BWR ECCS Pump 

Suction Concerns 

following a LOCA

Office of Nuclear Regulatory Research

NUREG-2196

BWR ECCS Pump 

Suction Concerns 

following a LOCA

Manuscript Completed: February 2016

Date Published: May 2016

Prepared by: 

W.J. Krotiuk, C. Boyd, T. Zaki, W. Wang

Office of Nuclear Regulatory Research

NUREG-2196

iii

ABSTRACT

This document describes a method to perform a plant-specific assessment to determine 

whether an emergency core cooling system (ECCS) pump will operate acceptably following a 

loss-of-coolant accident (LOCA) in a boiling water reactor (BWR) pressure suppression pool. 

The assessment approach uses the ECCS pump strainer location, the ECCS pump operation 

criteria, the ECCS pump startup and full operation timing information, and an estimate of the 

transient pool void fraction information to make the determination.

To support the developed assessment method, this report describes the acceptable ECCS 

pump operating range necessary to prevent pump damage if noncondensable gas ingestion 

occurs or to permit recovery from noncondensable gas ingestion. Additionally, this report uses 

computational fluid dynamics (CFD) analyses to define a noncondensable gas bubble 

“exclusion zone” and to provide the time-dependent noncondensable void fraction distribution in 

the suppression pool outside the large gas bubble following a LOCA. The report also lists the 

test facilities used to test BWR pressure suppression systems and summarizes the available 

test data that are used to verify the acceptability of the CFD analyses.

v 

TABLE OF CONTENTS

ABSTRACT...............................................................................................................................iii

TABLE OF CONTENTS ............................................................................................................. v

LIST OF FIGURES ...................................................................................................................vii

LIST OF TABLES......................................................................................................................ix

EXECUTIVE SUMMARY ........................................................................................................... xi 

ACRONYMS AND ABBREVIATIONS................................................................................... xxiii

1. INTRODUCTION...................................................................................................................1

2. BWR POST-LOCA CONTAINMENT ANALYSIS AND SUPPRESSION POOL

BEHAVIOR ...........................................................................................................................3 

2.1 Containment Analysis .................................................................................................... 3 

2.1.1 Vent Discharge Flow........................................................................................... 5 

2.2 Post-LOCA Suppression Pool Behavior......................................................................... 9 

2.2.1 Downcomer Vent Water Clearing........................................................................ 9 

2.2.2 Noncondensable Gas Flow and Bubble Behavior ............................................... 9 

2.2.3 Condensation Oscillation and Chugging............................................................10

3. LITERATURE SURVEY......................................................................................................13

3.1 Pressure Suppression Experimental Facilities and Test Data ...................................... 13 

3.1.1 General Electric Pressure Suppression Testing ................................................ 13 

3.1.2 NRC 1/5 Scale Mark I Pressure Suppression Testing ....................................... 14 

3.1.3 Massachusetts Institute of Technology Mark I Blowdown Tests ........................ 14 

3.1.4 University of California Mark I Scaled Suppression Pool Tests ......................... 14 

3.1.5 Lappeenranta University Scaled Suppression Pool Experiments ...................... 14 

3.1.6 Nordic Owners Group Testing for Swedish BWRs ............................................ 18 

3.1.7 Purdue University Mark I Scaled Air and Air/Steam Suppression Pool 

Tests ................................................................................................................20

3.1.8 GKSS Pressure Suppression Research Program .............................................21

3.1.9 Marviken Test Facility Testing...........................................................................24

3.1.10 Paul Scherrer Institute Tests ............................................................................. 24 

3.1.11 Alden Research Laboratory Testing ..................................................................25

4. ECCS PUMP OPERATION................................................................................................. 29

4.1 ECCS Pump Timing Considerations ............................................................................ 29 

4.2 Acceptable ECCS Pump Operation ............................................................................. 29 

4.2.1 ECCS Pump Performance in BWRs (NUREG/CR-2772)....................................30

4.2.2 Pump Performance under Air Ingesting Conditions (NUREG/CR-2792).............30

4.2.3 Pump Operation and Acceptable gas Quantity (Generic Letter 2007-02) ...........31

4.2.4 Lappeenranta University Pump Experiments with Noncondensable Gas ...........31

4.2.5 BWR Owners’ Group ECCS Pump Suction Void Fraction Study ........................32

4.2.6 University of Minnesota Centrifugal Pump Performance Tests with Air 

Ingestion ............................................................................................................32

