher power core was installed. 

• The core flow area was adjusted. 

• The decay power algorithm was modified. 

A.3.1.1 Maximum Core Power 

Based on the power scaling ratio given in Table A.4, the new core power needed in APEX to 

match a 3% decay power in the AP1000 was: 

 (A.17) 

This resulted in a desired power in APEX of 1.06 MW thermal. A new 48 rod bundle core with a 

[ ]a,b,c thermal power was installed. This represents a [ ]a,b,c decay power. 

A.3.1.2 Core Flow Area 

Based on the core flow area ratio given in Table A.4, the new core flow area needed in APEX to 

properly match the scaled core flow area in AP1000 was: 

 (A.18) 

The AP1000 effective core flow area is [ ]a,b,c. Using the scaling ratio in equation 

(A.18), the desired APEX-AP1000 core flow area is [ ]a,b,c or [ ]a,b,c. Using 48 

heaters, each having a [ ]a,b,c diameter, results in an actual core flow area of [ ]a,b,c. 

A.3.1.3 Decay Power Algorithm 

Establishing the initial conditions for ADS4 operation and long-term cooling requires that the 

total energy input into the system, to the point of reaching the transition pressure be properly 

scaled. The integrated core power ratio is written as follows: 

q dt

q dt

E =

core,p p

t

o

core,m m

t

o

core,R p

m

∫

∫ (A.19) 

where tm is the time it takes in APEX-AP1000 to reach the ADS4 opening pressure, and tp is the 

time it takes in AP1000 to reach the ADS4 opening pressure. Since the ideal core power 

scaling ratio given in Table A.4 is [ ]a,b,c and the ideal time scaling ratio is 1:2, the integrated 

core power ratio should be [ ]a,b,c. 
A-19 

The 1979 Decay Heat Standard for Light-Water Reactors was used to model the decay power 

curve for the AP1000 [Ref. 11]. Figure A.3 shows the decay power algorithm for the APEXAP1000 test facility. The power algorithm is designed to match the ideally scaled AP1000 

decay power after ~160 seconds. Similarly, as shown in Figure A.4, the total integrated decay 

power matches after 160 seconds. 

Figure A.3 Comparison of the APEX-AP1000 Decay Power Algorithm 

and the Ideally Scaled AP1000 Decay Power Curve 

a,b,c 
A-20 

Figure A.4 Comparison of the APEX-AP1000 Integrated Decay Power 

to the Ideally Scaled AP1000 Decay Power 

a,b,c 
A-21 

A.3.2 IRWST, CMT, Sump, and Pressurizer Geometric Scaling Analysis 

Table A.6 presents the results of applying the scaling ratios of Table A.5 to the AP1000 IRWST, 

CMT, sump, and PZR. 

Table A.6 Results of IRWST, CMT, Sump, and Pressurizer Geometry Scaling 

Parameter Scaling 

Ratio AP1000* APEX Ideal 

APEX 

Actual Units AP1000* APEX Ideal 

APEX 

Actual 

Units

IRWST Liq. Volume 

(min.) 

 

IRWST Liq. Level (min.) 

IRWST Liq. Surface Area 

IRWST Level Setpoint for 

Sump Recirculation 

 

IRWST Liq. Head on DVI 

CMT Volume 

CMT Internal Diameter 

CMT Internal Height 

CMT Volume for ADS-1 

Actuation 

 

CMT Volume for ADS4 

Actuation 

Sump Liq. Volume Below 

DVI Elevation 

Difference Between DVI 

and Sump Flood-up 

Elevations** 

Sum Liq. Surface Area 

Pressurizer Volume 

Pressurizer Liq. Volume 

Pressurizer Inside 

Diameter 

 

* AP1000 Plant Parameters Document [Ref. 8] 

** Assumes break occurs in loop compartment 

a,b,c 
A-22 

A.3.3 ADS4, IRWST, CMT, PRHR, and Sump Recirculation Line Scaling 

A detailed analysis of the pressure drops in the passive safety system balance and injection 

lines was provided in Section A.7 of WCAP-14270 [Ref. 1]. The pressure drop for a section of 

pipe having a constant diameter is given by the following equation: 

+ K v D

fl

2

1 P = 2 ⎟ρ

⎠

⎞ ⎜

⎝

⎛ ∆ (A.20) 

where f is the Darcy friction factor, l/D is the piping length to diameter ratio, K is the form loss 

coefficient due to fittings and bends, and v is the fluid velocity in the section of pipe. For singlephase, non-choked, fluid flow with fluid property similitude, the ratio of the model to the 

prototype can be written as: 

P = F v

2 ∆ R T,R R (A.21) 

In this equation, FT,R is the ratio of the total friction and form loss coefficients. To preserve the 

line pressure drops, it is required that the pressure drop ratio, ∆PR, be 0.25 and that the velocity 

ratio, vR, be 0.5. Substituting these scaling ratios into the equation above yields the following 

requirement: 

 FT,R = 1 (A.22) 

For the piping under consideration, the form loss coefficients, K, dominate. That is, the 

presence of valves, elbows, and other fittings make the value of K much larger than the value of 

fl/D. In light of this fact, the line resistance can be adjusted using a flow orifice. 

The AP1000 line resistance data, R (ft/gpm2

), is expressed in terms of head loss over the 

square of volumetric flow rate. That is, the line resistance is expressed in terms of measured 

quantities as follows: 

g Q

g P

R = 2

c

ρ

∆

 (A.23) 

The resistance is related to the friction and form loss factors as follows: 

a

1

+ K

D

fl

2

1 R = 2 ⎟

⎠

⎞ ⎜

⎝

⎛ (A.24) 

where a is the flow area of the pipe. Therefore, the units of line resistance are often expressed 

as 1/ft4

. The flow area scaling ratio is given as 1:48. Therefore, the line resistance scaling ratio 

for single-phase fluid under non-choked flow conditions is given as: 

 RR = 482

 = 2304 (A.25) 
A-23 

Table A.7 lists the range of line resistance values required for the IRWST and sump 

recirculation lines in the APEX-AP1000 test facility. Table A.8 presents similar values for the 

CMT and PRHR lines. 

The ADS4/HL nozzle diameter and resistance was scaled as with previous lines. However, the 

two ADS4 branch lines that comprise a single ADS4 train on a HL require a different scaling 

approach because the intent is to model two ADS4 branch lines with a single line in the APEXAP1000 test facility. That is, the flow area of one APEX ADS4 branch line must equal the 

scaled flow area of two AP1000 branch lines: 

48

2a 1000

a = AP

APEX (A.26) 

In terms of a diameter scaling ratio, the ratio of a single ADS4 branch line diameter in APEX to a 

single ADS4 branch line in the AP1000 is: 

24

1

DA DS,R = (A.27) 

Table A.9 presents the ADS4 line diameters and resistance assuming that two AP1000 ADS4 

branch lines are modeled with a single branch line in APEX. 
A-24 

Table A.7 IRWST and Sump Recirculation Line Resistance Values (ft/gpm2) [Ref. 8]

 AP1000 IDEAL APEX 

Path/Line I.D. (in) Min Nom Max I.D. (in) Min Nom Max 

IRWST TO SUMP TEE 

Line A 

Line B 

SUMP TEE TO MOV ISOLATION VALVE 

Line A 

Line B 

MOV ISOLATION VALVE TO CHECK SQUIB VALVE 

Line A 

Line B 

CHECK/SQUIB VALVES PARALLEL PATHS 

Line A 

Line B 

CHECK/SQUIB VALVES TO DVI INJECTION LINE TEE 

Line A 

Line B 

IRWST INJECTION LINE TO DVI NOZZLE 

Line A 

Line B 

SUMP RECIRC. PATH TO IRWST INJECTION LINE TEE 

Line A 

Line B 

IRWST DRAIN TO CONTAINMENT 

Line A 

Line B 

a,b,c 
A-25 

Table A.8 CMT and PRHR Line Resistance Line Values (ft/gpm2) [Ref. 8]

 AP1000 IDEAL APEX 

Component I.D. (in) Min Nom Max I.D. (in) Min Nom Max 

CORE MAKEUP TANK 

CL Balance Line 

CMT to ACC Tee 

ACC to DVI 

PRHR 

RCS HL to ADS4 

Reducer 

 

ADS4 Reducer to 

PRHR Reducer 

 

PRHR Reducer to 

PRHR HX Inlet 

PRHR HX Inlet 

PRHR HX Tube 

PRHR HX Outlet 

PRHR HX Outlet to 

SG 

 

Table A.9 ADS-4 Line Resistance Values(ft/gpm2

) [Ref. 8] 

ADS4 AP1000 IDEAL APEX 

I.D. (in) Min Nom Max I.D. (in) Min Nom Max 

ADS4/HL Nozzle 

ADS4 Branch Line 

ADS4-1 (50%) 

ADS4-1 100%) 

ADS4-2 (50%) 

ADS4-2 (100%) 

a,b,c 

a,b,c
A-26 

A.4 Pressurizer Surge Line Scaling

This section represents a change to the original PZR surge line scaling analysis presented in 

WCAP-14270, which focused on preserving the two-phase flow regime transitions. However, 

a recent NRC analysis performed by Bessette and Di Marzo suggests that preserving the PZR 

draining behavior is more important to the outcome of SBLOCA transients [Ref. 2]. Based on a 

thorough review of the NRC analysis, the PZR surge line was modified to properly simulate the 

AP1000 PZR draining behavior. 

Bessette and Di Marzo expressed the governing momentum balance equation for the PZR 

surge line filling and draining process, in dimensionless form, as follows: 

Z& Z& + AZ = B (A.28) 

where: 

⎟

⎠

⎞ ⎜

⎝

⎛

⎟

⎟

⎠

⎞

⎜

⎜

⎝

Φ ⎛

ρ

+ K

D

Lf

S

S

2

+1

g

C

A =

SL

PZR

2 2

LO

 (A.29) 

⎟

⎠

⎞ ⎜

⎝

⎛

⎟

⎟

⎠

⎞

⎜

⎜

⎝

Φ ⎛

α

ρ ρ

+ K

D

Lf

S

S

2

H

H - (1- ) gHP - g

C

B =

SL

PZR

2 2

LO

PZR

SL

PZR

PZR

 (A.30) 

and 

S

S C = g

SL

PZR ρ (A.31) 

Integrating equation (A.28) yields the PZR surge line filling and draining period: 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

+ ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

τ

α= α=0.8

1/ .2

0

1/ 2 1/ 2

PZR

SL A

B

A

B

g

H = 2 (A.32) 

For the APEX-AP1000, a 1:2 scale time period is required. Preserving a one-half time scale 

PZR surge line filling and draining period requires that: 

2

1 (H ) = 1/ 2

PZR R (A.33) 

and that: 

=1

A

B1/ 2

R

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ (A.34) 
A-27 

The APEX-AP1000 PZR liquid height, HPZR, is [ ] a,b,c. Therefore, the scaling criterion given 

by equation (A.33) is met. 

Table A.10 presents the results obtained for equation (A.34). The value is close to unity. It has 

been assumed that the ratio of the Martinelli-Nelson two-phase flow multiplier, ΦLR , is unity. 

In addition to preserving the surge line drain and fill rates, it is desired that the pressure drop 

ratio in the surge line be [ ] a,b,c. That is: 

 (A.35) 

This requires: 

 (A.36) 

As shown in Table A.10, for the line size selected, the friction factor ratio is [ ] a,b,c and the 

velocity ratio in the surge line is [ ] a,b,c. This yields the desired DP ratio of [ ] a,b,c. 

Table A.10 Results of Pressurizer Surge Line Scaling Analysis 

Parameter AP1000 APEX 

Roughness (in) 

ID (in) 

Relative Roughness 

Reynolds Number 

Friction Factor 

Calculated Pressure Drop Ratio 

Friction Factor Ratio 

(fL/d) Ratio 

Pressurizer ID (in) 

HPZR(ft) 

Head at 1 atm (ft) 

HsL (ft) 

AR (Bessette & di Marzo) 

BR (Bessette & di Marzo) 

(BR)

1/2/A 

a,b,c 
A-28 

A.5 Reactor Coolant System Depressurization Scaling Analysis

This section presents a revised depressurization scaling analysis for the APEX-AP1000 test 

facility. It represents a more comprehensive analysis than that presented in the original 

WCAP-14270. Figure A.5 presents the scaling analysis flow diagram for primary system 

depressurization. First, the governing set of equations for the depressurization of a two-phase 

fluid system was developed. This resulted in obtaining a depressurization rate equation. Next, 

a top-down system level scaling analysis was performed for the RCS, assuming multiple vent 

and injection paths. This included the development of scaling criteria for sizing the break and 

ADS flow areas. Last, a bottom-up analysis was performed to describe the scaling of local 

transport processes, such as the critical flow through the breaks and ADS valves. 

Figure A.5 Flow Chart of the Primary Side Depressurization Scaling Analysis 
A-29 

A.5.1 Reference Pressure for LOCA Transients 

A variety of primary loop blowdown tests will be performed in APEX. The integral system tests 

will be performed under pressure-scaled conditions to observe all of the passive safety system 

functions. A series of full-pressure tests shall also be performed to examine the transition from 

ADS4 operation to IRWST injection. 

Figure A.6 presents typical trends for the RCS and SG pressures during a SBLOCA transient in 

the APEX facility. The opening of a break is immediately followed by a period of subcooled 

blowdown, during which the primary system depressurizes to a pressure equivalent to the SG 

PORV setpoint. Hence, the primary and secondary system pressures will meet at the onset of 

saturation conditions in the primary as shown in the figure. The AP600 scaling analysis was the 

first research to recognize that the onset of primary loop saturation offered a common reference 

point for the purpose of scaling. This fact, coupled with the concept that the depressurization of 

systems under phase equilibria conditions exhibits self-similarity, offered the possibility of 

performing pressure-scaled testing. 

Figure A.6 Primary and Secondary Side Pressures During a Typical 2-Inch Break LOCA in APEX 

For the AP1000, the reference pressure for the onset of saturation conditions in the primary loop 

coincides with the SG PORV pressure setpoint of 1,200 psia. For the APEX-AP1000, this value 

will be set at 345 psia. The goal of the pressure-scaled testing is to provide the full range of 

passive safety system behavior for a given transient in the proper event sequence. This data 

can be used to benchmark the thermal-hydraulic safety analysis codes. However, by scaling 

the break and ADS1-3 flow areas relative to the saturation reference pressure, the timing of the 

subcooled blowdown portion of the transient will not proceed at one-half time scale. For rapid 

blowdowns, such as the DEDVI line break, this effect is negligible. This was demonstrated by 

the excellent comparisons obtained between the APEX and the full-pressure ROSA AP600 

DEDVI tests. 
A-30 

The AP600 test program revealed that the transition from ADS4 operation to the start of IRWST 

injection was important to core cooling. Therefore, this test program will include a series of fullpressure ADS4 blowdown tests to carefully examine the ADS4 transition under prototypic fluid 

pressures and temperatures. 

A.5.2 Governing Equations for RCS Two-Phase Fluid Depressurization 

The mass conservation equation for an RCS control volume undergoing a depressurization 

event is given by: 

= m - m dt

dM ∑ & in ∑ & out (A.37) 

where M is the fluid mass within the RCS and m& represents the mass flow rate entering or 

leaving the RCS. The energy conservation equation for the RCS fluid is expressed as follows: 

() () dt

dV = mh - mh + q + q + q - P dtdU

in out SG core loss ∑ & ∑ & (A.38) 

where U is the bulk internal energy of the fluid within the RCS; h is the enthalpy of the fluid 

entering or leaving the RCS; qSG, qcore and qloss are the SG energy transfer rate, core power, and 

heat loss, respectively; P is the RCS pressure; and V is the RCS volume. The specific internal 

energy and specific volume are defined, respectively, as follows: 

M

U

e = (A.39) 

M

V

v = (A.40) 

The total change in specific internal energy is written in terms of partial differentials with respect 

to pressure and specific volume, as follows: 

 dv

v

e dP + Pe de =

v p

⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ (A.41) 

Substituting equation (A.39) into (A.38) yields: 

() () dt

dV = mh - mh + q + q + q - P dtdMe

in out SG core loss ∑ & ∑ & (A.42) 

Expanding the term on the left-hand side of equation (A.42), substituting equation (A.37), and 

rearranging yields: 

( )( ) ( )( ) dt

dV = m h - e - m h - e + q + q + q - P dtde M in in out out SG core loss ∑ & ∑ & (A.43) 
A-31 

In equation (A.43), it has been assumed that hin is the same for all the injection locations, and 

hout is the same for all the vent paths. Substituting equation (A.41) into (A.43), and rearranging 

yields: 

( )( ) ( )( )

dt

dv

 v

e - M

dt

dV + q + q + q - P

= m h - e - m h - e

dt

dP

P

e M

p

SG core loss

in in out out

v

⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

∑ ∑ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ & &

 (A.44) 

Using equation (A.40) and the mass conservation equation, the last term on the right-hand side 

of equation (A.44) is written as: 

 ( ) m - m

v

e - v

dt

dV

 dv

e = dtdv

 v

e M in out

p p p

⎟ ∑ & ∑ & ⎠

⎞ ⎜

⎝

⎛

∂

∂ ⎟

⎠

⎞ ⎜

⎝

⎛ ∂ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ (A.45) 

Substituting back into equation (A.44) yields: 

( )

( )

dt

dV

v

e

+ q + q + q - P + 

v

e - m h - e + v 

v

e = m h - e + v dt

dP

 P

e M

p

SG core loss

p

out out

p

in in v

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ ∑

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ ⎟ ∑ ⎠

⎞ ⎜

⎝

⎛

∂

∂

&

&

 (A.46) 

which is the “depressurization rate equation.” For the RCS control volume, which has rigid 

boundaries, equation (A.46) becomes: 

( )

( ) SG core loss

p

out out

p

in in v

q q q

v

e - m h - e + v 

v

e = m h - e + v dt

dP

 P

e M

+ + +

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ ∑

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ ⎟ ∑ ⎠

⎞ ⎜

⎝

⎛

∂

∂

&

&

 (A.47) 

The net energy transfer rate is given as: 

q = q + q + q net SG core loss (A.48) 

Equation (A.47) is the governing equation for depressurization behavior in the RCS. 
A-32 

A.5.3 Top-Down Reactor Coolant System Depressurization Scaling Analysis 

The mass conservation equation, equation (A.37), is expressed in dimensionless form by 

dividing each term by its respective initial condition and further dividing by the mass flow rate of 

the fluid leaving the break. This results in the following dimensionless mass balance equation: 

= m - m dt

dM +

Brk + m DVI

+

τRCS Π Σ & Σ & (A.49) 

where the superscript + indicates normalization with respect to initial conditions. 

The residence time constant, (τrcs), for the depressurization transient is given by: 

m

M = Brk,o

o

rcs Σ & τ (A.50) 

and the characteristic time ratio is given by: 

m

m = Brk,o

INJ,o

m

&

&

Σ

Σ

Π (A.51) 

Πm is the system mass flow rate ratio. For a constant injection flow rate, Πm represents the total 

liquid mass injected into the RCS during the residence time (τrcs). 

Equation (A.47) can be expressed in dimensionless form by dividing each term by its respective 

initial condition. The normalized terms are as follows: 

M = M M+ o (A.52) 

P = P P+ o (A.54) 

q = q q+

net net,o net (A.53) 

m = m m+ Σ & DVI Σ & DVI,oΣ & DVI (A.55) 

m = m m+ Σ & Brk Σ & Brk,oΣ & Brk (A.56) 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

v

e

h - e + v v

e = h - e + v ve

h - e + v 

p

DVI

+

p

DVI

o p

DVI (A.57) 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

v

e

h - e + v v

e = h - e + v ve

h - e + v 

p

Brk

+

p

Brk

o p

Brk (A.58) 
A-33 

Substituting these equations into (A.47) and dividing through by 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

v

e

m h - e + v 

p

Brk

o

& Brk,o yields the dimensionless depressurization rate equation: 

+ q

v

e h - e + v m - 

v

e = m h - e + v t

P

P

e

M

+

net

p

Brk

+ +

Brk

p

DVI

+

+

DVI

h

+

RCS

+

v

+

Π

Π

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

Π

Σ

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ Σ

Π

Π

∂

∂ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

Γ

ε

ε

ε

τ

&

&

 (A.59) 

where the characteristic time ratios are given by: 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ Σ

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ Σ

Π

v

e

mdot h - e + v 

v

e

m h - e + v 

=

p

Brk

o

Brk,o

p

DVI

o

DVI,o

h

&

 (A.60) 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ Σ

ΠΓ

v

e

m h - e + v 

q = 

p

Brk

o

Brk,o

net,o

&

 (A.61) 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

Π = =

v

e

h - e + v 

 

p

Brk

o

,

o

v o

o

P

e

p

ε ε (A.62) 

Πh is the energy flow rate ratio. It represents the ratio of the total energy change due to fluid 

injection to the energy change caused by the break flow. ΠΓ is the power ratio. It represents 

the ratio of the net heat into the system to the rate of fluid energy transport through the break. 