4.2.7 Expert Panel Study of Gas Accumulation on Pumps – 

 “The Pump Roadmap”........................................................................................33
vi

TABLE OF CONTENTS

5. POST-LOCA SUPPRESSION POOL ANALYSIS...............................................................35

5.1 Scaling of Test Data to Plant Conditions and Geometry...............................................35

5.1.1 Scaling Approach using Experiment Modeling ...................................................35

5.1.2 Scaling Approach using Downward Gas Jet Behavior........................................36

5.1.3 Assessment of Scaling Approached Using CFD Analysis...................................37

5.1.4 Conclusions Regarding Attempt to Scale Test Data to Plant Geometry.............. 38 

5.2 CFD Analysis of Suppression Pool ..............................................................................38

5.2.1 PPOOLEX Tests ................................................................................................40

5.2.2 PUMA Tests.......................................................................................................48

5.2.3 Full-Scale Simulations........................................................................................57

6. ASSESSMENT APPROACH FOR ECCS PUMP OPERATION IN BWR PRESSURE

SUPPRESSION CONTAINMENT SYSTEMS .....................................................................69

6.1 Post-LOCA Suppression Pool Behavior.......................................................................70

6.1.1 Downcomer Vent Water Clearing.......................................................................70

6.1.2 Noncondensable Gas Flow and Bubble Behavior ..............................................70

6.1.3 Condensation Oscillation and Chugging.............................................................71

6.2 Other Considerations Regarding Suppression Pool Conditions ...................................71

6.2.1 Consideration of Dissolved Noncondensable Gas in the Suppression Pool .......71

6.2.2 Consideration of Noncondensable Gas Transport to and Accumulation at 

the ECCS Pump Inlet before Pump Startup .......................................................73

6.2.3 Consideration on Noncondensable Gas Transport in the Pump Inlet Piping

 during Pump Operation ...................................................................................... 75 

6.3 Determination of ECCS Pump Strainer Location in a Suppression Pool....................... 75 

6.4 ECCS Pump Operation................................................................................................75

6.4.1 ECCS Pump Timing Considerations ..................................................................75

6.4.2 ECCS Pump Acceptable Operation Criteria ....................................................... 76 

6.5 Procedure for Performing a GI-193 Plant Assessment................................................. 77 

6.6 Possible Corrective Actions .........................................................................................78

6.7 Sample Calculation for Performing a GI-193 Plant Assessment................................... 81 

7. SUMMARY AND CONCLUSIONS......................................................................................85

7.1 Determination of Suppression Pool Exclusion Zone and Void Fraction 

Distribution Assessment Method.................................................................................. 85 

7.2 Acceptable ECCS Pump Operation Criteria ................................................................. 85 

8. REFERENCES....................................................................................................................87
vii

LIST OF FIGURES

2.1-1 BWR Pressure Suppression System Designs............................................................... 3 

2.1-2 Cross Section and Dimensions of Typical (a) Mark I, (b) Mark II and (c) Mark III 

Suppression Chambers (Dimensions in meter (foot)) ................................................... 4 

2.1-3 Typical ECCS Pump Strainer Locations in (a) Mark I, (b) Mark II and (c) Mark III 

Suppression Chambers (Dimensions in meter (foot)) ................................................... 4 

2.1-4 Predictions of Mark I Containment Drywell (DW) and Wetwell (WW) Pressures 

following a DBA LOCA ................................................................................................. 6 

2.1-5 Predictions of Mark I Containment Drywell (DW), Wetwell (WW) and Suppression 

Pool (SP) Temperatures following a DBA LOCA .......................................................... 6 

2.1-6 MELCOR Predictions of Mark I Post-DBA LOCA Constituents of Vent System 

Entrance Flow .............................................................................................................. 7 

2.1-7 MELCOR Predictions of Mark I Post-DBA LOCA Constituents of Downcomer

Exit Flow....................................................................................................................... 8 

2.1-8 MELCOR Predictions of Mark I Post-DBA LOCA Noncondensable Gas Mass 

Fraction of Downcomer Exit Flow ................................................................................. 8 

2.2-1 Raleigh Gas Bubble Oscillation in Liquid (Ref. 3) .......................................................10

2.2-2 Schematic of Typical Condensation Mode Regions during LOCA 

Blowdown (Ref. 3)......................................................................................................11

2.2-3 Steam Condensation Oscillation at Vent Exit (Ref. 3).................................................11

2.2-4 Sampling Chugging Map for Vertical Vents (Ref. 3)....................................................12

3.1-1 Schematic of the Olkiluoto Type BWR Pressure Suppression 

Containment (Ref. 31) ................................................................................................ 15 

3.1-2 POOLEX (left) and PPOOLEX (right) Test Facilities (Ref. 31) .................................... 17 

3.1-3 Vent Air Bubble Comparisons between CFD Calculation and PPOOLEX 

Experiment (Ref. 29) .................................................................................................. 18 

3.1-4 Video Frames from a GKSS Full-Scale Test of Mark II Suppression Pool

Behavior following a LOCA.........................................................................................22

5.1-1 Model for Downward Gas Jet in Liquid .......................................................................37

5.2.1-1 Overview of PPOOLEX Test Facility...........................................................................40

5.2.1-2 CFD PPOOLEX Geometry .........................................................................................42

5.2.1-3 PPOOLEX Computational Mesh.................................................................................42

5.2.1-4 Mass Flow at Inlet for PPOOLEX Tests......................................................................43

5.2.1-5 Comparison of Predicted Bubble Size and Location for Test 09-01 ............................44

5.2.1-6 Comparisons of Predicted and Measured Pressures, Test 09-01 ...............................45

5.2.1-7 Comparisons of Predicted Bubble Size and Location for Test 09-03 ..........................46

5.2.1-8 Comparisons of Predicted and Measured Pressures, Test 09-03 ...............................47

5.2.2-1 Overview of PUMA Vessel Used for Tests..................................................................48

5.2.2-2 Video Frames Showing Downcomer Pipe and Instrumentation Supports ................... 49 

5.2.2-3 CFD PUMA Geometry ................................................................................................50

5.2.2-4 PUMA Computational Mesh .......................................................................................50

5.2.2-5 Mass Flows at the Inlet Pipe Boundary for PUMA Tests.............................................51

5.2.2-6 Comparison of Test Video and CFD Predictions for Test A1 ......................................52

5.2.2-7 Comparison of Test Video and CFD Predictions for Test A2 ......................................53
viii

LIST OF FIGURES

5.2.2-8 Comparison of Test Video and CFD Predictions for Test A3 ......................................53

5.2.2-9 Penetration Depth vs Time Comparisons for PUMA Tests..........................................55

5.2.3-1 Generic BWR Mark I Containment Layout ..................................................................57

5.2.3-2 CFD Torus Section Model .......................................................................................... 59 

5.2.3-3 CFD Torus Section Mesh ........................................................................................... 59 

5.2.3-4 Mass Flows at Inlet for Full-Scale Simulation..............................................................60

5.2.3-5 Contour Plots of Time-Dependent Bubble Interface (VOF Method) ............................61

5.2.3-6 Contour Plots of Time-Dependent Void Fraction (Eulerian Method)............................63

5.2.3-7 Time-Dependent Maximum Penetration Depth...........................................................64

5.2.3-8 Time-Dependent Radial Width of Main Bubble ...........................................................65

5.2.3-9 Average Pool Void Fraction below the Downcomer Exit Using the VOF and 

Euler Methods ............................................................................................................66

6.1-1 Downcomer Flow/Suppression Pool Conditions and ECCS Pump Operation 

Comparison................................................................................................................70

6.5-1 Approach for Assessing Effects of Noncondensable Gas on ECCS Pump 

Operation ...................................................................................................................80