Πε is the fluid mixture dilation property group. 

Equation (A.62) reveals that the fluid dilation property group couples the system intensive 

energy change to the intensive energy at the break. For high pressure systems venting to 

ambient, the fluid properties at the break are determined at critical flow conditions. 
A-34 

Evaluating the RCS time constant, τrcs, and the dimensionless groups Πm, Πh, Πr, requires 

knowledge of the pressure-scaled fluid properties and the critical mass flux. These parameters 

are evaluated in the bottom-up scaling analysis that follows. 

A.5.3.1 Similarity of Pressure Trajectories in Dimensionless Phase Space 

One of the goals of the test program is to operate the APEX facility such that the pressure 

trends that evolve in APEX for a given scenario would be the same for a similar scenario in 

AP1000 when the results are plotted in dimensionless phase space. That is, plotting scenario 

pressure histories as P/Po versus t/τRCS, or P/Po versus M/Mo, would yield overlaying curves for 

the two facilities. This condition can be achieved by satisfying the following requirements: 

The scenarios are initiated from the same initial condition in dimensionless phase space. In this 

case P+

 at t

+

=0, is 1. The rate of change, (i.e., slope), is preserved in dimensionless phase 

space. This imposes the following scaling criterion: 

=1

dt

dP+

+

R

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ (A.63) 

Satisfying the requirement given by equation (A.63) means preserving the dimensionless Π

groups on the right-hand side of equation (A.59). If the two requirements listed above are 

satisfied, the following is true: 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

P

P = PP

o APEX o AP1000

 (A.64) 

This means that the dimensionless pressure at any point along the scenario trajectory will be 

the same in APEX and AP1000. 

Figure A.7 shows how the saturation pressures in APEX would relate to the saturation 

pressures in AP1000. In this figure, Po for AP1000 is 8.27 MPa (1200 psia), while Po for APEX 

is 2.36 MPa (342 psia). These pressures correspond to the turbine bypass pressure relief 

setpoints on the secondary side. 
A-35 

(P/Po)AP1000 = (P/Po)APEX

R2 = 1

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

APEX (P/Po)

AP1000 (P/Po) Figure A.7 Scaling Relationship Between AP1000 and APEX Saturation Pressure 

A.5.4 Bottom-Up Scaling Depressurization Scaling Analysis 

The objective of the bottom-up scaling analysis was to obtain the closure relations needed to 

evaluate the scaling ratios developed through the top-down analysis. These closure relations 

are dependent on fluid properties. Therefore, the first step in the bottom-up analysis was to 

develop the method to relate fluid properties in AP1000 to those at reduced pressure in APEX. 

A.5.4.1 Self-Similarity of Fluid Properties in Phase Equilibria (Psat Scaling) 

The equation of state for many of the important thermodynamic properties, ψ (P), of the 

saturated fluid can be described by a simple power function as follows: 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

ψ ψ

ξ

P

P = o

o (A.65) 

where ψo can be evaluated at an arbitrary reference pressure, Po, within the range of 

applicability of equation (A.65) and ξ is an empirically determined constant. For saturated water, 

having a pressure between atmospheric and approximately 15 MPa (2175 psia), the following 

thermodynamic properties are well-correlated by equation (A.65) as demonstrated in Figures 

A.8 through A.11. 
A-36 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

P

P

h

Pv = h

Pv

o

0.11

fg

fg

o fg

fg (A.66) 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

P

P

s = s o

0.20

f f,o (A.67) 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

P

P

v = v o

-1.02

g g,o (A.68) 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

P

P

v = v o

-1.06

fg fg,o (A.69) 

The form of equation (A.65) has the special property of being “invariant” with respect to a twoparameter transformation in scale. That is, rescaling the variables such that: 

ψ ψ* = C1 (A.70) 

ψ ψ*

o C1 o = (A.71) 

P = C P* 2 (A.72) 

P = C P* o 2 o (A.73) 

where C1 and C2 are constants, and substituting these rescaled variables into equation (A.65) 

yields: 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

ψ ψ

ξ

P

P = *

o

* *

o

* (A.74) 

The form of the equation of state remains unchanged. This type of transformation is known as a 

“stretching” or “similarity” transformation. The special feature of invariance with scale implies 

that the points and slopes of any curves in the (ψ,P) phase space can be related to 

corresponding points and slopes in the (ψ*

, P*) phase space [Ref. 12]. This type of scale 

invariance is known as “self-similarity.” Self-similarity can be defined as a repetition of detail at 

descending scales. If the dimensionless pressure ratios are preserved in the scale model, selfsimilarity dictates that the dimensionless fluid properties (ψ/ψo) will also be preserved. 
A-37 

Pvfg/hfg = 0.0964P0.1144

R2 = 0.999

0.06

0.07

0.08

0.09

0.1

0.11

0.12

0.13

0.14

0 2 4 6 8 10 12 14 16

Pressure (MPa)

Pvfg/hfg

Figure A.8 Power Law for Pvfg/hfg as a Function of Pressure 

sf = 2.1209P0.2011

R2 = 0.9995

0

0.5

1

1.5

2

2.5

3

3.5

4

0 2 4 6 8 10 12 14 16

Pressure (MPa)

Saturated Liquid Entropy (kJ/kg/C)

Figure A.9 Power Law for sf as a Function of Pressure 
A-38 

vg = 0.1925P-1.0246

R2 = 0.9968

0.01

0.1

1

10

0 2 4 6 8 10 12 14 16

Pressure (MPa)

vg (m3/kg)

Figure A.10 Power Law for vg as a Function of Pressure 

vfg = 0.1911P-1.0582

R2 = 0.9939

0.001

0.01

0.1

1

10

0 2 4 6 8 10 12 14 16

Pressure (MPa)

vfg (m3/kg)

Figure A.11 Power Law for vfg as a Function of Pressure 
A-39 

A.5.4.2 Equations of State for Saturated Pressure and Temperature (Tsat Scaling) 

The Clausius-Clapeyron equation is the classical differential equation that defines the slope 

dP/dT, for a phase equilibrium curve. It is derived assuming that the Gibbs free energies for 

the two-phases being considered are equal [Ref. 12]. Using the saturated definition for the 

Gibbs free energy and relating the change in entropy to the latent heat of vaporization and the 

saturation temperature yields the well-known Clausius-Clapeyron equation: 

v T

h = dT

dP

fg

fg (A.75) 

An analytical expression for an equation of state relating saturation pressure to saturation 

temperature can be obtained as follows. Substituting equation (A.66) into (A.75), where ξ is 

0.11 for water, and rearranging yields: 

dT

T

1

v

h dP = P

P

fg

fg

o o

-1

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ ξ

 (A.76) 

Integrating from Po to Psat and from To to Tsat yields the desired relationship between Psat and 

Tsat. 

( ) ⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

+ ξ

ξ

T

T Ln

Pv

h

P T = P 1 o

sat

1/

o fg

fg

sat sat o (A.77) 

Figure A.12 shows that equation (A.77) is quite accurate for the range of conditions examined. 

350

400

450

500

550

600

650

0 2 4 6 8 10 12 14 16

Pressure (MPa)

Temperature (K)

T(K) Data

Equation

Figure A.12 Comparison of Equation (A.77) to Steam Tables 
A-40 

Equation (A.77) can be rearranged to obtain saturation temperature in terms of saturation 

pressure. The result is: 

( ) ⎪⎭

⎪

⎬

⎫

⎪⎩

⎪

⎨

⎧

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

ξ

ξ

 P

P 1- h

Pv 

1

T P = T exp - o

sat

fg

fg

o

sat sat o (A.78) 

This equation can be accurately approximated using the following identity: 

exp (β) ≈1 + β (A.79) 

for β << 1 which is applicable to the current situation. Thus equation (A.78) is approximated by: 

⎪⎭

⎪

⎬

⎫

⎪⎩

⎪

⎨

⎧

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

ξ

ξ

 P

P 1 - h

Pv 

1

T = T 1 - o

sat

fg

fg

o

sat o (A.80) 

Rearranging equation (A.80) yields: 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

ξ ⎥

⎦

⎤ ⎢

⎣

⎡

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ ξ

P

P 1 - 1 = TT 1 - Pv

h

o

sat

o

sat

fg

fg

o

 (A.81) 

This equation is valid for both APEX and AP1000 for the range of saturated conditions being 

examined. Writing this equation for APEX yields: 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⎥ ξ ⎥

⎦

⎤

⎢

⎢

⎣

⎡

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ ξ

P

P 1 - 1 = TT 1 - P v

h

o

sat

o APEX

sat

o fgo APEX

fgo

APEX

 (A.82) 

and for AP1000: 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⎥ ξ ⎥

⎦

⎤

⎢

⎢

⎣

⎡

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ ξ

P

P 1 - 1 = TT 1 - P v

h

o

sat

AP1000 o

sat

fgo AP1000 o

fgo

AP1000

 (A.83) 

Recalling that: 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

P

P = PP

o

sat

o AP1000

sat

APEX

 (A.84) 

it is recognized that the right-hand side of equation (A.82) and (A.83) are identical. Therefore, 

they can be set equal to obtain: 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

T

T 1 - P v

h = T

T 1 - P v

h

o

sat

o fgo AP1000

fgo

AP1000 o

sat

o fgo APEX

fgo

APEX

 (A.85) 
A-41 

Rearranging this equation yields the desired saturation temperature scaling relation: 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ -1

P v

h - T

T

P v

h = T

T

o fgo

fgo

R o

sat

o fgo APEX

fgo

R o

sat

AP1000

 (A.86) 

Equation (A.86) is linear as shown in Figure A.13 for the APEX and AP1000 reference 

temperatures (i.e., the saturation temperatures for the corresponding APEX and AP1000 

reference pressures). Figure A.13 compares the approximate solution given by equation (A.86) 

to the exact result obtained using the steam tables. 

(T/To)AP1000 = 1.1248(T/To)APEX - 0.1258

R2

 = 0.9999

0.7

0.75

0.8

0.85

0.9

0.95

1

0.7 0.75 0.8 0.85 0.9 0.95 1

APEX (T/To)

AP1000 (T/To)

Figure A.13 Scaling Relationship Between AP1000 and APEX Saturation Temperature 

A.5.4.3 Dilation Property Group for Saturated Liquid-Vapor Blowdowns 

This section demonstrates how the dilation property group, Πε, given by equation (A.62), is 

evaluated for different fluid conditions. For a two-phase fluid mixture in phase equilibrium, the 

following thermodynamic relations are applicable for the mixture specific internal energy and the 

mixture specific volume, respectively: 

e = ef + xefg (A.87) 

v = vf + xvfg (A.88) 

Eliminating the fluid mixture equilibrium quality, x, between equations (A.87) and (A.88) yields 

the following expression for specific internal energy: 
A-42 

e

v

v - v

e = e + fg

fg

f

f (A.89) 

The partial derivative of this equation with respect to specific volume, at constant pressure, 

yields: 

v

e = v

e

fg

fg

P

⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ (A.90) 

Lastly, the latent heat of vaporization is given by: 

hfg = Pvfg + efg (A.91) 

Substituting equations (A.87) through (A.91) into equation (A.62) yields the following expression 

for the saturated mixture dilation property group: 

[ ] P

e

h

P v

v + x v

1 = fg v

fg

f Brk fg

⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

Πε (A.92) 

where the equilibrium vapor quality at the break is given by xBrk: 

h

h - h

x = fg

Brk f

Brk (A.93) 

Let us examine the pressure dependencies of the terms that comprise the fluid mixture dilation 

property group given by equation (A.92). The property group (Pvfg/hfg) has already been 

evaluated as shown in Figure A.8 and found to be described by equation (A.66), a power law 

that exhibits self-similarity for a wide range of saturation pressures. 

The last term on the right-hand side of equation (A.92), (Me/MP)v , is a partial derivative that 

requires examining the change in the fluid mixture internal energy with respect to pressure while 

holding the mixture specific volume, v, constant. It should not be confused with the total specific 

internal energy change with respect to pressure. An adequate model for this term can be 

obtained by developing an expression for the mixture internal energy as a function of pressure 

while at constant specific volume and taking the derivative with respect to pressure. Therefore, 

the term (Me/MP)v shall be evaluated assuming that the system expands along a trajectory which 

maintains the specific volume of the mixture at a constant value, vo, which is the initial specific 

volume of the saturated mixture. That is, 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

v

v - v

e = e + e

fg

o f

f fg (A.94) 

For the depressurization transients of interest to this analysis, the saturated mixture blowdown 

begins with the system nearly filled with saturated liquid with the exception of a small vapor 

volume located in the pressure vessel head. Since the density of the liquid phase is much 

greater than that of the vapor phase, the initial system vapor quality is very low and hence the 
A-43 

initial fluid mixture specific volume is essentially the same as the saturated liquid phase specific 

volume. Thus the value of vo, implemented in equation (A.94) is closely approximated by vfo. 

Figure A.14 shows how the mixture internal energy varies with pressure assuming a constant 

specific volume, vo = vfo, during the depressurization process. The plot has been made 

dimensionless to collapse the trends for a variety of initial pressures. 

e/eo = (P/Po)

0.2603

R2 = 0.9999

0

0.2

0.4

0.6

0.8

1

1.2

0 0.2 0.4 0.6 0.8 1 1.2

P/Po

e/eo

Po = 2000 psia

Po = 1700 psia

Po = 1400 psia

Po = 1100 psia

Po = 800 psia

Po = 500 psia

Po = 300 psia

Constant (de/dp)v

Power (Po = 1700 psia)

Figure A.14 Plot of Dimensionless Mixture Specific Energy as a Function of Dimensionless 

Pressure 

Note that for pressure ratios greater than ~0.5, the trend is linear, indicating that the slope, 

(Me/MP)v, is a constant in this region. As shown in the figure, the following power law (R2

 = 

0.9999) accurately describes the mixture-specific internal energy ratio for a wide range of 

normalized pressures. 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

P

P = e

e

o

0.2603

o vfo

 (A.95) 

Taking the derivative of this equation with respect to pressure yields the expression for the 

desired partial derivative: 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

P

P

P

e = 0.2603

P

e

o

0.2603

fo

vfo

 (A.96) 
A-44 

It should be noted that the numerical constants arising in equations (A.95) and (A.96) are 

particular to the case where the initial fluid mixture-specific volume and specific energy are 

given by vfo and efo , respectively. Different constants are obtained at different initial vapor 

qualities for the saturated mixture. 

Substituting equations (A.66) and (A.96) into (A.92) yields the following equation for the dilation 

property group. 

[ ] ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

Πε

P

P h

e v v + x v

0.2603 = o

-0.6257

fg

f fg

o f Brk fg

 (A.97) 

Equation (A.97) can be readily evaluated for two bounding cases, saturated liquid breaks and 

saturated vapor breaks. 

The fluid conditions for breaks located at low points in the RCS can be approximated assuming 

saturated liquid at the break. For these conditions, xBrk would be approximately zero, and 

equation (A.97) would become: 

Saturated Liquid Breaks: 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

Πε

P

P

v h

e v = 0.2603 o

-0.6257

f fg

f fg

o

f (A.98) 

Here it has also been assumed that vf does not change significantly over the range of pressure 

of interest. 

The fluid conditions for breaks located at high points in the RCS (e.g., ADS1-3) can be 

approximated assuming saturated vapor at the break. For these conditions, xBrl would equal 

one and equation (A.97) would become: 

saturated vapor breaks: 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

Πε

P

P

h

e = 0.2603 o

0.4443

fg

f

o

g (A.99) 

To obtain this equation, use has been made of equation (A.68), the power law for the vaporspecific volume. Equations (A.97) and (A.99) are written in terms of scaling ratios, as follows: 

( ) ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

Πε

v h

e v = f fg

f fg

o,R

f R (A.100) 

and 

( ) ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

Πε

h

e = fg

f

o,R

g R (A.101) 

where equation (A.64) has been implemented to eliminate the pressure ratios. 
A-45 

A.5.4.4 Critical Flow Models 

For the blowdown cases of interest to this study, all of the break flows are assumed to be 

choked. This section presents two critical flow models (a saturated vapor model and a 

saturated liquid model) used to estimate the break mass flow rate: 

In general, the break mass flow rate is expressed as: 

m& Brk = CD GcABrk (A.102) 

where CD is the discharge coefficient, Gc is the critical mass flux and ABrk is the break flow area. 

A.5.4.4.1 Saturated Steam 

This initial critical mass flow rate for saturated steam, approximated as a perfect gas, is given by 

Moody [Ref. 13]:

( ) [ ] P + 1

2

m = C A go o

1/ 2 2 -1

+1

Brk,o D Brk γρ ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

γ

γ

γ

& (A.103) 

where CD is the discharge coefficient, ABrk is the break flow area and γ is the ratio of specific 

heats, which is approximately 1.33 for steam. 

The transient, pressure-dependent, critical mass flow rate for the steam would be: 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ γ

γ

P

P

m = m o

2

1+

& Brk & Brk,o (A.104) 

Once again, we find a power law, self-similar in pressure. 

A.5.4.4.2 Saturated Liquid 

The models typically implemented for the case of saturated liquid choke flow are the 

homogeneous equilibrium model or the Henry-Fauske model [Ref. 14]. An earlier study of 

APEX data indicated that the Henry-Fauske model at saturated liquid conditions best fit the 

APEX data [Ref. 15].

The Henry-Fauske critical flow model for saturated liquid is given by the following set of 

equations: 

{ } ( ) ( )( ) ⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

dP

ds

0.14 s

x G = v - v

f

fg Brk Brk

Brk

g Brk f

-1/ 2

c (A.105) 

and 

2P

v G =1 - P

P 2

Brk f c (A.106) 
A-46 

In these equations, the parameters with the subscript “Brk” are evaluated at the break plane. 

All other parameters are evaluated at the system conditions. Thus, PBrk is commonly known as 

the throat pressure, while P is the stagnation pressure. A typical value of equilibrium quality for 

the saturated liquid breaks observed in APEX was xBrk equal to 0.03. 

The saturated liquid entropy derivative on the right-hand side of equation (A.105) was obtained 

by taking the derivative of equation (A.67) with respect to pressure. This yielded the following 

expression for (dsf/dP): 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

P

P

P

s = 0.2

dP

ds

o

0.2

f f,o (A.107) 

Equations (A.105) and (A.106) were solved iteratively to obtain the critical flow rates for 

saturated liquid breaks in APEX and AP1000. Figure A.15 presents the result of this analysis. 

Furthermore, it indicates that the critical flow model also obeys a power law that is self-similar 

in pressure. That is: 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ ≈ P

P

G G o

0.7063

c co (A.108) 

Gc = 14064P0.7063

R2 = 0.9981

0.E+00

1.E+04

2.E+04

3.E+04

4.E+04

5.E+04

6.E+04

7.E+04

8.E+04

9.E+04

1.E+05

0 2 4 6 8 10 12 14

Stagnation Pressure (MPa)

Critical Mass Flux (kg/(m2-s))

Figure A.15 Critical Mass Flux for Saturated Liquid Breaks as Predicted by the Henry Fauske 

Model (xBrk = 0.03) 

Using equation (A.102), the mass flow rate for saturated liquid breaks can be expressed as: 
A-47 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ ≈ P

P

m m o

0.7063

& Brk & Brk,o (A.109) 

Equations (A.104) and (A.109) served as the closure relations for the top-down scaling 

dimensionless groups, Πm, Πh, ΠΓ, and the time constant τRCS. 

A.5.5 RCS depressurization and Inventory Scaling Criteria 

It is desired that APEX-AP1000 depressurization rate and the rate at which the RCS mass 

inventory changes be preserved in dimensionless phase space, (P+

,M+). That is, the following 

similarity criteria shall be imposed: 

=1

dt

dM+

RCS

R

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

τ (A.110) 

=1

dt

d P+

RCS

R

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

τ (A.111) 

where: 

⎥

⎦

⎤ ⎢

⎣

⎡

τ

m

M = Brk

RCS

o,R

RCS,R

& (A.112) 

and the one-half time scale requirement has been set. 