6.7-1 Plot of Suppression Pool Void Fraction Sample Cases for GI-193 Assessment.......... 82 

6.7-2 Log Plot of Suppression Pool Void Fraction Sample Cases for GI-193 

Assessment ................................................................................................................ 82 
ix

LIST OF TABLES

2.1-1 Post-LOCA Downcomer Flow Phases after Vent Water Clearing ................................. 5 

3.1-1 Pressure Suppression Test Facilities.......................................................................... 26 

4.1-1 Allowable Average Noncondensable Void Fractions to Preclude Damage to 

Typical BWR Pumps...................................................................................................33

5.2.1-1 CFD Model Basic Solver Settings...............................................................................41

5.2.1-2 Boundary and Initial Conditions for PPOOLEX Predictions.........................................43

5.2.2-1 Boundary and Initial Conditions for PUMA Models .....................................................51

5.2.3-1 Boundary and initial Conditions for Torus Model.........................................................60

5.2.3-2 Eulerian Multi-Phase CFD Model Basic Solver Settings .............................................62

6.7-1 Suppression Pool Void Fraction Sample Cases for GI-193 Assessment..................... 81 

xi

EXECUTIVE SUMMARY

This report summarizes the work and recommendations developed to support the technical 

basis for resolving U.S. Nuclear Regulatory Commission (NRC) Generic Safety Issue (GI) 193, 

“Boiling Water Reactor (BWR) Emergency Core Cooling System (ECCS) Suction Concerns.” 

The following text from Reference 1 provides the background for this activity: 

GSI-193 addresses the possible failure of the ECCS pumps due to unanticipated, 

large quantities of entrained gas in the suction piping from BWR suppression 

pools. The issue applies to Mark I, II, and III containments during large- and 

medium-break LOCAs, and could potentially cause pump failure or degraded 

performance due to gas binding, vapor locking, or cavitation.

BWRs in the United States use containment pressure suppression (PS) systems to minimize the 

containment pressure rise following a loss of coolant accident (LOCA). The containment 

systems include a drywell that contains the reactor vessel and piping, a suppression chamber or 

wetwell that contains a gas atmosphere and a pool of water to provide the pressure suppression 

function, and a downcomer vent system that connects the drywell to a submerged position in 

the wetwell suppression pool. Following a LOCA the drywell atmosphere and the break mass 

flows through the downcomer vent system and enters the suppression pool. The pool 

condenses the break steam flow and removes energy from any noncondensable gas (NCG).

This report describes the technical specifics of the GI-193 problem and summarizes the 

assessment of two analytical approaches, a scaling method and computational fluid dynamics 

(CFD) analyses, to support the development of the technical basis for developing criteria to 

perform a plant assessment. 

Section 3 lists the facilities used to test BWR pressure suppression systems and summarizes 

the test data available from these facilities, which are applicable to the GI-193 activity. The 

primary tests and facilities which provide data specific to the GI-193 activities are the tests 

performed at Lappeenranta University and Purdue University.

Section 4 of this report describes the acceptable ECCS pump operating range necessary to 

prevent pump damage if NCG ingestion occurs or to permit recovery from condensable gas 

ingestion.

Section 5 of this report summarizes the CFD analyses used to define an NCG large bubble 

“exclusion zone” and to provide the time dependent noncondensable void fraction distribution in 

the suppression pool outside the large gas bubble following a LOCA. 

Section 6 describes a method to perform a plant-specific assessment to determine whether an 

ECCS pump will operate acceptably following a design-basis accident (DBA) LOCA in a Mark I 

suppression pool using the analyses results and criteria presented in this report. The 

assessment approach uses the ECCS pump strainer location, the ECCS pump operation 

criteria, the ECCS pump startup and full operation timing information, and the pool void fraction 

information to make the determination.
xii

Post-LOCA Suppression Pool Behavior

Following a LOCA, the downcomer vent flow into the suppression pool can be divided into four

different phases: (1) the vent water clearing phase, (2) the large noncondensable gas flow and 

small noncondensable gas bubble phase, and (3) the steam condensation oscillation and 

chugging phase.