If these criteria are preserved and the same initial conditions, (P+

, M+)o are imposed in APEXAP1000, 

( ) 2

1

RCS = τ R (A.113) 

the same process trajectory will result when plotted in dimensionless phase space. Applying 

these conditions to the dimensionless balance equations given by equations (A.49) and (A.59) 

means that the following scaling criteria must be on the order of unity. 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

Π

m

m = Brk,o

INJ,o

R

m,R

&

& (A.114) 

⎥

⎥

⎥

⎥

⎥

⎦

⎤

⎢

⎢

⎢

⎢

⎢

⎣

⎡

⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ Σ

⎟

⎟

⎠

⎞ ⎜

⎜

⎝

⎛

Π

Π

P

e

m P 

v

e

m h - e + v 

=

v,o

Brk,o o

p

DVI

o

DVI,o

R

h

R &

&

ε

 (A.115) 
A-48 

⎥

⎥

⎥

⎥

⎦

⎤

⎢

⎢

⎢

⎢

⎣

⎡

⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ ⎟

⎟

⎠

⎞ ⎜

⎜

⎝

⎛

Π

ΠΓ

P

e

m P 

q =

v,o

Brk,o o

net,o

R

R & ε

 (A.116) 

( )

⎥

⎥

⎥

⎥

⎥

⎦

⎤

⎢

⎢

⎢

⎢

⎢

⎣

⎡

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

Π

v

e

h - e + v 

P

e

P 

=

p

Brk

o

v,o

o

R

ε R (117) 

The SBLOCA cases considered in the analysis are the case of saturated vapor at the ADS1-3 

location and saturated liquid at the break and ADS4 locations. 

A.5.5.1 Saturated Vapor Break Flow Areas 

For saturated vapor breaks, the following would be applicable: 

hBrk = hg (A.118) 

Thus, implementing equations (A.89) through (A.91) yields: 

v

v h = v

e

h - e + v fg

g fg

p

Brk

o ⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ (A.119) 

Similarly: 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

∆ ⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ - h

v

h v = v

e

h - e + v sub

fg

fg f

o p

DVI

o

 (A.120) 

The injected liquid subcooling is defined as hf - hDVI. Using these fluid property relations, it is 

now possible to evaluate equations (A.113) through (A.117). Substituting the critical flow 

equation for saturated vapor, equation (A.103), and the RCS time constant ratio given by 

equation (A.110) into (A.111) yields. 

[ ] 2

1 =

goPo

0.5

ABrk

MRCS

R⎥

⎥

⎥

⎦

⎤

⎢

⎢

⎢

⎣

⎡

ρ

 (A.121) 

The MRCS ratio shall be set such that: 

192

1

MRCS,R = (A.122) 
A-49 

Substituting this condition into equation (A.120) and solving for the break area ratio yields: 

( ) [ ] ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

ρ P

1

96

1

A =

go o

1/ 2

R

Brk R (A.123) 

A.5.5.2 Saturated Liquid Break Flow Areas 

For saturated liquid breaks, the following would be applicable: 

hBrk = hf (A.124) 

Thus, implementing equations (A.89) through (A.91) results in: 

v

v h = v

e

h - e + v fg

f fg

p

Brk

o ⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ (A.125) 

Substituting the critical flow equation for saturated liquid, equation (A.109), and the RCS time 

constant ratio given by equation (A.110), into (A.111) yields: 

2

1 = C A P

M

0.7063

D Brk o

RCS

R

⎥

⎦

⎤ ⎢

⎣

⎡ (A.126) 

Assuming similar discharge coefficients, substituting equation (A.122) and rearranging, yields 

the break flow area ratio: 

( ) ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

P

1

96

1

A = 0.7063

o R

Brk R (A.127) 

Equations (A.123) and (A.127) become identical for the case of fluid property similitude. That is: 

( ) 96

1

ABrk = R (A.128) 

A.5.5.3 ADS Valve Flow Area Scaling Ratios 

Operation of the ADS is initiated when the CMT liquid level drops below a prescribed setpoint. 

This typically occurs subsequent to PZR draining. Therefore, most of the flow is single-phase 

steam. As a result, the ADS1 and ADS2 valves will be scaled in accordance with equation 

(A.123). That is: 

( ) [ ] ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

ρ P

1

96

1

A =

go o

1/ 2

R

ADS1,2 R (A.129) 
A-50 

The opening of the ADS3 valve shall be used to transition to fluid property similitude conditions. 

This procedure was first implemented in the NRC’s AP600 test program to allow actuation of the 

ADS4 valves at higher operating pressures. The flow area for the APEX ADS3 valve was 

designed such that: 

( ) ( ) A 1000

96

1

AADS-3 APEX = ADS1,2,3 AP Σ (A.130) 

The procedure was such that ADS1 and ADS2 would open sequentially as specified by the 

safety logic. Then ADS1 and ADS2 would be closed and APEX ADS3, representing the sum of 

flow areas for all three stages, would be opened. Thus the ADS1-3 train would be configured 

for fluid property similitude and long-term cooling. 

The ADS4 valve flow area was sized based on fluid property similitude, as follows: 

( ) 96

1

AADS4 = R (A.131) 

A.5.5.4 DVI Mass Flow Rate Scaling Ratio 

All of the injection flow rates will be scaled as follows for the one-half time scale. This is done 

by preserving the line resistance and relative elevation in the IRWST, CMT, and ACCs. Thus: 

( ) 96

1

mDVI,o = R & (A.132) 

A.5.5.5 Net Power Scaling Ratio 

Substituting equation (A.96) into (A.116) and setting the result equal to one yields the net power 

scaling ratio. Use has been made of the fact that (P/Po)R is unity. The core decay power 

scaling has already been set in Section A.3 to be 1:96. 

(q ) = [m e ] net,o R & Brk fo R (A.133) 

A.5.6 Initial Conditions and Flow Areas 

Table A.11 presents the initial conditions for the APEX test facility for a LOCA transient. These 

parameters will be modified for other transients such as the station blackout. 
A-51 

Table A.11 APEX-AP1000 LOCA Initial Conditions 

Parameters AP1000 APEX Units 

Decay Power (2.8%) 

RCS Pressure 

SG Operating Pressure 

CL Temperature 

SG Operating Pressure 

SG PORV Setpoint (Reference Pressure) 

SG Shell Side Liquid Level 

SG Shell Side Liquid Mass 

SG Shell Side Vapor Mass 

PZR Liquid Volume 

ACC Gas Pressure 

Table A.12 presents the flow areas required for the ADS valves as used for LOCA transients, 

Table A.13 presents the ADS1 and 2 flow areas for the special case of a transient initiated by 

the inadvertent opening of the ADS1 valve, Table A.14 presents the scaled break sizes, and 

Table A.15 presents the numerical values for the RCS time constant and P ratios. 

a,b,c 
A-52 

Table A.12 Summary of APEX-AP1000 ADS1-4 Valve Flow Areas for LOCA 

Component 

Scaling 

Ratio AP1000 APEX Units AP1000 APEX

Units

ADS1-2 FLOW AREAS FOR LOCA (SATURATED STEAM) 

ADS1 Single Valve Effective Flow Area 

(min) 

ADS1 Single Valve Effective Throat I.D. 

ADS1 Two Valves Effective Flow Area 

(combined) 

ADS1 Two Valves Effective Throat I.D. 

(combined) 

ADS2 Single Valve Effective Flow Area 

(min) 

ADS2 Single Valve Effective Throat I.D. 

ADS2 Two Valves Effective Flow Area 

(combined) 

ADS2 Two Valves Effective Throat I.D. 

combined 

 ADS1-3 FLOW AREAS COMBINED (FLUID PROPERTY SIMILITUDE) 

ADS1,2,3 Single Train Effective Flow Area 

(min) 

ADS1,2,3 Single Train Effective Throat I.D. 

ADS1,2,3 Two Trains Effective Flow Area 

(combined) 

ADS1,2,3 Two Trains Effective Throat I.D. 

(combined) 

 ADS4 FLOW AREAS PER HOT LEG (FLUID PROPERTY SIMILITUDE) 

ADS4 Single Train (50% Effective Flow 

Area) 

ADS4 Single Train Effective Throat I.D. 

ADS4 Two Trains (100% Effective Flow 

Area) 

ADS4 Two Trains Effective Throat I.D. 

(combined) 

a,b,c 
A-53 

Table A.13 Summary of APEX-AP1000 ADS1-2 Valve Flow Areas for Inadvertent Opening 

 of ADS 

Component 

Scaling

Ratio AP1000 APEX Units AP1000 APEX

Units

 ADS1-2 FLOW AREAS FOR INADVERTENT ADS OPEN IN TRANSIENT (SATURATED LIQUID)

ADS1 Single Valve Effective Flow Area 

(mini) 

ADS1 Single Valve Effective Throat I.D. 

ADS1 Two Valves Effective Flow Area 

(combined) 

ADS1 Two Valves Effective Throat I.D. 

(combined) 

ADS2 Sin le Valve Effective Flow Area (min) 

ADS2 Single Valve Effective Throat I.D. 

ADS2 Two Valves Effective Flow Area 

(combined) 

ADS2 Two Valves Effective Throat I.D. 

(combined) 

A,b,c 
A-54 

Table A.14 APEX-AP1000 Break and Steam Generator PORV Sizes. LOCA FLOW AREAS 

 (Pressure Scaled Two Phase Fluid) 

Component 

Scaling 

Ratio AP1000 APEX Units 

2-INCH COLD LEG BREAK 

Flow area 

Diameter 

DOUBLE-ENDED DVI LINE BREAK 

DVI Nozzle Flow Area 

Diameter 

STEAM GENERATOR PORV FLOW AREA (SINGLE-PHASE STEAM) 

*PORV Equivalent Flow Area 

PORV Equivalent Throat Diameter 

*Based on 1.02 x 106 lbm/hr at 1200 psia for AP1000 

Table A.15 Time Constants and Π Groups for APEX Pressure Scaled LOCAs

Π Group Saturated Vapor Breaks Saturated Liquid Breaks 

(t/τRCS),R 

Πm,R 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

Π

Π

ε

h

R

 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

Π

Π

ε

Γ

R

 

( ) Πε R 

A.5.7 Summary and Conclusions 

This section presents the results of the SBLOCA scaling analysis. The goal was to scale the 

APEX initial conditions and break flow areas such that the RCS depressurization rate and 

corresponding cooldown rate would be preserved in dimensionless phase space. The two types 

of break conditions considered were (1) all-saturated vapor at the break, such as the stuck-open 

ADS1-3, and (2) an all-saturated liquid break, which was the assumed condition for the HL break. 

a,b,c 

a,b,c 
A-55 

The scaling analysis provided the RCS initial conditions and the break flow areas. All of the Π

groups, as shown in Table A.14, were found to have values on the order of 1, with the exception 

of the fluid property group under pressure-scaled conditions. 

An examination of the Π groups ratios indicates that for saturated vapor breaks (i.e., stuck-open 

ADS1-3) the depressurization and cooldown rate will be reasonably simulated in APEX. Some 

distortion may occur in simulating the HL break because conditions at the break will likely be 

two-phase rather than saturated liquid as assumed in the analysis. The validity of this 

assumption will depend on whether the injection flows keep the break conditions near the 

saturated liquid state. 

For the conditions of fluid property similitude, all of the Π group ratios listed in Table A.14 will be 

one. Thus, the full-pressure tests that will be performed in APEX to assess the transition from 

ADS4 operation to IRWST injection are expected to yield depressurization behavior that is 

similar to the AP1000. The next chapter presents the scaling analysis for the full-pressure 

ADS4 blowdown phenomena. 

A.6 Scaling Analysis of the Full-Pressure ADS4 Blowdown to IRWST 

Injection Transition Period 

The APEX-AP1000 test facility was modified to conduct full-pressure studies of the thermalhydraulic phenomena that arise during the transition from ADS4 blowdown to the onset of 

IRWST injection. This transition period was previously studied under full-pressure conditions in 

the SPES and ROSA-AP600 test facilities. This period was deemed to be very important by 

analysts because there is a limited amount of liquid available for core cooling while the system 

depressurizes to the IRWST injection setpoint. Operating the APEX facility under reducedpressure conditions during this phase of a transient results in distorting the conditions at which 

the IRWST begins to drain. 

In the AP1000, IRWST injection would initiate when the primary system pressure drops below 

~0.198 MPa (28.7 psia), based on the minimum initial IRWST liquid level. At this pressure, the 

flow at the ADS4 valves would be still be choked (i.e., essentially single-phase vapor flow) and 

independent of the ADS4 line resistance. However, for the one-quarter height IRWST in APEX, 

the onset of flow would not begin until the primary system pressure dropped below ~0.125 MPa 

(18.2 psia). At this pressure, the ADS4 flow would be unchoked and, therefore, the ADS4 flow 

rate would be dependent on ADS4 line resistance. 

In an effort to eliminate the ADS4 flow distortions in APEX during this transition period, a series 

of full-pressure ADS4 blowdown tests were performed. This requireed configuring APEX to 

provide pressure control for ADS4 blowdowns initiating from steady-state conditions, pressuring 

the IRWST to simulate full-height conditions and adjusting the IRWST line resistance to achieve 

properly scaled injection flow rates under full-height IRWST conditions. This section presents 

the scaling analysis performed to design the ADS4 blowdown to IRWST injection transition tests. 
A-56 

A.6.1 Description of the ADS4 Blowdown to IRWST Injection Thermal-hydraulic 

Phenomena 

Figure A.16 presents an overview of the thermal-hydraulic phenomena that arise during a 

typical ADS4 blowdown. This figure characterizes the phenomena in terms of RPV mixture 

level (or liquid mass) versus the system pressure. The ADS4 valves open when the liquid 

volume in either of the two CMTs drops below 20% of its initial volume. At PADS4, the pressure 

at which the ADS4 valves open, two-phase fluid choke flow will exist at the ADS4 valves. All of 

the liquid located above the top of the HLs, (i.e., the entrance to the ADS4 line), will eventually 

be swept out of the ADS4 valves. The amount of liquid located above the HLs is dependent on 

the amount of level swell in the core. When the fluid mixture level drops below the top of the 

HLs, the flow out of the ADS4 valves can be best described as a high-quality vapor choke flow. 

Liquid boil-off in the core, liquid flashing to vapor, and liquid carryover from the upper plenum 

will act to reduce the liquid inventory in the RPV. Because there are no other sources of liquid 

during this phase of the transient, primary system pressure must drop below the IRWST liquid 

level head pressure, PIRWST, before the RPV mixture level reaches the top of the core. When 

the primary system pressure drops below PIRWST, less some small check valve cracking 

pressure, IRWST liquid flow will act to reflood the RPV. 

PADS-4

Sweepout of

Liquid Above

Hot Leg

Liquid Boil-off,

Flashing, and

Carryover

RPV MIXTURE LEVEL

PRESSURE

PIRWST

Top of Hot Leg

Top of Core

Two-Phase Fluid

ADS-4 Flow

High Quality Vapor

ADS-4 Flow

IRWST

Injection

Figure A.16 Description of the Transition from ADS4 Blowdown 

to the Onset of IRWST Injection 

Figure A.17 shows how the thermal-hydraulic phenomena described in this section have been 

addressed in the top-down and bottom-up scaling analyses. First a top-down scaling analysis 

was performed to obtain a single governing equation in terms of dP+

/dM+. Next, a bottom-up 

scaling analysis was performed to describe the local thermal-hydraulic processes, including 

critical flow, core fluid boil-off, and core mixture level swell. 
A-57 

A.6.2 Top-Down RCS Depressurization and Mass Inventory Scaling Analysis 

(Pressure Similitude) 

The mass balance equation and depressurization rate equation developed in Section A.5 are 

directly applicable to this analysis. However, the initial conditions and processes to be 

examined are specific to the transition period from ADS4 blowdown to IRWST injection under 

full-pressure conditions. The objective of this analysis is to develop the criteria for preserving 

the AP1000 depressurization and mass inventory behavior for the transition period as described 

in dimensionless phase space (M+

, P+). 

The following mass balance equation was developed for the RCS in Section A.5.2. 

= m - m dt

dM ∑ & in ∑ & out (A.134) 

During the ADS4/IRWST transition period, there will be no sources of liquid injection; therefore, 

the first term on the right-hand side of equation (A.134) can be deleted. Furthermore, the only 

vent paths to be considered in this analysis shall be the ADS4 valves. 

ADS-4 Blowdown to IRWST

Injection Transition

Phenomena

Top-Down/System

Scaling

y RCS Depressurization

y Mass Inventory

Bottom-Up/Process

Scaling

y Core Power

y ADS-4 Flow Rate

y IRWST Injection Setpoint

y IRWST Injection Flow Rate

y Core Fluid Mixture Level

ADS-4 Blowdown

Similarity Criteria

Evaluate Scaling

Distortions

Specify RCS and IRWST

Initial Conditions, ADS-4

Valve Sizes, IRWST Flow

Rates

Level: Primary Loop Process Level

Figure A.17 Scaling Analysis Flow Chart for the Full Pressure ADS4 Blowdown Experiments 

Applying these assumptions results in the following equation: 
A-58 

= - m dt

dM & ADS4 (A.135) 

The following depressurization rate equation was also developed for the RCS in Section A.5.2. 

( )

( ) q q q v

e

m h - e + v

 v

e

m h - e + v = dt

dP

 P

e M

SG core loss

p

out out

p

in in v

+ + +

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ − ∑

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ ⎟ ∑ ⎠

⎞ ⎜

⎝

⎛

∂

∂

&

& (A.136) 

During the ADS4/IRWST transition period, the SGs will be isolated. Therefore, the qSG term can 

be deleted. The heat loss will be negligible compared to the core decay heat. Therefore, the 

qloss term can also be neglected. Applying these assumptions and those for the mass balance 

equation yields: 

 + q v

e - m h - e + v = dt

dP

 P

e M core

p

ADS4 ADS4 v ⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ & (A.137) 

It is noted that the key behavior of interest to the transition period can be described in terms of 

system pressure versus RCS mass inventory (or RPV mixture level) as shown in Figure A.20. 

The process could also be described in terms of two time scales. The first is the time required 

for the RCS to depressurize from PADS4 to PIRWST. The second is the time required to deplete the 

RCS mass inventory from its initial condition, MRCS,o, to its final mass inventory state which may 

or may not be sufficient to keep the core covered. However, rather than working with time 

scales, a departure from previous analyses, the present analysis was simplified by developing a 

single governing equation in terms of RCS mass and pressure. This was done by dividing 

equation (A.137) by the mass balance equation (A.135). The result is: 

m

q - v

e h - e + v = dM

dP Pe M

ADS4

core

p

ADS4 v ⎥ & ⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ (A.138) 

This equation can be expressed in non-dimensional form by dividing each term by its respective 

initial conditions. The normalized terms are as follows: 

M = M M+ RCS,o (A.139) 

P = P P+ RCS,o (A.140) 

q = q q+

core core,o core (A.141) 

m = m m+ & ADS4 & ADS4,o & ADS4 (A.142) 
A-59 

⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

P

e

P

e = Pe +

v v,o v

 (A.143) 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

v

e

h - e + v v

e = h - e + v ve

h - e + v 

p

ADS4

+

p

ADS4

o p

ADS4 (A.144) 

Substituting these dimensionless initial conditions into (A.138) and dividing through by 

⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

P

e

P v,o

o yields the non-dimensional governing equation: 

o ADS4

core

p

ADS4

o v m

q

v

e

h - e + v 1 = dM

dP

P

e

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

∏

∏−

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

∏ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

ε

Γ

ε +

+ +

& (A.145) 

where the dimensionless property group Πε is given as in the previous chapter as: 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

Π

v

e

h - e + v 

P

e

P 

=

p

ADS4

o

v,o

o

ε (A.146) 

and the power-to-energy flow rate ratio is: 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂

ΠΓ

v

e

m h - e + v 

q = 

p

ADS4

o

ADS4,o

core,o

&

 (A.147) 

Hence, to simulate the transition behavior in APEX-AP1000, the dimensionless groups defined 

in equations (A.146) and (A.147) must be preserved. For the full-pressure conditions proposed 

for the ADS4 blowdown test series, fluid property similitude reduces the ratio of these groups to: 

Πε,R =1 (A.148) 

and 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

ΠΓ

m

q = ADS4,o

core,o

R

,R

& (A.149) 

The property group, Πε in AP1000 and APEX-AP1000 are identical and, therefore, the similarity 

criterion is automatically satisfied. 
A-60 

A.6.3 Bottom-Up Scaling Analysis 

This section describes the scaling analyses performed for the ADS4 blowdown/IRWST 

transition phenomena. The core power and ADS4 choked flow rate behavior have already been 

assessed in Section A.5. The core power has been scaled as [ ]a,b,c. Similarly, when fluid 

property similitude exists, the choked flow conditions at the ADS4 valves are preserved, and the 

mass flow rate is scaled at [ ]a,b,c because the ADS4 flow areas have been scaled by 

[ ]a,b,c. Thus, equation (A.149) becomes: 

ΠΓ,R = 1 (A.150) 

which indicates that this dimensionless group will also be preserved in APEX-AP1000. 

A.6.3.1 Flashing and Liquid Boil-off Rate 

The flashing phenomena is automatically preserved by establishing the initial pressure and 

subcooling conditions in APEX as those in the AP1000 for the ADS4 blowdown phase. 

The liquid mass boil-off rate in the core, under steady-state conditions is described as follows: 

h + h

q

m = fg subcooled

core & Boil-off (A.151) 

For fluid property similitude, the ratio of this equation (i.e., APEX to AP1000) becomes: 

96

1

m = q = Boil-off,R core,R & (A.152) 

which is the required value to preserve the one-half time scale requirement. However, it is 

noted that the volume of the APEX downcomer is large on a scaled basis. Of particular interest 

to the full-pressure ADS4 blowdown tests is the downcomer volume between the bottom of the 

CL and the top of the heated fuel elevation. This region represents a source of liquid that 

resides in the downcomer at an elevation above the top of the heated fuel. For APEX, this 

region has a length of [ ]a,b,c and a corresponding volume of [ ]a,b,c. 

For AP1000, this region has a length of [ ]a,b,c and a volume of [ ]a,b,c. 

The length scale ratio is 1:4, as desired. However, the volume scale ratio is 1:100, which is 

approximately 2 times the desired value of 1:192. 

At a system pressure of [ ]a,b,c, the amount of saturated liquid mass that could be 

contained in this region for AP1000 would be [ ]a,b,c. For the similar region in 

APEX, this corresponds to [ ]a,b,c. 

During the transition period of the ADS4 blowdown, the primary source of water will be the 20% 

liquid volume remaining in the CMT. For the fluid conditions stated earlier, this corresponds to a 

volume of [ ]a,b,c in AP1000 and [ ]a,b,c in APEX-AP1000. It is 

desired that a 1:2 boil-off time scale be simulated. Therefore, the following scaling criterion 

must be preserved: 
A-61 

= 0.5

q

(V + V )(h + h ) = 

core

f CMT DC fg subcooled

R

Boil-off,R ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ ρ

τ (A.153) 

For the case of fluid property similitude and a fixed core power ratio, this criterion becomes: 

192

1 (V + V ) = CMT DC R (A.154) 

Because the downcomer volume in APEX is oversized, this criterion can be satisfied by initiating 

the ADS4 blowdown in APEX with a reduced liquid inventory in the CMT starting the ADS4 

blowdown with a 12.5% volume rather than a 20% volume would satisfy this requirement. 

A.6.3.2 IRWST Full-Height Liquid Head and Injection Flow Rate 

Simulating the full-pressure ADS4 blowdown transition in APEX will require a full-height liquid 

head in the IRWST. This will be achieved by pressurizing the APEX IRWST to a pressure of 

[ ]a,b,c to simulate full-height AP1000 conditions. Similarly, the IRWST line resistance 

shall be increased such that the initial IRWST mass flow rate at the simulated full-height 

conditions is [ ]a,b,c. Flow tests shall be performed to fine-tune the IRWST line resistance. 

A.6.3.3 Core Fluid Mixture Level Swell 

The initial liquid mass distribution in the RCS will determine the conditions at which the HL uncovers. 

As shown in Figure A.20, when the HL uncovers, the ADS4 flow transitions from two-phase 

choked flow to essentially single-phase vapor choked flow. This choked flow transition is 

important because the RCS depressurization rate will depend on the type of choked flow 

conditions at the ADS4 valves. 

Properly modeling the initial RCS liquid mass distribution requires preserving the core fluid 

mixture level swell. The equation for level swell is given by: 

1- < >

H

H = c

collapsed

mix

α

 (A.155) 

where Hmix is the fluid mixture level, Hcollapsed , is the collapsed liquid level and <αc> is the 

average core void fraction. To preserve the level swell in the core, the following similarity 

criterion must be preserved: 

= (1- < > )

H

H c R

mix

collapsed

R

⎟

⎟ α ⎠

⎞

⎜

⎜

⎝

⎛ (A.156) 

which implies that the average core void fraction ratio must also be preserved. That is: 

(< > ) =1 αc R (A.157) 
A-62 

The void fraction as a function of position in the core was obtained using the drift-flux 

relationship [Ref. 15]: 

( ) ( )

C j( ) z + V

j z

z = o gj

g α (A.158) 

At stagnant loop conditions, jR is zero, so equation (A.158) becomes: 

( ) ( )⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

α

j z

v

z = C + 

g

gj

o

-1

 (A.159) 

The volumetric vapor flux, jg, was found using the following relation: 

( ) q ( ) z dz

a h

1 j z =

Z

L g c fg

g

NB

′ ρ ∫ (A.160) 

where LNB is the non-boiling height. The core averaged void fraction, <αc>, was obtained by 

integration as follows: 

( ) z dz

L

1 =

c

NB

L

Lc

α α ε ∫ (A.161) 

APEX-AP1000 will implement a constant axial power profile. Therefore, equation (A.158) is 

given by: 

( ) ( )

a h L

q z - L j z =

g c fg c

core NB

g ρ

 (A.162) 

Substituting equations (A.157) and (A.160) into (A.159) yields the core averaged void fraction 

for constant axial power: 

( )

( )

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ ∏ − +

∏ − αε −

c

o z c NB

o z c NB

2

o L

C L L

Ln 1

C L L

L 1

C

1 = (A.163) 

where Πzuber is the Zuber number, rearranged as: 

 v a h

q =

g gj c fg

core zuber

ρ

Π (A.164) 

The AP1000 is assumed to have a chopped cosine shape axial power profile defined as follows. 
A-63 

( ) dz

L

z

q z = q’ cos c

c

z

-Lc / 2

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ Π

∫ (A.165) 

Integrating equation (A.163) yields: 

( ) ⎥

⎦

⎤ ⎢

⎣

⎡

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ π

π L

z 1+ sin L

q z = q’ c

c c (A.166) 

Substituting equation (A.164) into equations (A.158) yields: 

( ) ⎥

⎦

⎤ ⎢

⎣

⎡

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ π

πρ L

z 1+ sin

a h

q’ L j z =

g c fg c

c c

g (A.167) 

Equation (A.165) was substituted into (A.159) and numerically integrated to obtain the core 

averaged void fraction for the AP1000. The churn-turbulent drift-flux correlation was used to 

assess equation (A.157) for both APEX and AP1000. That is: 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ∆

ρ

σ ρ

2

liter

1/ 4

gj

g

v =1.53 (A.168) 

Table A.16 presents the data used to calculate the core average void fraction for the case of an 

ADS4 blowdown from [ ]a,b,c and a decay heat of [ ]a,b,c. The core averaged 

void fraction was found to be [ ]a,b,c for APEX and 60% for AP1000. 

Figure A.18 compares the axial void fraction profiles in AP1000 and APEX-AP1000 for an 

average subchannel for this ADS4 blowdown case. The difference in void fraction profiles 

arises because the power profiles have different shapes. The core exit void fractions were 

calculated to be [ ]a,b,c and [ ]a,b,c for AP1000 and APEX respectively. 
A-64 

Table A.16 Comparison of APEX and AP1000 Core Averaged Void Fractions 

Parameter AP1000 APEX-AP1000 Units 

Core Flow Area (ac) 

Core Heated Length 

Power (2.1% Decay) 

Average Linear Power at Decay Heat 

Pressure 

ρg 

Ρl

 

σ 

Hg 

HL

Co 

vgi 

<αc> 

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

Normalized Axial Position

Void Fraction

AP1000

APEX-1000

Figure A.18 Comparison of AP1000 and APEX Axial Void Fraction Profiles 

for an Average Subchannel 

a,b,c 
A-65 

A.6.3.4 Initial Conditions for the Full-Pressure ADS4 Blowdown Tests 

The following table presents the initial conditions for a full-pressure ADS4 test in APEX-AP1000. 

The test shall be initiated from steady-state conditions. The decay power and system pressure 

were selected based on the results of Chapter 15 of the AP1000 Safety Analysis Report. The 

highest decay power observed while actuating ADS4 was ~2.0% for the DEDVI case. The 

typical pressure at which ADS4 actuates is [ ]a,b,c. 

Table A.17 Proposed Initial Conditions for a Full Pressure ADS4 Blowdown Test In APEXAP1000 

Parameter AP1000 APEX-AP1000 Units 

RCS Pressure 

2% Core Decay Power 

Initial CMT Volume 

ADS4 Flow Areas – 100% 

ADS4 Flow Areas – 50% 

IRWST Nominal Liquid Level 

IRWST Head Pressure 

a,b,c 
A-66 

A.7 Upper Plenum Entrainment Scaling

Top-down component-level and bottom-up process-level scaling analyses have been performed 

to establish the upper plenum geometry and flow conditions that should be modeled in the 

APEX-AP1000 integral system test facility. Figure A.19 describes the scaling analysis process. 

Upper Plenum Pool

Entrainment and DeEntrainment Phenomena

Top-Down/System

Scaling

y Upper Plenum Mass

Depletion Rate

Bottom-Up/Process

Scaling

y Pool Entrainment (Near

Surface, Momentum

Controlled and Deposition

Regions)

y Liquid De-Entrainment on

Upper Plenum Structures

y CCFL (Upper Head and

Upper Core Plate Draining)

Upper Plenum

Similarity Criteria

Evaluate Scaling

Distortions

Specify Flow Conditions,

Upper Plenum, Upper

Head Plate, and Upper

Core Support Plate Flow

Areas and Hole Sizes,

Guide Tube Number and

Diameter

Figure A.19 Upper Plenum Scaling Analysis Flow Chart 

The main objective of the analysis was to obtain the scaling ratios that would preserve, on a 

scaled basis, the rate at which the liquid level in the upper plenum decreases during pool boiling. 

The top-down scaling analysis resulted in a non-dimensional mass balance equation for the 

upper plenum control volume. The bottom-up scaling analysis provided the closure relationships 

for the top-down analysis. In particular, pool boiling entrainment was examined for the nearsurface, momentum-dominated and deposition regions of the upper plenum. 
A-67 

A.7.1 Top-Down Scaling Analysis 

The control volume shown in Figure A.20 was defined for the upper plenum for the purpose of 

writing the mass balance equation. 

m& Carryover

m&INJ

HOT LEG HOT LEG

m& ENT

m&UPD

m& GT

CORE

H

+z

Figure A.20 Upper Plenum Control Volume Located Below the Hot Legs 

The liquid mass conservation equation for the upper plenum is written as follows: 

= m - m dt

dM

IN OUT

UP & & (A.169) 

The upper plenum flow cross-sectional area, aUP, is constant inside the control volume and the 

liquid is assumed to be incompressible; therefore, liquid density rL is uniform and constant. The 

vertical upward direction is defined as the z-coordinate. Thus, removing the constants from the 

derivative on the left-hand side of equation (A.169) yields: 

= m - m dt

dz

aUP ρL & IN & OUT (A.170) 

One source of liquid to the control volume is liquid mass flow from the core region due to safety 

injection. During the later part of ADS4 operation, IRWST injection would be supplying liquid to 
A-68 

the core. A second source of liquid to the upper plenum is de-entrainment on the vertical guide 

tubes. A third source of liquid mass into the control volume is draining from the upper head 

through the upper support plate. 

The primary path of liquid mass loss from the upper plenum is through the HL. When the liquid 

level is below the bottom of the HL, pool entrainment results, causing liquid droplets to be 

carried with the steam out the HL to the ADS4 line. Another source of mass loss is attributable 

to draining from the upper plenum through the upper core plate back into the core. Loss of 

upper plenum liquid mass as a result of flashing or boiling as a result of stored energy release 

from upper plenum structures are not considered applicable during this phase of the transient. 

Equations for the liquid mass loss and injection terms are presented in the bottom-up scaling 

analysis. Expanding equation (A.170) to include the individual liquid mass source and sink 

terms yields: 

= m + m - m - m dt

dz

aUP ρL & INJ & GT & ENT & UPD (A.171) 

The first term on the right-hand side represents the sum of the DVI injection flows that tend to fill 

the upper plenum via the downcomer (i.e., CMT, ACC, IRWST, and sump flows). It is the mass 

flow rate of liquid entering the control volume through the core. The second term on the righthand side represents the guide tube de-entrainment rate. It is the mass flow rate of liquid that is 

de-entrained by the upper plenum structure and that eventually drains as a liquid film into the 

control volume. The third term on the right-hand side is the liquid entrainment rate. It is the 

mass flow rate of liquid produced by the vapor interaction with the liquid (i.e., near surface, 

momentum-controlled or deposition pool entrainment). This liquid is carried out of the control 

volume, where a portion is de-entrained by the guide tubes and the remainder is carried out to 

the HL. The last term on the right-hand side represents mass flow rate of liquid that drains out 

of the control volume by gravity through the upper core plate back into the core region. 

A.7.1.1 Dimensionless Liquid Mass Balance 

Equation (A.171) can be made dimensionless using the following initial and boundary 

conditions: 

H

z

z = +

INJ (A.172) 

where H is the vertical length between the top of the core and the bottom of the HL. 

The dimensionless injection mass flow rate is given as: 

m

m

m = INJ,O

+ INJ

INJ

&

& & (A.173) 

The dimensionless guide tube de-entrainment rate is given as: 

m

m

m = GT,O

+ GT

GT

&

& & (A.174) 

The dimensionless liquid entrainment rate is given as: 
A-69 

m

m

m = ENT,O

+ ENT

ENT

&

& & (A.175) 

The dimensionless upper plenum draining rate is given as: 

m

m

m = UPD,O

+ UPD

UPD

&

& & (A.176) 

Substituting these ratios into equation (A.171) and dividing both sides of the equation by the 

initial entrainment mass flow rate yields the dimensionless mass balance equation for the upper 

plenum liquid. 

= m + m - m - m dt

d z + UPD UPD +

ENT + GT GT + INJ INJ

+

τENT Π & Π & & Π & (A.177) 

In this equation, the liquid entrainment reference time scale has been defined as: 

m

a H = ENT,O

UP L

ENT

&

ρ

τ (A.178) 

Thus the dimensionless time coordinate for upper plenum emptying due to liquid entrainment is 

given by: 

τENT

+ t

t = (A.179) 

All of the remaining filling and draining processes are therefore, expressed in terms of the liquid 

entrainment process. That is, the dimensionless P groups of equation (A.177) are written as: 

m

m = ENT,O

INJ,O

INJ

&

& Π (A.180) 

m

m = ENT,O

GT,O

GT

&

& Π (A.181) 

m

m= ENT,O

UPD,O

UPD

&

& Π (A.182) 

Equations (A.178) through (A.182) will be evaluated in the bottom-up scaling analysis. 

A.7.1.2 Preserving the Upper Plenum Draining Rate During Integral System Testing 

The desired outcome of this scaling analysis is to establish the geometry and flow conditions in 

APEX that would preserve the upper plenum draining rate when plotted in dimensionless phase 

space. That is, plotting the upper plenum liquid level history in terms of the dimensionless 
A-70 

spatial coordinate, z+, and the dimensionless time coordinate, t

+

 , would yield overlaying curves 

for the model and the prototype. This condition can be achieved by satisfying the following 

requirements. 

The scenarios are initiated from the same initial condition in dimensionless phase space. In this 

case, z+

 at t

+

 = 0 is 1. 

The rate of change (i.e., slope) is preserved in dimensionless phase space. This imposes the 

following scaling criterion: 

=1

dt

d z

+

+

R

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ (A.183) 

Note that (z+

)R and (t+)R are unity. Satisfying the requirement given by equation (A.183) can 

only be accomplished by preserving the ratios of the individual P groups on the right-hand side 

of equation (A.177). The APEX facility transient response was originally designed to operate at 

one-half time scale (tR = 0.5) which means that the reference time scale ratio, τR, must also be 

one-half to keep (t+

)R at unity. Because the APEX facility height is fixed at one-quarter scale, 

and the flow area is roughly fixed at [ ]a,b,c , certain physical limitations have already been 

established. Nonetheless, minor modifications to the upper plenum structures are possible. 

To preserve the one-half scale timing of integral system transient behavior, it is desired that: 

2

1 = dtdz

R

⎟

⎠

⎞ ⎜

⎝

⎛ (A.184) 

Substituting equations (A.172) and (A.179) into (A.184) yields: 

2

1 = dt

H d z

+

+

ENT R R

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

τ

 (A.185) 

Imposing the requirement given by equation (A.183) yields: 

2

1 = H

ENT R

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

τ

 (A.186) 

This is a direct result of the one-half time scale requirement. Substituting equation (A.178), 

assuming fluid property similitude, and rearranging yields the following simple scaling criterion 

for the upper plenum: 

2

1 = a

m

UP

ENT,O

R

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ & (A.187) 

For the case where liquid entrainment is the dominant mechanism (i.e., the largest mass 

transfer term on the right-hand side of the balance equation), equation (A.187) coupled with a 

correlation for liquid entrainment represents the governing scaling criterion. The next section 

presents the bottom-up scaling analysis and includes liquid entrainment equations that can be 

used to provide closure to equation (A.187). 
A-71 

A.7.2 Bottom-up Scaling Analysis 

This section examines the following upper plenum filling and draining mechanisms: 

• pool entrainment 

• guide tube de-entrainment 

• the effects of counter-current flow limitations (CCFL) on upper plenum draining 

The phenomenon of interest to this study is upper plenum pool entrainment. The following 

section presents a review of the phenomenon and the existing predictive models. 

A.7.2.1 Pool Entrainment 

The basic equations of pool entrainment will be derived to provide a theoretical context for their 

application to the scaling analysis. Here, the discussion will consist of a brief derivation of Ishii’s 

pool entrainment equations. Following this derivation, certain questions regarding the 

application of these equations in scaling the AP1000 experiments will be addressed. 

In particular, it will be necessary to examine the vertical extent of various entrainment regions, 

the flow regime in the pool, the question of flooding, the phenomena of re-entrainment from a 

falling film, and the issue of enhanced deposition as a result of the lateral motion of the flow. 

In Ishii’s pool entrainment model, the entrainment is defined as the ratio of the droplet to vapor 

mass flux, and represented phenomenologically by the following equation [Ref. 16]:

 f (D, j ) g(v ,D, j )dv dD

j

(j ) = j

j E (h, j ) = i g i D g v

D

0

g g

g

g g

f f

fg g h

max ∫ ∫

ρ

ε

ρ

ρ ∞ & (A.188) 

Here ε-dot is the mass flux of droplets at the liquid surface which is a function of the vapor 

superficial velocity, jg. As the droplets generated at the surface vary in size from zero to some 

maximum diameter, a distribution function for the droplet diameters is employed. This function, 

fD , gives the fraction of droplets within any given range, dD , and is itself a function of jg. 

Obviously, since all the droplets must fall within the range, 0 #D #Dmax: 

 f (D, j ) dD =1 D g

D

0

max ∫ (A.189) 

Furthermore, the droplets so generated will have an initial velocity, vi , which depends upon both 

the superficial velocity, jg , and the droplet diameter, D. Hence, a distribution function, g , is 

needed for the initial velocity, such that: 

g(v ,D, j )dvi =1 ∫0 i g

∞ (A.190) 

As written, equation (A.189) depends upon the maximum droplet diameter to define an upper 

limit of integration. By using the pool entrainment data to determine various experimental 

constants, Ishii found the best agreement for: 

D = 7.24 j

*

g

-1 *

max (A.191) 

In this equation, the asterisk denotes a dimensionless variable, such that: 
A-72 

∆ρ

σ

g

D

D =* (A.192) 

and 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

ρ

σ ∆ρ

2

g

1/ 4

* g

g

g

j j = (A.193) 

With regard to the second integral in equation (A.188), the lower limit, vh , is the minimum initial 

velocity required for a droplet of diameter, D , to reach height, h. With regard to its dependence 

on the droplet diameter, vh = 0 , for droplets in the range 0 # D # Dc, where the critical diameter, 

Dc , is the diameter of the largest droplet which can be suspended by vapor drag. Thus, the 

droplets within this range can reach height h by virtue of drag alone and hence need no initial 

velocity. On the other hand, droplets in the range Dc < D #Dmax are too large to be suspended 

by vapor drag and hence fall back into the pool after losing their initial momentum. 