1. Vent water clearing—A water column exists in the downcomer vent before the start of a

LOCA. As the drywell pressure increases following a LOCA the water column is pushed

into the suppression pool because of the increased pressure differential between the

drywell and wetwell. The water clearing phase lasts about 0.3 seconds for a double-end

DBA LOCA, but can take longer for smaller break sizes.

2. Noncondensable gas flow and small gas bubble phase—The noncondensable gas flow

behavior can be divided into two phases, a large gas bubble phase followed by a longer

small gas bubble phase. After the completion of vent water clearing, the NCG-steam

mixture in the drywell begins to flow out of the vent and into the suppression pool. The

initial flow in this phase contains a significant amount of NCG. This initial large gas “jet”

can penetrate a significant depth into the pool as it follows the initial water jet downward.

The amount of NCG in the vent flow decreases quickly and reaches very small levels

after 2–3 seconds for a DBA LOCA. The size of the initial large NCG “jet” can define an

“exclusion zone” volume at the vent exit which possesses a large fraction of

noncondensable gas. The ECCS strainer and pump suction should not extend into the

“exclusion zone” during ECCS pump operation. After 2–3 seconds the size of the NCG

volume decreases as the initial large gas bubble rises and exits the top of the

suppression pool. After the large NCG bubble exits, small gas bubbles are still present in

the suppression pool water; however the void fraction of this gas is significantly less than

the initial large gas “jet”. The average suppression pool void fraction is expected to

decrease to low levels by about 10 seconds, but much lower levels of noncondensable

gas void fractions levels may last in the suppression pool for about 30–40 seconds.

3. Steam condensation oscillation (CO) and chugging—After the NCG flow reaches very

low levels, the bulk of the flow exiting the vent is a steam-water mixture. Initially this flow

rate can be significant but will decrease as a function of time following the LOCA. These

steam condensation modes, which generally exist at times later than 10 seconds after

the LOCA, are not expected to impair ECCS pump operation because the fraction of

NCG exiting the vent and entering the suppression pool is small, and because any

steam in the suppression pool is expected to condense before entering the ECCS pump

inlet.

The determination of vent flow rates into the suppression pool following a LOCA is dependent 

on the break mass and energy flow and the resultant drywell pressure. This information is 

analytically provided by a containment analysis computer code. This report concentrates on the 

post-LOCA predictions for a typical Mark I plant design. A Mark I containment results in the 

most limiting NCG suppression pool conditions because of its relatively small suppression 

chamber (wetwell) volume and the small downcomer vent submergence when compared to the 

larger Mark II and Mark III suppression pool designs (Figure 2.1-2). Additionally, the ECCS 

pump strainers are generally located closer to the downcomer vent exits in a Mark I design.
xiii

Figure 2.1-2: Cross Section and Dimensions of Typical (a) Mark I, (b) Mark II and 

(c) Mark III Suppression Chambers (Dimensions in meters (feet))

Considerations for ECCS Pump Operation

ECCS pump operation could primarily be affected if the pump operates within the first 

2–3 seconds following a DBA LOCA. Concerns regarding NCG ingestion by the ECCS pumps 

would be expected to last until the bulk of the noncondensable gas rises and exits the pool 

surface. 

For a DBA LOCA, the ECCS (low-pressure core spray (LPCS) and low-pressure core injection 

(LPCI)) pump startup signal is typically generated at the high drywell pressure set point. A 

survey by the BWR Owners Group for Mark I containment designs (Ref. 4) reveals that a 

high-containment pressure signal can occur as early as 1 second following the start of a DBA 

large-break LOCA (LBLOCA). An additional 3-to-10 second pump startup delay is to be 

expected if a loss of offsite power is considered concurrent with the DBA LBLOCA to account 

for emergency diesel generator startup and ECCS pump initiation sequencing. Therefore, the 

earliest start time of the ECCS pumps with the availability of offsite power is about 1 second, but 

is about 4 seconds with an associated loss of offsite power.

During the initial period after receipt of a startup signal the LPCS and LPCI pumps are operated 

to reach full speed. Additionally, the initial pump flow is recirculated to the suppression pool and 

is not injected into the core. Regardless of the availability of offsite power, the LPCS and LPCI 

systems are expected to pump flow into the core at about 40 and 50 seconds after appropriate

valve reconfiguration. Consequently, during the first 40 to 50 seconds following a LOCA, when 

the ECCS pumps circulate flow within the suppression pool, the effects of NCG ingestion do not 

affect the ability of the ECCS pumps to provide flow to the core. The primary concerns in this 

early time frame concerns the possibility of incurring pump damage or, if pump damage does 

not occur, the ability of the ECCS pumps to recover to full operation after possible 

noncondensable gas ingestion during the early period. 
xiv

Figure 6.1-1 presents a summary of the NCG behavior in the suppression pool. Information 

regarding the typical startup and full operation timing of the ECCS pumps of a Mark I 

containment design following a DBA LOCA is also presented on Figure 6.1-1.

Figure 6.1-1: Downcomer Flow/Suppression Pool Conditions 

 and ECCS Pump Operation Comparisons

Prediction of Post-LOCA Suppression Pool Flow Conditions 

Although many of the earlier tests did not provide specific measurements of suppression pool 

thermal-hydraulic conditions, many of the programs included activities to develop analytical and 

computer models for the noncondensable gas phenomena in the suppression pool. CFD

computer codes have been used to simulate the thermal-hydraulic behavior in the suppression 

pool following a LOCA. These analytical models and methods must be verified for application to 

an actual plant geometry. Consequently, the CFD analysis methods were compared to test data 

from several of the test facilities. 

Considering the experience with CFD analyses for analyzing the vent clearing and 

noncondensable gas portions of the downcomer vent flow into the suppression pool following a 

LOCA, it is reasonable to assume that a CFD analysis could be used to analyze the initial vent 

clearing and noncondensable gas flow transient in a full-scale plant geometry. In Section 5, the 

analysis techniques for a current CFD code were benchmarked against data from two scaled 

test facilities and qualitatively compared to a video of the Mark II full-scale tests. The CFD 

techniques were then used to develop models of a full-scale plant geometry and used to 

simulate the noncondensable gas behavior in a suppression pool. Section 5 of this report 

describes the CFD analyses used to define a noncondensable gas bubble “exclusion zone” and 

also provides an estimate of the time dependent noncondensable void fraction distribution in the 

suppression pool around the vent pipe location following a LOCA.

It should be noted that the CFD models are primarily applicable to the analyses of the vent 

clearing and noncondensable bubble phase. CFD analysis of the steam bubble CO and 

chugging phases require modeling of contact condensation at the steam bubble surface. The 

contact condensation phenomenon is not considered in the current analysis.

Pump Start Signal ≥1 to 20s

Vent Water 

Clearing≈0.3s Small NCG bubble Void Fraction ~0.3 to ~20s Downcomer steam Flow after ~20s

~0.3 s 20 s 40 s

Start Pump Core Injection at ~40s

0 s

Large NCG Jet

~0.3 to <~2-3s

Largest NCG 

Concentration

~2-3 s

Pump Startup & Pool Recirculation ≥4 to ~40s

With offsite power

With loss of offsite power Start Pump Core Injection at ~40s

Pump Startup & Pool Recirculation ≥1 to ~40s

Pump Start Signal ≥4 to 20s
xv 

Figures 5.2.3-5, 5.2.3-6, 5.2.3-7 and 5.2.3-8 provide a time dependent definition of the 

“exclusion zone” and the suppression pool void fraction.

Figure 5.2.3-5: Contour Plots of Time Dependent Bubble Interface (VOF Method)
xvi

Figure 5.2.3-6: Contour Plots of Time Dependent Void Fraction (Eulerian Method)
xvii

Figure 5.2.3-7: Time Dependent Maximum Penetration Depth

Figure 5.2.3-8: Time Dependent Radial Width of Main Bubble

Section 4 of this report addresses the acceptable operating range of the ECCS pumps to define 

the operating conditions necessary to prevent pump damage if noncondensable gas ingestion 

occurs or to determine the ability of the pump to recover from noncondensable gas ingestion. 
xviii

The pump requires additional time to reach full speed after receipt of the startup signal; 

however, for conservatism, the pump is assumed to instantaneously reach full speed.