Consequently, these droplets must possess a minimal initial velocity, vh , to reach the height in 

question. With regard to these calculations, the critical droplet diameter, Dc , is first obtained by 

balancing the weight of the droplet with the vapor drag: 

j 2

1

C 4

D = 6D

g

2

D g g

2

c 3c ∆ρπ π ρ (A.194) 

where 

Re

10.76

C = D 1/ 2 (A.195) 

Solving equation (A.194) for the critical diameter subject to the drag relation given by equation 

(A.195) yields: 

D = 4 j (N ) * 1/ 3

g

*

c μg (A.196) 

where the viscosity number is given by: 

( /g )

N = 1/ 2

g

g

g

ρ σ σ ∆ρ

μ

μ (A.197) 

As mentioned above, vh = 0 for diameters within the range 0 # D # Dc. For diameters within the 

range Dc < D # Dmax, vh is determined by neglecting drag compared to gravity and calculating 

the velocity needed to reach height h. 

v = 2gh ;for Dc < D Dmax

f

h ≤

∆

ρ

ρ (A.198) 
A-73 

Thus, in terms of dimensionless parameters: 

v = 0 ; for 0 D < D *

c * * h ≤ (A.199) 

and 

v = 2 h ;for D < D D*

max * * c

f

g

1/ 2

* * h ⎥ ≤

⎦

⎤ ⎢

⎣

⎡

ρ

ρ (A.200) 

Having discussed the basic equation and the limits of integration, it remains to determine 

expressions for the surface mass flux, ε dot, the diameter distribution function, fD, and the 

velocity distribution function, g. To this end, observe first of all that near the water surface, vh

goes to zero for all droplet diameters in accordance with equations (A.199) and (A.200). Hence, 

by virtue of equations (A.188), (A.189), and (A.190): 

j

(j ) f (D, j ) g(v ,D, j ) dv dD = j(j )

E (0, j ) =

g g

g

D g 0 i g i

D

0

g g

g

fg g

*

max

ρ

ε

∫ ∫

ρ

ε ∞ & & (A.201) 

Using the limited data available for near surface entrainment, Ishii determined that [Ref. 16]:

⎥

⎦

⎤ ⎢

⎣

⎡

∆ρ

ρ

ρ

ε g

-1.0

-3

g g

g

fg g = 4.84x10 j

(j )

E (0, j ) = & (A.202) 

Thus: 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

∆ρ

ρ

ε ρ g

-1.0

g g

-3

g &(j ) = 4.84 x10 j (A.203) 

To determine the distribution function, fD, observe that after the droplets in the supercritical 

range have fallen back into the pool, one is left with droplets in the subcritical range. Since 

these droplets are sufficiently small to be suspended by vapor drag, they do not fall back into 

the pool. Consequently, the entrainment above a certain height remains constant except for the 

phenomenon of deposition on solid surfaces. Moreover, since these droplets are in the range 

0 # D # Dc, vh = 0 by virtue of equation (A.199), and equation (A.188) can be integrated for a 

subrange of droplet sizes subject to equation (A.190) to yield: 

 f (D, j ) dD

j

(j ) f (D, j ) g(v ,D, j ) dv dD = j(j )

D g

D

0

g g

g

D g 0 i g i

D

0

g g

g

∫

ρ

ε

∫ ∫

ρ

ε ∞ & & (A.204) 

Now, on the basis of experimental data, it has been determined that: 

 f (D, j ) dD = C(D ) (j ) j

(j ) * 1.5

g

* 1.5

D g

D

0

g g

g

∫

ρ

ε& (A.205) 
A-74 

By differentiating this equation with respect to D, one obtains: 

(D ) (j )

/g

1.5C f (D, j ) = j

(j ) * 1.5

g

* 0.5

D g

g g

g

ρ σ ∆ρ

ε& (A.206) 

By comparing to the experimental data and utilizing equation (A.203), Ishii obtained: 

(D ) (j )

/g

0.077

f (D, j ) = * 1.5

g

* 0.5

D g

σ ∆ρ

 (A.207) 

To obtain the velocity distribution function, Ishii first made a rough model based on the breakup 

of liquid ligaments to determine the initial velocity of a droplet of diameter D in a vapor stream of 

superficial velocity jg. Such a breakup mechanism is consistent only with churn flow in the pool 

and, therefore, is inapplicable for pools in the bubbly or annular flow regimes. Accordingly, 

Ishii's entire model is limited to situations where there is churn flow in the pool because the 

initial velocity distribution derived from this assumption impacts all the subsequent correlations. 

With respect to this initial velocity, Ishii set up a rough equation based on the above mechanism 

and utilized experimental data to obtain [Ref. 16]:

⎥

⎦

⎤ ⎢

⎣

⎡

∆ρ

ρ ⎥

⎦

⎤ ⎢

⎣

⎡

ρ

ρ

μ

g

 0.115

*0.21

H

f

g

 1/ 2

* 1/ 4 1/ 4 * -1/ 4

g

*

i v = 75.17 j N D D g

 (A.208) 

Moreover, since no information was available on the distribution of initial velocities, Ishii 

assumed that a single initial velocity characterizes a droplet of a given size, D , such that the 

distribution reduces to a delta function: 

⎟

⎟

⎟

⎠

⎞

⎜

⎜

⎜

⎝

⎛

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

∆ρ

ρ

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

ρ

ρ δ

μ

g

 0.115

*0.21

H

f

g

 1/ 2

* 1/ 4 1/ 4 * -1/ 4

g

* i g i g(v ,D, j ) = v - 75.17 j N D D g

 (A.209) 

Having determined expressions for all parameters in equation (A.188), the entrainment as a 

function of height can now be determined. To this end, observe from equation (A.208) that the 

initial velocity of a given droplet decreases with increasing size. Nonetheless, since there is a 

maximum droplet size, even the heaviest droplets will possess a minimal initial velocity and, 

therefore, will rise a definite height before falling back into the pool. Thus, there will be no fall out 

before a minimum height is attained, so there will be a near-surface region where the entrainment 

is constant. Above this height, entrainment will decrease as first the heavier and then the lighter 

droplets within the supercritical range fall back into the pool. This second region is, therefore, 

called the momentum-controlled region because the entrainment at any given height is 

dependent upon the droplets having sufficient initial momentum to reach the height in question. 

Finally, after all the supercritical droplets have fallen out of the flow, the entrainment will consist 

only of those droplets small enough to be suspended by the vapor drag. Thus, above a certain 

height the droplets stop falling back into the pool and are taken out of the flow entirely as a 

result of deposition on solid surfaces. For this reason, the third region is referred to as the 

deposition-controlled region. 
A-75 

The extent of the near-surface region can be determined by setting the initial velocity of the 

largest droplets equal to the initial velocity, vh , needed to attain a particular height, h, and then 

solving for this height. The velocity of the largest droplets can be obtained by substituting the 

maximum diameter from equation (A.50) into equation (A.208). As usual, vh is given by 

equation (A.200): 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

∆ρ

ρ

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

ρ

ρ

⎥⎦

⎤ ⎢⎣

⎡

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

ρ

ρ

μ

g

 0.115

*0.21

H

f

g

 1/ 2

*

g

-1 -1/ 4 * 1/ 4 1/ 4

g

f

g

1/ 2

* 2 h = 75.17 j N 7.24 j D g

 (A.210) 

From this relation the near surface region can be shown to extend from: 

⎥

⎦

⎤ ⎢

⎣

⎡

∆

≤ ≤

ρ

ρ

μ

g

 0.23

* 0.42

H

* 1/ 2

g

* 3 0 h 1.038 x10 j N D g (A.211) 

where the upper limit on h* is obtained by solving equation (A.210). Within this region, vh is 

effectively zero since vh < vi for all droplet sizes. Thus, using equations (A.188) and (A.202), 

the entrainment becomes: 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

∆ρ

ρ

ρ

ε

∫ ∫

ρ

ε ∞ g

-1.0

-3

g g

g

i g i D g v

D

0

g g

g

fg g = 4.84 x10 j

(j ) f (D, j ) g(v ,D, j ) dv dD = j(j )

E (h, j ) = h

max

& & (A.212) 

Above the near-surface region is the momentum-controlled region where the heavier droplets 

within the supercritical range progressively fall back into the pool. The lower limit for this region 

is given by the right-hand side of equation (A.211). The upper limit is determined by the point at 

which all the supercritical droplets have fallen out of the stream so that only subcritical droplets 

remain. To determine this limit, it is first necessary to substitute equation (A.196) into equation 

(A.208) to get the initial velocity of a critically sized droplet and then to set the resulting 

expression equal to the initial velocity, vh , given by equation (A.200). When these steps are 

performed, the following expression is obtained: 

[ ] ⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

∆ρ

ρ

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

ρ

ρ

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

ρ

ρ

μ μ

g

 0.115

*0.21

H

f

g

 1/ 2

* 1/ 3

g

 -1/ 4 * 1/ 4 1/ 4

g

f

g

1/ 2

* 2 h = 75.17 j N 4 j (N ) D g g (A.213) 

On the basis of this relation and equation (A.196), the momentum-controlled region can be 

shown to extend from: 

⎥

⎦

⎤ ⎢

⎣

⎡

∆ ⎥ ≤ ≤

⎦

⎤ ⎢

⎣

⎡

∆ ρ

ρ

ρ

ρ

μ μ

g

 0.23

* 0.42

H

g * 3 1/ 3

 0.23

* 0.42

H

* 1/ 2

g

3 1.038 x10 j N D h 1.413x10 N D g g (A.214) 

To determine the entrainment as a function of height within these limits, it is first necessary to 

determine the diameter of the largest supercritical droplet having the initial velocity needed to 

reach the point in question. Denoting this diameter Dh, substituting the variable Dh into equation 

(A.208), and setting the result equal to equation (A.200), yields the following expression: 
A-76 

[ ] ⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

∆ρ

ρ

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

ρ

ρ

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

ρ

ρ

μ

g

 0.115

*0.21

H

f

g

 1/ 2

*

h

* 1/ 4 1/ 4 -1/ 4

g

f

g

1/ 2

* 2 h = 75.17 j N D D g

 (A.215) 

In this expression, Dh should not be confused with the hydraulic diameter DH. If equation 

(A.215) is solved for Dh, one obtains: 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

∆ρ

ρ

μ

g

 0.46

*0.84

* 2 H

*

g

2

4

*

h N D h

j

2

(75.17 ) D = g (A.216) 

Now, since all droplets within the range Dh < D # Dmax lack sufficient initial momentum to reach 

height h, only diameters within the range 0 # D # Dh remain within the flow. Moreover, since all 

droplets remaining within the flow have initial velocities greater than or equal to vh, vh is 

effectively zero for these droplets. Hence, the entrainment at height h is given by: 

 f (D, j ) dD

j

(j ) f (D, j ) g(v ,D, j ) dv dD = j

(j )

E (h, j ) = D g

D

0

g g

g

i g i D g v

D

0

g g

g

fg g

h

h

h

∫

ρ

ε

∫ ∫

ρ

ε ∞ & & (A.217) 

Utilizing equations (A.202) and (A.207), and switching to dimensionless coordinates yields: 

E (h, j ) = 4.84 x10 0.077(D ) (j ) dD * 1.5 *

g

D * 0.5

0

g

-1.0

-3

fg g

*

h

∫ ⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

∆ρ

ρ (A.218) 

Finally, integrating this equation subject to the expression for Dh given by equation (A.216) 

yields the entrainment as a function of height in the momentum-controlled region: 

⎥

⎦

⎤ ⎢

⎣

⎡

∆ρ

ρ

μ

g

 -0.31

* 1.25

H

1.5

* 3

* 3

6 g

fg g N D h

j

E (h, j ) = 5.417 x10 g (A.219) 

Above the momentum-controlled region lies the deposition-controlled region, which only 

droplets within the subcritical diameter range can reach. Since these droplets can be 

suspended entirely by vapor drag, vh = 0 for these droplets in accordance with equation (A.191). 

Hence, the entrainment is given by: 

 f (D, j ) dD

j

(j ) f (D, j ) g(v ,D, j ) dv dD = j

(j )

E (h, j ) = D g

D

0

g g

g

i g i D g v

D

0

g g

g

fg g

c

h

c

∫

ρ

ε

∫ ∫

ρ

ε ∞ & & (A.220) 

where Dc is the diameter of a critically sized droplet. Utilizing equations (A.202) and (A.207), 

and switching to dimensionless coordinates yields: 

E (h, j ) = 4.84 x10 0.077(D ) (j ) dD * 1.5 *

g

D * 0.5

0

g

-1.0

-3

fg g

*

c

∫ ⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

∆ρ

ρ (A.221) 
A-77 

Integrating this equation subject to the expression for Dc given by equation (A.196) gives the 

entrainment in the deposition-controlled region: 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

∆ρ

ρ

μ

g

 -1.0

* 3 1/ 2

g

-3

fg g E (h, j ) =1.988 x10 j N g

 (A.222) 

As written, equation (A.222) predicts that entrainment remains constant with height. However, 

because of deposition on solid surfaces, the entrainment will decrease. Ishii makes the case 

that the deposition rate should be proportional to the entrainment itself, so that the entrainment 

decreases exponentially with height. On the basis of experimental data, Ishii determines an 

exponential attenuation term and adjusts the leading constant to yield: 

E (h, j ) = 7.13 x10 j N exp (-0.205 (h/ DH)) g

 -1.0

* 3 1/ 2

g

-4

fg g g

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

∆ρ

ρ

μ (A.223) 

Then, to ensure a smooth transition at the boundary and hence consistency among his 

equations, Ishii sets equation (A.223) for the deposition-controlled region equal to equation 

(A.219) for the momentum-controlled region to obtain the following equation for the transition 

height [Ref. 16]:

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

∆ρ

ρ

μ

g

 0.23

* 0.42

H * 3 1/ 3

H * * h exp (-0.068 (h / D )) =1.97 x10 N D g

 (A.224) 

When the deposition term is small, the equation can be approximated as: 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

∆ρ

ρ

μ

g

 0.23

* 0.42

H * 3 1/ 3 h =1.97 x10 N D g

 (A.225) 

The entrainment correlations, Efg, listed above can be expressed in terms of mass flow rate 

using the following definition: 

j

j E =

g g

f f

fg

ρ

ρ (A.226) 

and the liquid mass continuity equation: 

m& ENT = ρf j

f aup (A.227) 

Substituting equation (A.226) into (A.227) yields: 

m& ENT = ρg j

g Efg aup (A.228) 

This places the entrainment equations in the form needed to evaluate the dimensionless groups 

obtained in the top-down scaling analysis. 
A-78 

In summary, based on Ishii's pool entrainment model, the following entrainment relations are 

predicted [Ref. 16]:

Near-Surface Region: 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

∆ρ

ρ

μ

g

 0.23

* 0.42

H

* 1/ 2

g

* 3 0 LE h LE1.038 x10 j N D g

 (A.229) 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

∆ρ

ρg

-1.0

-3 Efg (h, j

g) = 4.84 x10 (A.230) 

m = 4.84 x10 j

g aup

-3 & ENT ∆ρ (A.231) 

Momentum-Controlled Region: 

⎥

⎦

⎤ ⎢

⎣

⎡

∆ ⎥ ≤ ≤

⎦

⎤ ⎢

⎣

⎡

∆ ρ

ρ

ρ

ρ

μ μ

g

 0.23

* 0.42

H

g * 3 1/ 3

 0.23

* 0.42

H

* 1/ 2

g

3 1.038 x10 j N D h 1.97 x10 N D g g (A.232) 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

∆ρ

ρ

μ

g

 -0.31

* 1.25

H

1.5

* 3

* 3

g 6

fg g N D h

j

E (h, j ) = 5.603 x10 g

 (A.233) 

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

∆ρ

ρ

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

ρ

σ ∆ρ ρ μ

g

 -0.31

* 1.25

H

1.5

2

g

.25

* 3

* 4

g

g UP

6

ENT N D g

h

j

m = 5.603x10 a g & (A.234) 

Deposition-Controlled Region: 

1.97 x10 N D h g *

 0.23

* 0.42

H

3 1/ 3

g ⎥ ≤

⎦

⎤ ⎢

⎣

⎡

∆ρ

ρ

μ (A.235) 

E (h, j ) = 7.13 x10 j N exp (-0.205 (h/ DH)) g

 -1.0

* 3 1/ 2

g

-4

fg g g

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

∆ρ

ρ

μ (A.236) 

N exp (-0.205 (h/ D )) g

m = 7.13 x10 a j H

g

 -1.0

1/ 2

2

g

.25

* 4

g UP g

-4

ENT g

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

∆ρ

ρ

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

ρ

σ ∆ρ ρ μ & (A.237) 

As previously mentioned, these equations are built upon a particular entrainment mechanism 

under the assumed conditions of churn flow in the pool. Thus, the validity of these equations 

hinges upon the satisfaction of this initial assumption. 
A-79 

A.7.2.2 Upper Plenum Liquid Entrainment Rate Scaling Criteria 

The upper plenum geometry and flow conditions can be determined by substituting 

the entrainment correlations into the scaling criterion given by equation (A.187). 

Near-Surface Region: 

Substituting equation (A.231) into (A.187) and assuming fluid property similitude, yields the 

following relation: 

2

1 j = g,R (A.238) 

This ratio is established by controlling the core power and fixing the upper plenum flow area. 

Performing a simple steady-state energy balance for the core yields: 

j aUP (hfg - hSC) = q g g ρ ∆ & (A.239) 

where DhSC is the core subcooling, and q is the core power. Rearranging and substituting into 

equation (A.238) yields: 

2

1 = a

q

UP R

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ & (A.240) 

Currently, the APEX upper plenum flow area, aUP, ratio is approximately equal to [ ]a,b,c. 

Thus, the core power scaling ratio becomes: 

96

1

q = R& (A.241) 

The requirements for liquid entrainment in the near-surface region can be easily met. 

Momentum-Controlled Region: 

Substituting equation (A.234) into equation (A.187) and assuming fluid property similitude 

yields: 

2

1 = h

j D

3

1.25

H

4

g

R

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

 (A.242) 

Fixing jg,R to 1:2 and hR to 1:4 yields the following scaling ratio for the hydraulic diameter of the 

upper plenum: 

= 0.1895

8

1

D =

0.8

H,R ⎟

⎠

⎞ ⎜

⎝

⎛ (A.243) 

This new scaling criterion shall be applied in the upper plenum design section. 
A-80 

Deposition-Controlled Region: 

Substituting equation (A.237) into equation (A.187) and assuming fluid property similitude 

yields: 

2

1 = Dh j exp -.205

H

4

g

R

⎥

⎦

⎤ ⎢

⎣

⎡

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ (A.244) 

The form of this criterion requires that it be assessed numerically for the model and the 

prototype. 

A.7.2.3 Scaling Criterion for Lateral De-Entrainment 

A study of droplet de-entrainment on vertical tubes in droplet cross-flow was performed by 

Dallman and Kirchner at Los Alamos Scientific Laboratory in 1980 [Ref. 17]. The study included 

both analytical and experimental research. Droplets of a known size were sprayed 

perpendicular to vertical conduits situated in a draft induced wind tunnel. The two geometries 

examined were square channels (76.2 mm square) and cylindrical tubes (25.4 mm, 63.5 mm, 

and 101.6 mm diameters). The air velocities were varied from 0 m/s to 14 m/s and the liquid 

droplet velocities were varied from 10 m/s to 15 m/s. It was found that the de-entrainment 

efficiency for a single tube was only weakly dependent on the tube diameter, air velocity, droplet 

mean velocity, total liquid flow rates for air velocities below 14 m/s, and droplet mass flux rates 

greater than 2 kg/(m2

-s). In fact, the de-entrainment rate was nearly a constant value of 0.19 for 

cylinders and 0.27 for square tubes for this range. That is: 

= 0.19 (For Cylinders) ηI (A.245) 

and 

= 0.27 (For a 76.2mmSquare Tube) ηI (A.246) 

The authors cited the work of Chen [Ref. 18] as providing a method of extrapolating the single 

tube data to an array of tubes. Namely, the de-entrainment efficiency for horizontal droplet flow 

through “k” rows of identical vertical tubes with a constant diameter to pitch ratio is: 

= A[1- C(1- (1+ 4.5(D/P ) ) ] 2 k

ηk ηI (A.247) 

where C is a geometric factor dependent on pitch to diameter ratios. The constant A is unity for 

a staggered array with no line of sight through the array. Equation (A.247) indicates that 

preserving the upper plenum guide tube de-entrainment efficiency requires preserving the guide 

tube diameter to pitch ratio and the number of guide tube rows. Therefore, the first criterion can 

be written as: 

(D/P ) =1 R (A.248) 
A-81 

For the diameter-to-pitch ratios in this study, very high entrainment efficiencies (i.e., greater than 

90%) are achieved with just three staggered rows of guide tubes. Therefore, the number of 

guide tube rows beyond three does not contribute much to the de-entrainment process. 

Furthermore, an examination of the AP1000 upper plenum structure geometry reveals the 

following [Refs. 16, 17, and 18]:

(1) The upper core plate includes a significant number of holes at its periphery. De-entrainment 

of liquid from the peripheral holes will occur on the core barrel walls, because there are no 

guide tubes to impede the flow. 

(2) There are several [ ]a,b,c diameter flow holes located adjacent to each HL. 

No de-entrainment will occur at those sites. Thus, the entrainment efficiency at those sites 

is zero. Modeling of those flow holes in the APEX test facility will be important. 

(3) Liquid from the central internal flow holes will be de-entrained at a relatively high de-entrainment 

efficiency (i.e., on the guide tubes). Thus, the flow through the peripheral flow holes will 

dominate the behavior. 

(4) The flow path of least resistance for the fluid is radially outward toward the periphery of the 

upper internals. 

The upper plenum can be effectively divided into two zones: a peripheral zone that exhibits no 

de-entrainment and an internal guide tube zone with a relatively high de-entrainment efficiency. 

Rather than preserving the row numbers, the scaling approach shall be to preserve the flow 

area averaged de-entrainment efficiency for the internal guide tube zone. That is, the following 

scaling criterion will be applied: 

=1 ηeff,R (A.249) 

Effective De-Entrainment Efficiency 

For purpose of this analysis, the AP1000 guide tube bundle has been divided into five flow zones, 

from I = 4 at the center of the upper plenum to i = 0 at the periphery of the upper plenum. 

The total liquid mass flow rate entering the flow zones through the upper core plate and 

traveling to the periphery of the guide tube bundle is given by: 

m = m (1- ) k i & Periphery Σ & i η (A.250) 

where k represents the number of guide tube rows between a flow zone and the guide tube 

bundle periphery. For the AP1000, the upper plenum flow zones have been selected such that 

k = i. Dividing by the total liquid mass flow rate through the upper core plate yields the fraction 

of entrained liquid reaching the periphery. One minus this value yields the equation for the 

desired effective de-entrainment efficiency. That is: 

m

m (1- ) =1- UCP

k i i

eff &

Σ & η

η (A.251) 
A-82 

Substituting equation (A.247) into (A.251) yields: 

[1- A(1- C )]

m

m =1- i

k

UCP

i

eff ⎟

⎟ Φ ⎠

⎞

⎜

⎜

⎝

⎛

η Σ &

& (A.252) 

where: 

=1- [1+ 4.5(D/P ) ]

2 Φ ηI (A.253) 

Furthermore, for the case of a constant inlet mass flux, ρv, equation (A.253) becomes a flow 

area averaged de-entrainment efficiency: 

[1- A(1- C )]

a

a =1- i

k

UCP

i

eff ⎟

⎟ Φ ⎠

⎞

⎜

⎜

⎝

⎛

η Σ (A.254) 

For the case of A and C equal to 1, and i = k, this equation reduces to: 

⎟

⎟Φ

⎠

⎞

⎜

⎜

⎝

⎛

η Σ i

UCP

i

eff

a

a =1- (A.255) 

Equations (A.247) and (A.249) shall be used to help scale the lateral de-entrainment 

phenomenon. 

A.7.2.4 Scaling Criterion for the Upper Core Plate 

The two key phenomena to be preserved by scaling the upper core plate flow holes are: 

• Upper plenum flow distribution 

• Counter-current flooding limitation (CCFL) at the core upper support plate holes. 

Flow through the upper core plate at the guide tube bundle periphery will dominate the 

entrainment behavior in the upper plenum. In particular, the flow area directly adjacent to the 

HLs has unimpeded access to the HLs. Thus, preserving the flow distribution in the upper core 

plate will be essential to proper simulation of the entrainment behavior in the upper plenum. 

The following scaling criterion will be implemented: 

=1

a

a

UCP

Periphery

R

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ (A.256) 

Similarly, the upper core plate flow area at the entrance to the HLs shall also be preserved as 

follows: 

=1

a

a

UCP

HL

R

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ (A.257) 
A-83 

Preserving the counter-current flooding behavior in the upper core plate requires that the 

Kutateladze number flooding limit, defined in terms of jg* in this analysis, be preserved in APEX. 

That is: 

= 4

g

j

2

g

1/ 4

g

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

ρ

σ ∆ρ

 (A.258) 

For the condition of fluid property similitude, properly matching the Kutateladze number requires 

that jg in APEX match that in AP1000. That is: 

(j ) =1

g,UCP R (A.259) 

Thus, the flow area of the APEX upper core plate must be scaled as 1:96, as follows: 

96

1 (a ) = UCP R (A.260) 

Furthermore, the minimum upper core plate hole diameter must be greater than ~2.0 inches, so 

that the onset of the flooding behavior is not dependent on the hole diameter. 

A.7.2.5 Scaling Criterion for the Upper Support Plate 

The upper support plate separates the upper plenum from the reactor vessel head. The plate 

provides support for the control rod guide tubes and includes 12 [ ]a,b,c diameter holes 

that permit the upper head to drain into the upper plenum. During the pool entrainment process, 

it is expected that the upper head will be empty of liquid. Nonetheless, it has been included to 

document the scaling basis for this component. 

Because the flow area represented by the drain holes is quite small, the flooding behavior and 

flow distribution cannot both be simultaneously preserved. Since the upper head is essentially 

a dead-ended volume (there is a small bypass flow path from the upper head to the upper 

downcomer), flooding behavior will not be as important as the flow distribution. That is, CCFL is 

not expected to occur in the plate. The dominant phenomenon will be the pressurized draining 

of the upper head as a result of the flashing of liquid during a depressurization event. As a result, 

the drain hole number and distribution will be preserved and the flow area will be scaled as follows: 

48

1 (a ) = USP R (A.261) 
A-84 

A.7.2.6 Design of APEX Upper Plenum Guide Tube Bundle, Upper Core Plate, and Upper 

Support Plate 

Based on the top-down and bottom-up scaling analyses presented in the previous sections, 

the following scaling approach was adopted for the upper plenum guide tubes, upper core plate, 

and upper support plate: 

• Preserve the upper plenum hydraulic diameter ratio using equation (A.243). 

• Preserve the local upper plenum subchannel hydraulic diameter ratio using equation (A.243). 

• Preserve the total upper plenum flow area based on [ ]a,b,c flow area scaling. 

• Preserve the guide tube diameter to pitch ratio using equation (A.248). 

• Preserve the effective de-entrainment efficiency for the guide tube bundle using equation (A.249). 

• Preserve the upper core plate flow area distribution using equations (A.256) and (A.257). 

• Preserve the vapor volumetric flux, jg, in the upper core plate holes using equation (A.260). 

• Preserve the upper support plate flow distribution and flow area using equation (A.261). 

With regard to the scaling basis for the upper plenum guide tube bundle, the momentumcontrolled entrainment scaling criterion given by equation (A.243) has been selected to guide 

the modification to the APEX upper internals. The basis for this selection is as follows. 

The near surface entrainment rates can be properly simulated in APEX solely by maintaining jgR

at a constant value of 0.5. In that sense, no modification to the facility would be needed. 

However, the momentum-controlled entrainment rates can only be matched by preserving both 

the jgR and the hydraulic diameter ratios. It will be shown that the deposition-controlled region 

is both unimportant and not achievable in the one-quarter length scale APEX facility. 

Figure A.21 depicts the upper core plate geometry. There are a total of 13 flow holes. The six 

holes on the periphery are 1.91 inch in diameter and represent approximately 37% of the total 

upper core plate flow area, compared to 38% in the AP1000. The flow holes at the entrance 

of each HL represent 6.2% of the total flow area as in the AP1000. 

Figure A.21 APEX Upper Core Plate Geometry 

a,b,c 
A-85 

Figure A.22 illustrates the APEX upper core plate guide tube pitch and diameter. Figure A.23 

presents the APEX upper support plate flow hole location. Table A.18 presents the revised 

upper plenum geometry design based on matching the scaling criteria given above. 

1.95 in

1.689 in

Figure A.22 APEX Upper Plenum Guide Tube Pitch and Diameter 

Figure A.23 APEX Upper Support Plate Flow Hole Locations 

a,b,c 
A-86 

Table 18 Comparison of the APEX and AP1000 Upper Plenum Geometries

Component AP1000 APEX Units 

UPPER PLENUM 

Core Barrel Inside Diameter 

Length* 

Flow Area with Internals 

Total Wetted Perimeter 

Total Hydraulic Diameter 

SUPPORT COLUMNS 

Number 

Diameter 

GUIDE TUBES 

Number 

Pitch 

Equivalent Diameter 

Pitch-to-Diameter Ratio 

Subchannel Flow Area 

Subchannel Wetted Perimeter 

Subchannel Hydraulic Diameter 

UPPER CORE PLATE 

Number of Large Thru Holes 

Total Thru Hole Flow Area 

Number of Guide Tube Flow Holes 

Total Guide Tube Flow Area 

Number of Support Column Flow Holes 

Total Support Column Flow Area 

Total Upper Core Plate Flow Area 

Peripheral Flow Area 

Ratio of Peripheral Flow Area to Total Area 

HL Entrance Flow Area 

Ratio of HL Entrance Flow Area to Total Area 

UPPER SUPPORT PLATE 

Number of Holes 

Hole Diameter 

Total Flow Area 

*Bottom of HL to Top of Active Fuel 

a,b,c 
A-87 

A.7.2.7 Assessment of Pool Entrainment Limiting Conditions 

This section examines the limiting conditions for the APEX entrainment tests. A core power of 

1 MW in APEX represents a 2.8% decay power in the AP1000. The greatest vapor generation 

rates would be achieved at saturation conditions at 1 atmosphere. These conditions would 

produce an upper plenum vapor volumetric flux, jg, of 11.2 m/s and 5.7 m/s in AP1000 and 

APEX, respectively. Applying these conditions to the entrainment transition and rate 

correlations presented in the previous section results in Figure A.24. 

0 0.2 0.4 0.6 0.8 1

Dimensionless Height (Bottom of Hot Leg to Top of Active Core)

APEX-AP1000 AP1000

NEAR 

SURFACE

REGION

MOMENTUM CONTROLLED

REGION

DEPOSITION CONTROLLED

REGION

NEAR 

SURFACE

REGION

MOMENTUM CONTROLLED

REGION

Bottom of 

Hot Leg

Top of Active Core

Figure A.24 Assessment of AP1000 and APEX Upper Plenum Pool Entrainment Regimes 

for a 2.8% Core Decay Power at Atmospheric Pressure 

This figure shows how the different pool entrainment regions would exist in the upper plenums 

for the prescribed set of conditions. As the level drops below the HL, both the AP1000 and 

APEX upper plenums would exhibit entrainment characteristic of the near surface region. 

The entrainment rates that would be observed in APEX in this region are well-scaled, as is the 

transition to the momentum-controlled region. It is very important to note that the entrainment 

rates in the near surface region are a factor of 150 times greater than the entrainment 

rates generated in the momentum-controlled region. 

The figure shows that well-scaled entrainment behavior in APEX is expected to a depth of about 

18.9 cm (0.55 in terms of dimensionless height in Figure A.24) below the HL. Below this depth, 

APEX is in the momentum-controlled region, whereas AP1000 would be in the depositioncontrolled region at the equivalent scaled depth. It is also important to note that the 

entrainment rates in the momentum-controlled region are a factor of 4–6 times greater 

than the entrainment rates generated in the deposition-controlled region. 
A-88 

With regard to the test facility scaling, the entrainment rate is well-scaled for depths to 

approximately the middle of the upper plenum. Below this depth, the APEX entrainment rates 

will range from 4 to 16 times greater, on a scaled basis, than what would be expected in the 

AP1000. Thus, the APEX facility will conservatively simulate the entrainment behavior when the 

level is below the midpoint of the upper plenum. 

In general, relative to the near-surface region and momentum-controlled region entrainment 

rates, the deposition-controlled region entrainment rates are quite small. 

On the basis of the conditions prescribed earlier, some preliminary calculations can be made to 

narrow down the flow conditions in the problem. First, since it is imperative that churn flow 

exists in the pool, it is necessary to determine if the flow is between the limits of bubbly and 

annular flow. For bubbly flow, Ishii gives the limit as [Ref. 17]:

j = 0.325 = 0.008

f

g

 1/ 2

*

g

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

ρ

ρ (A.262) 

where the previous density values have been substituted into Ishii's formula to give the 

numerical result. For annular flow, the criterion is given by [Ref. 17]:

j = 3.1 *

g (A.263) 

For these conditions, the value of jg* was found to be 1.75 for AP1000, and 0.89 for APEX. 

Thus, the jg* values fall between these limits for both the model and the prototype. Thus, churn 

flow exists in the pool for both systems over the full range of operating conditions, making Ishii's 

equations applicable to the analysis. 

The next issue to be resolved is whether the flooding criterion is exceeded. After all, if the 

flooding criterion is exceeded, all of the liquid will be quickly expelled from the pool as a rising 

film, making the entrainment of droplets irrelevant. As is well known, the flooding criterion 

depends upon the diameter of the channel. For diameters in the range 3 # DH* # 20, Ishii 

restates the Wallis criterion in the following form [Ref. 14]:

= 0.5

D

j

*

H

*

g (A.264) 

Furthermore, for diameters within the range DH* $ 30, Ishii gives the Kutateladze criterion as 

[Ref. 17]:

j = 4 *

g (A.265) 

The dimensionless hydraulic diameters, DH*, for AP1000 and APEX were for these conditions 

were found to be 135 and 29, respectively. Using equation (A.248) for APEX and AP1000 

or equation (A.247) for APEX, the criterion indicates that flooding in the upper plenum will not 

occur for the 2.8% power conditions described. 
A-89 

The next question to be addressed is whether the deposition on the wall is sufficiently large that 

one must consider re-entrainment from a falling film back into the vapor. In addressing this 

question, observe that only a fraction of the droplets are taken out of the flow by deposition. 

Furthermore, only a fraction of this fraction is re-entrained from the film back into the flowing 

vapor. Consequently, since only a fraction of the liquid removed by deposition is re-entrained 

back into the flow, it would seem that conservative results could be obtained by ignoring 

deposition in the first place. The only objection to this conservatism would be if deposition 

removes large droplets from the flow which are replaced by smaller droplets as a result of 

re-entrainment. Since this effect could shift the droplet distribution to subcritical diameters 

that are more easily carried by the vapor, the entrainment might be increased at large heights. 

However, a second and quite convincing argument is as follows. Since the data used by Ishii 

were based on flow in tubes, the phenomena of deposition and re-entrainment was necessarily 

present in the experiments themselves and, therefore, already included in the correlations. 

Given this fact, the equations may be used straightforwardly without having to account for this 

added effect. 

A.7.3 Evaluation of Scale Distortion 

Properly simulating the upper plenum pool entrainment process requires that the following ratios 

be on the order of unity: 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

a ρ H

m t

t = UP L

ENT,O

R

+

R

& (A.266) 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

Π

m

m = ENT,O

INJ,O

R

INJ,R

&

& (A.267) 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

Π

m

m = ENT,O

GT,O

R

GT,R

&

& (A.268) 

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

Π

m

m = ENT,O

UPD,O

R

UPD,R

&

& (A.269) 

The numerical values for each of the non-dimensional ratios are presented in Table A.17 for the 

near-surface entrainment and momentum-controlled entrainment regions. The data obtained 

from the assessment of limiting conditions presented in the previous section were used in the 

evaluation. The desired timescale ratio, tR, of 0.5 was used to assess equation (A.266). All of 

the injection flow rates have been scaled to 1:96. Furthermore, equation (A.269) can be 

interpreted as the ratio of effective de-entrainment efficiencies. That is: 

ΠGT,R ηeff,R = (A.270) 

The facility has been scaled to match the onset of upper plenum draining behavior when CCFL 

conditions are present at the upper core plate flow holes. 

The ratio of the gravity-driven draining rates for the APEX and AP1000 upper plenums is given 

by: 

( ) m = [a (H - h ) ] 1/ 2

UCP UP R UPD,O R & (A.271) 
A-90 

Table A.19 indicates that the APEX-AP1000 upper plenum geometry and long-term cooling 

operating conditions match the pool entrainment scaling criterion very well with the exception of 

PUPD,R. 

Table A.19 Ratio of Π Groups for Near Surface and Momentum Controlled Upper Plenum 

Entrainment 

Non-Dimensional Ratio Near Surface Momentum Controlled 

(t+)R 

Πinj,R 

ΠGT,R 

ΠUPD,R 

(jg

+

)R 

A.7.3.1 Impact of Reduced Flow Area in the Upper Core Plate 

The greatest source of distortion will be the upper plenum draining rate with a Π ratio of 0.5. 

This distortion arises because matching the flooding behavior at the upper core plate required 

reducing the upper core plate flow area such that the Kutateladze number was identically 

preserved. This delayed draining effect is conservative relative to core cooling. In addition, the 

vapor jetting behavior at the upper core plate produces a local vapor volumetric flux jg equivalent 

to what would be observed in the AP1000. Preserving a 1:1 vapor velocity in a 1:4 length upper 

plenum will produce conservative results. It is expected that the proposed set of upper internals 

will overestimate the amount of entrained liquid on a scaled basis. This is a conservative 

feature of the experiment. 

a,b,c 
A-91 

A.8 Summary and Conclusions

This document presents the scaling analysis performed to guide the OSU APEX-AP1000 test 

facility modifications. The purpose of the test program is to provide high-quality benchmark data 

for the reactor safety analysis codes. Previous testing in APEX has demonstrated that all of the 

thermal-hydraulic phenomena observed in the large-scale AP600 facilities such as ROSAAP600 or SPES were also observed in APEX, thereby providing a broad data base for code 

assessment. However, the most faithful representation of the AP600 phenomena was for the 

very long-term cooling phase, which occurs at low system pressure. The APEX-AP1000 test 

program will be modified to include full-pressure ADS4 blowdown tests to investigate the 

transition period that leads to the onset of IRWST injection. This scaling analysis provides the 

basis for the following plant modifications: 

Component Modification to APEX 

Reactor Power Increase core power by 67%. (Maximum of 1MW). 

PZR Increase PZR volume. 

Reduce PZR Surge Line Diameter. 

CMTs Increase CMT volumes by 25%. 

Reduce line resistance to 64% of original value. 

IRWST Increase IRWST liquid level. 

ADS4 Increase ADS4 flow area by 76%. 

Reduce line resistance to 28% of original value. 

PRHR HX Increase PRHR flow capacity by 74% by reducing line 

resistance. No change in surface area/tube number 

required for testing. 

Containment Sump Flood-Up Elevation Change flood-up elevation in primary sump tank. 

The following paragraphs briefly summarize the results of the study. 

(1) The original AP600 Scaling Analysis, WCAP-14270, which was performed in support of the 

design of the existing Advanced Plant Experiment (APEX) facility at OSU was found to be 

applicable to the APEX-AP1000 design for two reasons. First, the thermal-hydraulic 

phenomena to be studied fell within the scope of the original AP600 test program, with the 

exception of upper plenum behavior. Second, the AP1000 is geometrically similar to the 

AP600. As a result, all of the tank volumes, flow areas, and line resistance scaling ratios 

found in WCAP-14270 were used in the present study. 

(2) An NRC PZR surge line study by Di Marzo and Bessette was used to guide the 

modifications to the APEX-AP1000 PZR surge line. The reduced line diameter is expected 

to produce a better simulation of the PZR draining behavior. 

(3) The RCS depressurization scaling analysis represents a more comprehensive approach 

than that implemented in the original AP600 study. Nonetheless, the similarity criteria 

derived in this report resulted in scaling ratios that were identical to those in the original 
A-92 

study. It was found that the characteristic time ratios for AP1000 would be well-matched in 

APEX-AP1000. The greatest distortion (a factor of 2) was the result of pressure-scaled fluid 

properties. 

(4) A scaling analysis has been performed to design the APEX-AP1000 to simulate the 

important phenomena that occur during the transition from ADS4 blowdown to the onset of 

IRWST injection. The key phenomena have been identified and described, and the scaling 

analysis indicates that full-pressure ADS4 blowdown tests can be performed using the 

modified APEX-AP1000 test facility. Chapter 15 of the Westinghouse AP1000 Safety 

Analysis Report indicates that ADS4 actuation typically occurs near 100 psia. The 

maximum decay power, ~2.0%, at the time of ADS4 valve opening was calculated for the 

DEDVI break. These initial conditions will be examined as part of the testing. The AP1000 

core axial void fraction profile and core averaged void fraction were shown to be preserved 

in APEX-AP1000. 

PADS-4

Sweepout of

Liquid Above

Hot Leg

Liquid Boil-off,

Flashing, and

Carryover

RPV MIXTURE LEVEL

PRESSURE

PIRWST

Top of Hot Leg

Top of Core

Two-Phase Fluid

ADS-4 Flow

High Quality Vapor

ADS-4 Flow

IRWST

Injection

(5) A detailed pool entrainment scaling analysis has been performed to design the upper 

plenum components in APEX-AP1000. The analysis indicates that well-scaled entrainment 

behavior in APEX is expected to a depth of about 18.9 (0.55 in terms of dimensionless 

height) cm below the HL. Below this depth, APEX is in the momentum-controlled region, 

whereas AP1000 would be in the deposition-controlled region at the equivalent scaled 

depth. It was found that the entrainment rates in the near-surface region are a factor of 150 

times greater than those generated in the momentum-controlled region, which are a factor of 

4-6 times greater than those generated in the deposition-controlled region. Thus, the APEX 

facility will conservatively simulate the entrainment behavior when the level is below the 

midpoint of the upper plenum. In addition, the vapor volumetric flow rate ratio, (jg,R) , at the 

upper core plate holes has been set to unity to preserve the Kutateladze number. This adds 

an additional conservatism to the pool entrainment process because the local vapor jets at 
A-93 

the upper core plate holes will extend into the upper plenum, thereby increasing liquid 

entrainment. 

0 0.2 0.4 0.6 0.8 1

Dimensionless Height (Bottom of Hot Leg to Top of Active Core)

APEX-AP1000 AP1000

NEAR 

SURFACE

REGION

MOMENTUM CONTROLLED

REGION

DEPOSITION CONTROLLED

REGION

NEAR 

SURFACE

REGION

MOMENTUM CONTROLLED

REGION

Bottom of 

Hot Leg

Top of Active Core
 

APPENDIX B. 

... 
B-1 

APPENDIX B. APEX COMPONENT DIMENSIONS, CONTROL 

LOGIC, AND SETPOINTS 

B.1 Component Dimensions

Tables B.1 through B.11 contain detailed component dimensions (and, in some cases, 

thermocouple locations) for the reactor pressure vessel (RPV), pressurizer (PZR), steam 

generators (SGs), in-containment refueling water storage tank (IRWST), core makeup 

tanks (CMTs), accumulators (ACCs), primary and secondary sumps, and safety system 

line dimensions and resistances. Figure B.1 is a map of the thermocouples in the RPV 

downcomer. 

Table B.1 Reactor Pressure Vessel Components

Component No. Dimensions Reference 

 

 

 

 

 

 

 

 

a,b,c 
B-2 

Table B.1 Reactor Pressure Vessel Components

Component Dimension Reference 

 

 

Figure B. 1 Thermocouple Map of Downcomer 

a,b,c 

a,b,c 
B-3 

Table B.2 Thermocouple Legend for Figure B.1

Thermocouple 

Map Number Tag Number 

Radial 

Position 

Thermocouple 

Map Number Tag Number 

Radial 

Position 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

a,b,c 
B-4 

Table B.2 Thermocouple Legend for Figure B.1

Thermocouple 

Map Number Tag Number 

Radial 

Position 

Thermocouple 

Map Number Tag Number 

Radial 

Position 

 

 

 

 

 

 

 

Table B.3 Pressurizer Components 

Component No. Dimensions Reference 

 

 

 

a,b,c 

a,b,c 
B-5 

Table B.4 Steam Generators (Each) Components 

Component No. Dimensions Reference 

 

 

 

 

a,b,c 
B-6 

Table B.5 IRWST Components 

Component No. Dimensions Reference 

 

 

 

 

 

Table B.6 Core Makeup Tank Components 

Component No. Dimensions Reference 

 

 

 

a,b,c 

a,b,c 
B-7 

Table B.7 Accumulator Components 

Component No. Dimensions Reference 

 

 

 

Table B.8 Primary Sump Components 

Component No. Dimensions Reference 

 

 

 

a,b,c 

a,b,c 
B-8 

Table B.9 Secondary Sump Components 

Component No. Dimensions Reference 

 

 

 

a,b,c 
B-9 

Table B.10 Summary of Safety System Line Dimensions 

Name Material Data and Dimensions Location and Connections (Valves and Equipment) 

 

 

a,b,c 
B-10 

B.2 Control Logic and Controller Setpoints

Tables B.12 through B.23 contain detailed control logic and setpoints for station 

blackouts; SBLOCAs; LBLOCAs; inadvertent ADS initiation; full-pressure ADS initiation; 

automatic actuation of PRHR and CMT on “S“ signal; automatic actuation of ADS1-4 on 

CMT level or timers; automatic control of SG pressure, liquid water level, feedwater flow, 

and main steam flow; automatic control of PZR pressure and level; programmed reactor 

decay power and control of core exit temperatures; and programmed CVS and RNS 

pump flow rates. 

Table B.12 Station Blackout Routine Logic 

Step Condition 

 

a,b,c 
B-11 

Table B.13 Small-Break LOCA Routine Logic 

Step Condition 

 

a,b,c 
B-12 

Table B.14 Large-Break LOCA Routine Logic (Much the same as Small-Break 

LOCA routine logic except the test initialization) 

Step Condition 

 

Table B.15 Inadvertent ADS Initiation Routine Logic 

Step Condition 

 

a,b,c 

a,b,c 
B-13 

Table B.16 Full-Pressure ADS Routine Logic 

Step Condition 

 

Natural Circulation Routine 

This routine permits operation of the reactor controller in manual or automatic mode, 

with or without RCPs and with no “S” signal or automatic trip functions. This routine 

allows flexibility to custom-tailor a test without requiring extensive re-programming. 

The setpoint in Table B.17 show each bi-stable device and its associated set/reset points. 

Most of these values can be readily modified by the operator and, in some cases, vary 

among the different logic routines. The table contains the tag name of the device, 

the specific point (e.g., HH for High High), the set/reset points applicable to the logic 

routine, and the logic diagram sheet number(s) showing the device. 

a,b,c 
B-14 

Table B.17 APEX Controller Set Point Values for Natural Circulation Routine

 Routine Trip Reset Action(s) 

 

 

 

 

 

 

 

 

 

 

 

a,b,c 
B-15 

Table B.17 APEX Controller Set Point Values for Natural Circulation Routine

 Routine Trip Reset Action(s) 

 

 

 

 

 

 

a,b,c 
B-16 

Table B.17 APEX Controller Set Point Values for Natural Circulation Routine

 Routine Trip Reset Action(s) 

 

 

 

 

 

 

 

a,b,c 
B-17 

Table B.17 APEX Controller Set Point Values for Natural Circulation Routine

 Routine Trip Reset Action(s) 

 

 

a,b,c 
B-18 

Table B.18 Automatic Actuation of PRHR and CMT on “S” signal 

Conditions (And) Actions 

 

 

Table B.19 Automatic Actuation of ADS 1-4 on CMT Level or Timers 

Conditions Actions 

 

 

a,b,c 

a,b,c 
B-19 

Table B.20 Automatic Control of SG Pressure, Liquid Water Level, 

Feedwater Flow, and Main Steam Flow 

Input Signal Output signal to 

 

a,b,c 
B-20 

Table B.21 Automatic Control of PZR Pressure and Level 

Input Signal Output signal to 

Table B.22 Programmed Reactor Decay Power and Control of Core Exit Temps 

Input Signal Output signal to 

Table B.23 Programmed CVS and RNS Pump Flow Rates 

Input Signal Output signal to 

 

 

 

 

 

 

 

 

a,b,c 

a,b,c 

a,b,c 
 

APPENDIX C. 

... 
C-1 

APPENDIX C. TEST TRANSIENT PLOTS AND SEQUENCE OF EVENTS 

This appendix contains transient plots, facility configuration, and sequence of events for tests 

NRC-AP1000-01, 02, 03, 04, 05, 06, 10, and 11. Facility information can be found in Chapter 3 

and Appendix B of this report. The main purpose of the APEX-AP1000 confirmatory testing was 

to examine the performance of passive safety systems. 

C.1 NRC-AP1000-01

The purpose of this test was to determine the impact on passive safety system behavior that 

resulted from a DEDVI break with a major failure of ADS system. It was similar in intent to 

Test NRC-26, which was run for AP600 design certification, in the sense that in NRC-26, a 

major failure of the ADS system was simulated. However, Test NRC-26 included delayed core 

heatup to simulate containment failure, and was configured for a smaller 1-inch CL break, and 

the delayed core heatup in NRC-26 resulted in a heater rod temperature excursion. 

Test NRC-AP1000-01 was run with a larger DEDVI break without delayed core heatup. 

The core remained covered during the long-term cooling period of Test NRC-AP1000-01. 

Both Test NRC-26 and Test NRC-AP1000-01 experienced short periods of CMT refill as a result 

of condensation-induced pressure drops. A similar test in the ROSA-AP600 facility also 

investigated a similar scenario. In Test ROSA AP-CL-05, there was no refilling during CMT 

draining. The difference in ROSA and APEX CMT behavior was attributed to the use of two 

separate CMT-PBLs (APEX) instead of one (ROSA) and the reduced height scale of APEX 

(permitting liquid to be drawn through the PBL into the CMTs, which further enhanced 

condensation). 

Test NRC-AP1000-01 simulated a DEDVI break with complete failure of ADS1, 2, and 3. 

Table C.1 summarizes the facility configuration. This test was defined as a beyond-DBA 

scenario because of the multiple valve failures. All four ADS4 valves were available and, 

therefore, this test investigated the ability of the ADS4 system to accomplish system 

depressurization. The PRHR HX was inactive during this test. The APEX facility models 

both trains of the AP1000 ADS1, 2, and 3 lines located at the top of the PZR after a flow nozzle 

that models the scaled choked flow area of the AP1000. 

Table C.1 Facility Configuration for Test NRC-AP1000-01 

Component Configuration 

ADS1 Failed Shut 

ADS2 Failed Shut 

ADS3 Failed Shut 

ADS4-1 (non-PZR side) 100% ADS4-1 flow 

ADS4-2 (PZR-side) 100% ADS4-2 flow 

Break DEDVI Line 1 
C-2 

Table C.2 summarizes the sequence of events for Test NRC-AP1000-01. After the break was 

opened at zero seconds, the PZR, CMTs, and ACC1 immediately began to inject. However, 

all Side 1 injection flows (ACC1 and CMT1 in Figure C.3), went out the DEDVI Line 1 break. 

The core collapsed liquid level reached a relative minimum at [ ]a,b,c as shown in 

Figures C.1 and C.2. The ACC2 injection began at [ ]a,b,c as shown in Figure C.4. 

The ADS4 valves opened just before the minimum RPV liquid level at around [ ]a,b,c. 

The RPV two-phase level dropped near the bottom of the HL at around [ ]a,b,c; however, 

as a result of CMT2 injection, the water level in the core began to rise at around [ ]a,b,c. 

At [ ]a,b,c CMT2 refilled as a result of condensation-induced pressure drop. 

At [ ]a,b,c, IRWST injection through DVI line 2 began, and the RPV liquid level began 

to increase. 

The RPV two-phase mixture level remained high throughout the entire test, dropping briefly 

below the bottom of the HL between [ ]a,b,c as shown in Figure C.5. This high 

liquid level can also been seen by examining the HL liquid level plots in Figure C.6. 

At approximately [ ]a,b,c, liquid began to preferentially entrain through the non-PZR side 

ADS4-1 line as shown in Figure C.7. From about [ ]a,b,c on, the flow quality in the 

non-PZR ADS4-1 line remained constant at approximately 0.05, while the PZR side ADS4-2 line 

remained constant at approximately 0.4. The core remained covered during the entire test as 

shown in Figure C.8. 
C-3 

Table C.2 Sequence of Events for Test NRC-AP1000-01 

Time (s) Event 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

a,b,c 
C-4 

Figure C.1 Vessel Wide Range Collapsed Water Level 

Figure C.2 Core Upper Half and Lower Half Collapsed Water Levels 

a,b,c 

a,b,c 
C-5 

Figure C.3 Accumulator, CMT, and IRWST Flows to DVI1 

Figure C.4 Accumulator, CMT, and IRWST Flows to DVI2 

a,b,c 

a,b,c 
C-6 

Figure C.5 Upper Plenum Two-Phase Level 

Figure C.6 Hot Leg Collapsed Water Levels 

a,b,c 

a,b,c 
C-7 

Figure C.7 ADS4 Flow Qualities 

Figure C.8 Core Heater Rod Temperatures 

a,b,c 

a,b,c 
C-8 

C.2 NRC-AP1000-02

The objectives of Test NRC-AP1000-02 were to determine system behavior during a loss of 

shutdown cooling, if the RCS would pressurize sufficiently to prevent IRWST injection, and if 

CCFL would prevent the PZR from draining. In this test, a loss of RNS cooling was simulated 

after the reactor had been shutdown for 24 hours. The RCS was at [ ]a,b,c with an initial 

PZR level of [ ]a,b,c. The ADS3 valve was opened (which simulated two trains of ADS1, 2, 

and 3 in AP1000) venting the RCS to the IRWST. The SG secondary sides were assumed to 

have been drained for maintenance. The ACCs and CMTs were assumed to be unavailable. 

One ADS4 valve was failed on demand. The facility configuration for this test is given in Table 

C.3. 

Table C.3 Facility Configuration for Test NRC-AP1000-02 

Component Configuration 

ADS1 Failed Shut 

ADS2 Failed Shut 

ADS3 Open to vent primary system 

ADS4-1 (non-PZR side) 1 out of 2 Failed Open 

ADS4-2 (PZR-side) 100% ADS4-2 flow 

ACCs Unavailable 

CMTs Unavailable 

RNS Cooling Unavailable 

Break None 

Table C.4 summarizes the sequence of events for this test. The test was started by increasing 

reactor power to 320 kW. When both HL liquid levels decreased to approximately 1⁄2-inch, one 

IRWST valve and one ADS4 valve were opened after a [ ]a,b,c time delay. ADS4B was 

opened at around [ ]a,b,c , which caused a rapid depressurization of the primary system 

and the IRWST began injecting makeup water by [ ]a,b,c as shown in Figure C.9. The 

sump recirculation valves were opened at [ ]a,b,c, as shown by a sharp increase in IRWST 

injection flow rate, and the facility transitioned into long-term cooling mode. 

As the heater rod power was increased in the beginning of the test, the RPV liquid level dropped 

significantly [ ]a,b,c as water was boiled off and released out of the open ADS3 valve 

as shown in Figure C.10. After ADS4 actuation, the RPV level increased as IRWST injection 

was provided to the core. This same trend is shown in Figure C.11, which plots the RPV mixture 

level. During the entire transient, the core never uncovered, and the mixture level was always 

above the level of the upper core plate. Figure C.14 plots the core collapsed liquid levels from 

the lower core plate to the mid-core spacer grid (LDP-109) and from the mid-core spacer grid to 

the upper-core spacer grid (LDP-110). 

Figure C.12 plots the HL levels on the PZR and non-PZR sides of the plant. At around 

[ ]a,b,c, when the HL levels dropped to 1⁄2-inch, ADS4B was opened. The HL liquid levels 
C-9 

briefly continued to drop, but quickly turned around as makeup water was provided from the 

IRWST, and the decrease in pressure caused the RPV mixture level to swell. At no time during 

the transient were the HLs emptied. 

One of the objectives of this test was to determine if the PZR would be able to drain as a result 

of CCFL. Figure C.13 plots PZR level versus time and shows that the PZR drained slowly from 

about [ ]a,b,c as primary water was boiled off and the steam exited the open ADS3 

valve. When the ADS4B valve was actuated, the PZR quickly drained and remained empty until 

about [ ]a,b,c when the plant refilled as a result of IRWST injection. Figure C.15 plots core 

heater rod temperatures, and no high rod temperature excursion was experienced during the 

entire test. 

Table C.4 Sequence of Events for Test NRC-AP1000-02 

Time (s) Event 

 

 

 

 

 

 

 

 

a,b,c 
C-10 

Figure C.9 IRWST (Lines 1 and 2) Injection Flow Rates 

Figure C.10 RPV Collapsed Liquid Level 

a,b,c 

a,b,c 
C-11 

Figure C.11 RPV Mixture Level 

Figure C.12 Hot Leg Collapsed Liquid Levels 

a,b,c 

a,b,c 
C-12 

Figure C.13 PZR Collapsed Liquid Level 

Figure C.14 Core Collapsed Liquid Levels 

a,b,c 

a,b,c 
C-13 

Figure C.15 RPV Heater Rod Temperatures 

a,b,c 
C-14 

C.3 NRC-AP1000-03

This test simulated a DEDVI DVI line 1 break with two of the four ADS4 valves failed closed 

(ADS4-2 closed and ADS4-1 opened), with the purpose of investigating a major failure of the 

ADS system. Both failed ADS4 valves were on the PZR side of the plant. Table C.5 

summarizes the facility configuration. This test is similar in intent to NRC-10, which was 

conducted for AP600 design certification. In NRC-10, failure of three of the four ADS4 valves 

produced core uncovery and heatup for a 1-inch CL break. 

Table C.5 Facility Configuration for Test NRC-AP1000-03 

Component Configuration 

ADS1 Normal actuation 

ADS2 Normal actuation 

ADS3 Normal actuation 

ADS4-1 (non-PZR side) 100% ADS4-2 flow 

ADS4-2 (PZR side) Failed Shut 

Break DEDVI Line 1 

Table C.6 summarizes the sequence of events for NRC-AP1000-03. Similar to NRC-AP1000-01, 

this DEDVI test depressurized quickly, causing a low inventory in the vessel as shown in Figure 

C.16, indicating that the break (rather than early actuation of the ADS1, 2, and 3 valves) was 

responsible for the initial rapid depressurization and mass loss. CMT1 began draining out of the 

DVI line 1 break immediately after the break opened. ADS1, 2, and 3 were opened in sequence 

from [ ]a,b,c, further depressurizing the plant. ACC2 was injecting by [ ]a,b,c, while 

ACC1 and CMT1 emptied by [ ]a,b,c. All Side 1 injection flow (ACC1, CMT1, and IRWST 

Line 1) was lost out of the DEDVI break. At [ ]a,b,c , the ADS4-1 valve was opened. 

Coupled with CMT2 and ACC2 injection, the vessel inventory increased until these two tanks 

were completely emptied by [ ]a,b,c. As shown in Figure C.19, there was a period 

[ ]a,b,c when there was no injection flow to the core. This delay in IRWST injection was also 

experienced in NRC-AP600-10 and is attributable to a known scaling distortion related to the 

reduced IRWST head, which is an artifact of the one-quarter height scaling in APEX and is 

conservative in nature. 

The RPV two-phase water level dropped below the bottom of the HL from [ ]a,b,c as shown 

in Figure C.20. IRWST injection increased from [ ]a,b,c , and the core water level 

rose to a relative maximum at approximately [ ]a,b,c. Then the vessel inventory decreased 

slowly over time, as shown in Figure C.16, as IRWST injection decreased until emptying at 

[ ]a,b,c. Sump recirculation began at about [ ]a,b,c, thus ending the IRWST injection 

phase, as shown in Figure C.21, and transitioning the plant into long-term cooling mode. 

The HL water levels are shown in Figure C.22. HL2 emptied from [ ]a,b,c and refilled from 

[ ]a,b,c while CMT2 and ACC2 injected water into the RPV. When CMT2 and 

ACC2 emptied at approximately [ ]a,b,c, HL2 was empty from [ ]a,b,c. IRWST 

injection raised the HL levels from [ ]a,b,c, but as injection flow decreased, the HL2 
C-15 

liquid level began to decrease until it emptied for the last time at approximately [ ]a,b,c. 

This is the time that IRWST injection ceased and sump recirculation began. Enough water was 

recirculated through the core that the quality out the ADS4-1 line was not unity (~0.8), as shown 

in Figure C.23. During this time, the RPV two-phase level was well below the bottom of the HL, 

as shown in Figure C.20, and a fair amount of water remained in HL1, as shown in Figure C.22. 

This behavior illustrates the importance of upper plenum and HL entrainment on passive safety 

system performance. 

Figure C.24 depicts the PZR wide range level. After the initial blowdown, substantial PZR 

injection began at approximately [ ]a,b,c and it ceased to inject water into the primary 

system at [ ]a,b,c. 

Several small heater rod temperature spikes are shown in Figure C.25. The first spike at 

around [ ]a,b,c corresponded to the lowest vessel level experienced during the test, which 

is shown Figure C.16. This spike occurred when all CMTs and ACCs were empty before 

IRWST injection flow began. The next significant temperature spike occurs at around 

[ ]a,b,c when the core water level was low during sump recirculation. However, even 

though the vessel level was low at certain times during the test, the core still remained covered. 

Consequently, these temperature spikes were most likely caused by relatively large voids in the 

core coming into contact with specific thermocouples and do not represent an “average” core 

temperature, but local “warm spots”. 
C-16 

Table C.6 Sequence of Events for Test NRC-AP1000-03 

Time (s) Event 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

a,b,c 
C-17 

Figure C.16 Vessel Wide Range Collapsed Water Level 

Figure C.17 Core Upper Half and Lower Half Collapsed Water Levels 

a,b,c 

a,b,c 
C-18 

Figure C.18 Accumulator, CMT, and IRWST Flows to DVI1 

Figure C.19 Accumulator, CMT, and IRWST Flows to DVI2 

a,b,c 

a,b,c 
C-19 

Figure C.20 RPV Two-Phase Level 

Figure C.21 Sump Recirculation Flows 

a,b,c 

a,b,c 
C-20 

Figure C.22 Hot Leg Collapsed Water Levels 

Figure C.23 ADS4-1 (Non-Pressurizer Side) Flow Quality 

a,b,c 

a,b,c 
C-21 

Figure C.24 Pressurizer Wide Range Collapsed Water Level 

Figure C.25 Core Heater Rod Temperatures 

a,b,c 

a,b,c 
C-22 

C.4 NRC-AP1000-04

Test NRC-AP1000-04 was a beyond-DBA test simulating a 1-inch CL break with one ADS4-1 

valve failed shut (50% line flow) on the non-PZR side of the plant. In addition, CMT1, ACC1, 

sump recirculation line 1, and IRWST injection line 1 were unavailable. Table C.7 summarizes 

the facility configuration for this test. Upon reaching 30 minutes of steady-state long-term 

cooling, the primary sump drain valve was opened to simulate a degraded containment sump. 

This test was similar in intent to NRC-AP600-27, where a similar break size with a degraded 

sump was simulated, in part, to determine the impact on long-term cooling as a result of 

depleting the primary sump inventory. In NRC-AP1000-04 and both AP600 tests, the two-phase 

level in the core remained above the heater rods. 

Table C.7 Facility Configuration for Test NRC-AP1000-04

Component Configuration 

ADS1 Normal actuation 

ADS2 Normal actuation 

ADS3 Normal actuation 

ADS4-1 (non-PZR side) 50% ADS4-1 flow 

ADS4-2 (PZR side) 100% ADS4-2 flow 

CMT1 Unavailable 

ACC1 Unavailable 

Sump Recirculation Line 1 Unavailable 

IRWST Injection Line 1 Unavailable 

Break 1-inch bottom of CL 4 

Table C.8 summarizes the sequence of events for NRC-AP1000-04. The break was opened at 

zero seconds, causing the CMTs to inject immediately and, during the initial blowdown, the RPV 

vessel inventory dropped to a relative minimum at around [ ]a,b,c as shown in Figures C.26 

and C.27. The ACC2 injection began at [ ]a,b,c as shown in Figure C.28. The RPV 

inventory remained relatively low until around [ ]a,b,c IRWST injection began, and the 

inventory in the RPV increased to a relative maximum at approximately [ ]a,b,c. This was 

around the same time IRWST injection reached maximum flow as shown in Figure C.28. 

At [ ]a,b,c the primary sump drain valve was opened to simulate a degraded sump (ending 

the IRWST injection phase of the transient), which created an increase in IRWST flow shown in 

Figure C.28. Primary sump overflow occurred at approximately [ ]a,b,c and water was 

recirculated from the sump, to the IRWST, through the DVI2 line and into the RPV. 

From [ ]a,b,c the two-phase RPV level never dropped below the bottom of the HL and 

remained high until test termination at [ ]a,b,c as shown in Figure C.29. The HL liquid 

levels also remained high after the initial plant blowdown as shown in Figure C.30. 
C-23 

High HL and two-phase RPV levels provided an increased amount of water for liquid 

entrainment out of the ADS4 valves. In Figure C.31, the quality out of ADS4-1 was very low 

(< 0.04) compared to ADS4-2 quality ([ ]a,b,c), illustrating a preference to entrain liquid 

through the non-PZR side HL. This same preference was also observed in Test NRC-AP1000-

01. This preferential entrainment to one side of the plant was unique, but did not raise any 

safety concerns and this phenomenon is currently being investigated in conjunction with 

thermal-hydraulic code development and assessment activities. 

The PZR began to drain at [ ]a,b,c and emptied by [ ]a,b,c as shown in Figure C.32. 

Adequate injection was provided by CMT2, ACC2, and the PZR during the initial blowdown to 

keep the heater rods covered. A 1-inch CL break along with a degraded sump did not challenge 

the passive safety systems enough to uncover the heater rods. The core remained adequately 

cooled during the entire test, as shown in Figure C.34. 
C-24 

Table C.8 Sequence of Events for Test NRC-AP1000-04 

Time (s) Event 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

a,b,c 
C-25 

Figure C.26 Vessel Wide Range Collapsed Water Level 

Figure C.27 Core Upper Half and lower Half Collapsed Water Levels 

a,b,c 

a,b,c 
C-26 

Figure C.28 Accumulator, CMT, and IRWST Flows to DVI2 

Figure C.29 Upper Plenum Two-Phase Level 

a,b,c 

a,b,c 
C-27 

Figure C.30 Hot Leg Collapsed Water Levels 

Figure C.31 ADS4 Flow Qualities 

a,b,c 

a,b,c 
C-28 

Figure C.32 Pressurizer Wide Range Collapsed Water Level 

Figure C.33 Core Heater Rod Temperatures 

a,b,c 

a,b,c 
C-29 

C.5 NRC-AP1000-05

 

This simulated a double-ended guillotine break of a DEDVI with two of the four ADS4 valves 

failed closed (ADS4-1 closed and ADS4-2 opened). Both ADS4 valves on the non-PZR side 

were failed closed. Table C.9 summarizes the facility configuration. This test compliments Test 

NRC-AP1000-03, in which the only difference was in the location of the valve failure. In NRCAP1000-03, both failed valves were on the PZR side. 

Table C.9 Facility Configuration for Test NRC-AP1000-05 

Component Configuration 

ADS1 Normal actuation 

ADS2 Normal actuation 

ADS3 Normal actuation 

ADS4-1 (non-PZR side) Failed shut 

ADS4-2 (PZR side) 100% ADS4-2 flow 

Break DEDVI Line 1 

Table C.10 summarizes the sequence of events for this test. When the break opened at zero 

seconds, CMT1, CMT2, and ACC2 immediately began to discharge out the break. 

By [ ]a,b,c, CMT1, ACC1, and ACC2 were completely empty. Figures C.36 and C.37 

depict the flow from the ACCs, CMTs, and IRWST. The vessel inventory was low from 

approximately [ ]a,b,c. 

The vessel inventory increased slightly, as shown in Figure C.38, as CMT2 injection was 

prolonged, but after CMT2 emptied at [ ]a,b,c, no injection water was provided to the core. 

The RPV water level in the core began to drop at approximately [ ]a,b,c into the test. 

With no flow from the IRWST, HL1 emptied at [ ]a,b,c while water continued to be 

entrained out the ADS4-2 line, as shown in Figure C.40. A significant water level in HL2 was 

maintained until [ ]a,b,c. The flow quality in ADS4-2 throughout most of the test remained 

fairly low (~0.5), even when the RPV two-phase level dropped below the bottom of the HLs. 

The minimum level that the two-phase probe can measure is approximately 2 inches above the 

upper core plate. Thus, the LT-120 measurement could not decrease after it “bottomed” at 

[ ]a,b,c. 

By [ ]a,b,c the PZR had still not drained due to CCFL, as shown in Figure C.41, and the 

heater rods began to uncover. At [ ]a,b,c a heater rod temperature excursion was 

observed, as shown in Figure 4.9, and the heater rod power was shut off at [ ]a,b,c. 

Immediately after the heater rods were turned off, IRWST injection flow began. 

The results of this test illustrate the sensitivity of ADS4 valve failure on core coolability. In this 

test, IRWST injection never began through DVI Line 2 as water was lost out ADS4-2 as a result 

of upper head and HL liquid entrainment. 
C-30 

Table C.10 Sequence of Events for Test NRC-AP1000-05 

Time (s) Event 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

a,b,c 
C-31 

Figure C.34 Vessel Wide Range Collapsed Water Level 

Figure C.35 Core Upper Half and Lower Half Collapsed Water Levels 

a,b,c 

a,b,c 
C-32 

Figure C.36 Accumulator, CMT, and IRWST Flows to DVI1 

Figure C.37 Accumulator, CMT, and IRWST Flows to DVI2 

a,b,c 

a,b,c 
C-33 

Figure C.38 Upper Plenum Two-Phase Level 

Figure C.39 Hot Leg Collapsed Water Levels 

a,b,c 

a,b,c 
C-34 

Figure C.40 ADS4-2 (Pressurizer Side) Flow Quality 

Figure C.41 Pressurizer Wide Range Collapsed Water Level 

a,b,c 

a,b,c 
C-35 

Figure C.42 Core Heater Rod Temperatures 

a,b,c 
C-36 

C.6 NRC-AP1000-06

The purpose of this test was to investigate a significant failure of the ADS4 system for a small 

CL break. It is similar in intent to Test NRC-10, which was conducted for AP600 design 

certification. In NRC-10, a 1-inch CL break was simulated with three of the four ADS4 valves 

failed closed. The AP600 test indicated that a failure of three of the four ADS4 valves to open 

resulted in a significant delay in the start of IRWST injection. The liquid level dropped below 

the top of the core, and a temperature excursion was observed. This result did not agree with 

those of the ROSA counterpart, Test AP- CL-07. The difference in the APEX and ROSA results 

was attributed to the reduced elevation head available for IRWST injection in APEX. 

Test NRC-AP1000-06 simulated a 2-inch equivalent diameter break at the bottom of CL4. This 

was a beyond-DBA test, with two of the four ADS4 valves failed closed. Table C.11 summarizes 

the facility configuration. Both of the failed valves were on the ADS4-1 line, which connects to 

the HL without the PZR. The PRHR HX was inactive during this test. 

Table C.11 Facility Configuration for Test NRC-AP1000-06 

Component Configuration 

ADS1 Normal actuation 

ADS2 Normal actuation 

ADS3 Normal actuation 

ADS4-1 (non-PZR side) Failed shut 

ADS4-2 (PZR side) 100% ADS4-2 flow 

Break 2-inch bottom of CL 

Test NRC-AP1000-06 investigated a larger break size than NRC-10, but failed two of the four 

ADS4 valves. Similar to NRC-10, the water level dropped below the top of the core, and a 

temperature excursion occurred. Table C.12 summarizes the sequence of events for 

NRC-AP1000-06. Figure C.43 shows the vessel wide range collapsed water level, which 

remained above the top of the core until shortly after [ ]a,b,c, which is when the ACCs 

emptied. As the ACC flows decreased, the collapsed water level in the top half of the core 

decreased rapidly, as shown in Figure C.44. Figures C.45 and C.46 depict the flow from the 

ACCs, CMTs, and IRWST. 

The CMTs both emptied by [ ]a,b,c. The system had not depressurized sufficiently to 

begin IRWST injection, so from this time until the test was terminated at [ ]a,b,c, there was 

no injection flow to the vessel. The two-phase level, shown in Figure C.47, remained above the 

bottom of the HLs until approximately [ ]a,b,c, as indicated by the two-phase level probe 

(LT-120). The two-phase level continued to decrease, and the upper plenum emptied at 

[ ]a,b,c (based on LDP-112). The minimum level that the two-phase probe can measure is 

approximately 2 inches above the upper core plate. Thus, the LT-120 measurement could not 

decrease after it “bottomed” at [ ]a,b,c. 
C-37 

Figure C.48 depicts the HL levels. Of particular interest, is the difference in levels at 

approximately [ ]a,b,c. The HL on the non-PZR side (HL1) emptied shortly before the twophase level dropped below the bottom of the HLs. A level in the PZR side HL (HL2) remained 

even after the upper plenum emptied at [ ]a,b,c. Figure C.49 depicts the PZR water level, 

which remained high following ADS4 opening at [ ]a,b,c. 

Flow quality to the ADS4-2 line, is low (x < 0.1) immediately following valve opening at 

[ ]a,b,c as shown in Figure C.49. The quality increased rapidly from approximately 0.1 

to 0.4 shortly after accumulator injection stopped, and then increased again after the CMTs 

emptied. The quality increased from approximately 0.8 as the two-phase level dropped below 

the bottom of the HL, and remained less than 0.9 even after the upper plenum became empty 

at [ ]a,b,c. 

At [ ]a,b,c, a core temperature excursion was observed, as shown in Figure C.51. 

The core power was manually scrammed. IRWST injection began immediately after to prevent 

heater rod damage. 
C-38 

Table C.12 Sequence of Events for Test NRC-AP1000-06 

Time (s) Events 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

a,b,c 
C-39 

Figure C.43 Vessel Wide Range Collapsed Water Level 

Figure C.44 Core Upper Half and Lower Half Collapsed Water Levels 

a,b,c 

a,b,c 
C-40 

Figure C.45. Accumulator, CMT, and IRWST Flows to DVI1 

FigureC.46 Accumulator, CMT, and IRWST Flows to DVI2 

a,b,c 

a,b,c 
C-41 

Figure C.47 Upper Plenum Two-Phase Level from Level Transmitter 

Figure C.48 Hot Leg Collapsed Water Levels 

a,b,c 

a,b,c 
C-42 

Figure C.49 ADS4-2 Flow Quality 

Figure C.50 Pressurizer Wide Range Collapsed Water Level 

a,b,c 

a,b,c 
C-43 

Figure C.51 Core Heater Rod Temperatures 

a,b,c 
C-44 

C.7 NRC-AP1000-10

The objective of this test was to obtain thermal-hydraulic data for PRA verification of a simulated 

AP1000 2-inch CL break off the bottom of CL4. The test was initiated from normal operating 

temperature and pressure. For this test, ADS1-3 and one of the four ADS4 valves failed to 

open on the non-PZR side of the plant. In addition, a [ ]a,b,c sharp edge orifice was 

installed in each SG PORV line. The SG PORVs were allowed to cycle between [ ]a,b,c

(open) and [ ]a,b,c (closed) to help remove decay heat during the initial blowdown stage of 

the transient. The IRWST sump recirculation set point was changed from [ ]a,b,c. 

The test was terminated after 30 minutes of steady-state long-term cooling. Table C.13 shows 

the facility configuration for this test. 

Table C.13 Facility Configuration for Test NRC-AP1000-10 

Component Configuration 

ADS1 Failed shut 

ADS2 Failed shut 

ADS3 Failed shut 

ADS4-1 (non-PZR side) 50% ADS4-1 flow 

ADS4-2 (PZR side) 100% ADS4-2 flow 

Break 2-inch bottom of CL 4 

Table C.14 summarizes the sequence of events for this test. The break was initiated at zero 

seconds, and the CMTs immediately began to drain, as shown in Figures C.52 and C.53. 

The ACCs began to inject around [ ]a,b,c and, shortly afterward at [ ]a,b,c , the ADS4 

valves opened. This caused the plant to further depressurize and increased ACC injection flow 

rates considerably, as shown in Figures C.52 and C.53. 

Between [ ]a,b,c , there was a noticeable oscillation of IRWST injection flow rate 

(Figures C.52 and C.53), core liquid level (Figure C.54), and ADS4 line flow qualities (Figure C.59). 

The RPV pressure was very close to atmospheric by [ ]a,b,c (Figure C.58), and the HLs 

were almost completely flooded (Figure C.56). This oscillatory behavior is consistent with 

separate effects experiments run in ATLATS, which investigated the intermittent nature of liquid 

entrainment from the HLs into the vertical ADS4 lines. 

The RPV liquid level reached a maximum around [ ]a,b,c , and then began to slowly 

decrease as the IRWST injection flow rate decreased. Oscillations were experienced again 

from [ ]a,b,c, because the core exit flow returned to saturated conditions as IRWST 

injection flow was considerably decreased. These RSOs are consistent with previous AP600 

testing, and are described in Section 4.1.4 of this report. The plant entered long-term cooling at 

[ ]a,b,c when the sump recirculation lines were opened. Figure C.60 plots the RPV heater 

rod temperature and shows that the core remained adequately cooled during the entire test. 
C-45 

Table C.14 Sequence of Events for Test NRC-AP1000-10 

Time (s) Events 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

a,b,c 
C-46 

Figure C.52 Side 1 Injection Flow Rates (ACC1, CMT1, and IRWST Line 1) 

Figure C. 53 Side 2 Injection Flow Rates (ACC2, CMT2, and IRWST Line 2) 

a,b,c 

a,b,c 
C-47 

Figure C.54 RPV Collapsed Liquid Level 

Figure C.55 RPV Mixture Level 

a,b,c 

a,b,c 
C-48 

Figure C. 56 Hot Leg Collapsed Liquid Levels 

Figure C.57 Core Collapsed Liquid Levels 

a,b,c 

a,b,c 
C-49 

Figure C.58 PZR Wide Range Collapsed Liquid Level 

Figure C.59 ADS4 Flow Qualities 

a,b,c 

a,b,c 
C-50 

Figure C.60 RPV Heater Rod Temperatures 

a,b,c 
C-51 

C.8 NRC-AP1000-11

The objectives of this test were to (1) evaluate long-term PRHR HX operation including heatup 

of the IRWST to saturated temperature, (2) determine the effect of system behavior caused by 

ADS1-3 discharge to the saturated liquid inside the IRWST, and (3) determine the effects on 

long-term cooling caused by the injection of saturated liquid into the RPV from the IRWST. 

The control logic was modified to simulate station blackout (loss of all AC power). The core 

was operated at reduced power (600 kW) and with a modified decay power curve to preserve 

integrated energy. The SG PORVs were set to cycle between [ ]a,b,c. When the 

SG wide range level decreased to [ ]a,b,c, an “S” signal was generated to open CMT 

injection valves and the PRHR HX outlet valve. The ADS was actuated [ ]a,b,c after 

initiation of the “S” signal. Table C.15 summarizes the facility configuration for this test. 

Table C.15 Facility Configuration for Test NRC-AP1000-11 

Component Configuration 

ADS1 Normal operation 

ADS2 Normal operation 

ADS3 Normal operation 

ADS4-1 (non-PZR side) 50% ADS4-1 flow 

ADS4-2 (PZR side) 100% ADS4-2 flow 

Break None 

Table C.16 summarizes he sequence of events for this test. The main feed pump was tripped 

at zero seconds, and the RCP coastdown command was given at 1 sec. For the next 

[ ]a,b,c, the SG PORVs cycled open and closed. The plant slowly depressurized for about 

[ ]a,b,c as shown in Figure C.61. The CMTs began to inject at a low flow rate around 

[ ]a,b,c (Figures C.62 and 63). 

At [ ]a,b,c ADS1-3 stages began to open, causing a rapid depressurization of the plant 

(Figure C.61). The ACCs began injecting about [ ]a,b,c later, and CMT injection flow 

increased significantly. Another jump in the ACC injection flow rate occurred at [ ]a,b,c, 

after ADS4 valve actuation. IRWST injection began by [ ]a,b,c , the CMTs and 

ACCs were emptied, illustrating about a 10-minute overlap in passive safety injection flow. 

The RPV collapsed liquid level remained high until ADS actuation, when it dropped just below 

the top of the core, as shown in Figure C.64. However, the two-phase mixture level remained 

well above the bottom of the HL throughout the entire test (Figure C.65). 

As the SG PORVs cycled open and closed between [ ]a,b,c, removing decay heat 

from the core, the RPV collapsed liquid level decreased (Figure C.64), while the PZR level 

increased, as shown in Figure C.67, as steam boiled off in the core filled the PZR. At about 

[ ]a,b,c , the PZR was completely filled (Figure C.67) and the primary system began to 
C-52 

pressurize (Figure C.61). At this time, the primary system went “solid” and the RPV liquid level 

remained constant (Figures C.64 and 65). 

The last phase of the transient, sump recirculation cooling, began upon reaching a low level in 

the IRWST corresponding to the containment sump flood-up elevation. The PRHR HX and 

CMT operation were capable of removing core decay energy (Figure C.69) during the prolonged 

operation expected for a station blackout. The ADS1-4 blowdown was not affected by the 

saturated conditions inside the IRWST. 

Table C.14 Sequence of Events for Test NRC-AP1000-11 

Time (s) Events 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

a,b,c 
C-53 

Figure C.61 RPV Head Pressure 

Figure C.62 Side 1 Safety Injection Flow (ACC1, CMT1, and IRWST Line 1) 

a,b,c 

a,b,c 
C-54 

Figure C.63 Side 2 Safety Injection Flow (ACC2, CMT2, IRWST Line 2) 

Figure C. 64 RPV Collapsed Liquid Level 

a,b,c 

a,b,c 
C-55 

Figure C.65 RPV Mixture Level 

Figure C.66 Hot Leg Collapsed Liquid Levels 

a,b,c 

a,b,c 
C-56 

Figure C.67 PZR Wide Range Collapsed Liquid Level 

Figure C.68 Core Collapsed Liquid Levels 

a,b,c 

a,b,c 
C-57 

Figure C.69 RPV Heater Rod Temperatures 

a,b,c