An ECCS pump can sustain damage caused by over speed and vibration if gas is ingested. The 

acceptable ECCS pump operating range for noncondensable gas inlet void fraction is discussed 

in Section 4 which also discusses the ability of the ECCS pump to recover after a limited 

amount of gas is ingested. 

The BWR ECCS operational criteria presented in Section 4 is summarized below.

1. Low transient flow rates (less than 70 percent of the best efficiency point) with an inlet

void fraction greater than 5 percent for more than 5 seconds will result in pump damage

unless pump specific testing can show otherwise.

2. Higher transient flow rates (between 70 percent and 120 percent of the best efficiency

point) with an inlet void fraction greater than 10 percent for more than 5 seconds will

result in pump damage unless pump specific testing can show otherwise.

3. Low flow operation (less than 40 percent of the best efficiency point) at greater than

1 percent by volume continuous suction noncondensable gas void fraction results in

pump damage unless pump specific testing can show otherwise.

4. Full flow operation (40 percent to 120 percent of the best efficiency point) at greater than

2 percent by volume continuous suction noncondensable gas void fraction results in

pump damage unless pump specific testing can show otherwise.

ECCS pumps that operate within the void fraction conditions specified in Table 4.1-1 and within 

the gas void fractions indicated above in items 1 to 4 can be considered capable of recovery 

without damage, and will operate at normal pump head and flow conditions if the inlet void 

fraction is reduced to a value below 1 percent. 

An ECCS pump assessment would be permitted to use other operation criteria if pump specific 

tests are performed to provide better operational guidance.

Assessment Approach for ECCS Pump Operation in BWR Pressure 

Suppression Containments

Section 6 describes a method to perform a plant-specific assessment to determine whether an 

ECCS pump will acceptably operate following a DBA LOCA in a Mark I suppression pool using 

the analyses results and criteria presented in this report. The assessment approach uses the 

ECCS pump strainer location, the ECCS pump operation criteria, the ECCS pump startup and 

full operation timing information, and the pool void fraction information to make its 

determination.

It should be noted that ultimately the plant operators will be responsible for applying the 

methods and criteria to their specific plant geometry to complete assessment for a particular 

plant. The assessment of ECCS pump operation following a LOCA would follow the following 

steps:
xix

1. Determine the post-LOCA noncondensable gas “exclusion zone” in the suppression

pool.

2. Determine the void fraction in the pool water outside the “exclusion zone”.

3. Determine the location of the ECCS pump strainer in relation to the “exclusion” zone.

4. Determine the ECCS pump performance if noncondensable gas is present at the pump

strainers.

5. If the assessment determines that the ECCS pump operation would be affected, the

plant operator would need to determine a course of action to correct the condition.

After pump startup, the criteria for acceptable pump operation will be determined using the flow 

logic presented in Figure 6.5-1, which provides a logic stream for this procedure to assess the 

post-LOCA operation of the ECCS pumps. The procedure proposes that the ECCS pump 

operation be assessed using the maximum and transient noncondensable gas void fraction at 

the strainer location using the graphs contained in Figures 5.2.3-5, 5.2.3-6, 5.2.3-7 and 5.2.3-8.

Additionally, Section 5.2.3.5 and Figure 5.2.3-9 show that the void fraction at a strainer located 

below the downcomer exit and outside of the large bubble will be 1 percent or less for times 

greater than about 2 seconds. This assessment approach is proposed because of the short 

duration of the large NCG large bubble which is less than 3 seconds. For this assessment it is 

assumed that the “exclusion zone” for the large NCG bubble is defined by the boundary where 

the pool void fraction exceeds 10 percent. This assumption is consistent with the ECCS pump 

acceptable operation criteria listed in Section 6.4.1. Other void fraction criteria are also 

consistent with the information listed in Section 6.4.2.

The assessment procedure presented in Figure 6.5-1 was developed using information for a 

Mark I containment. Because of similar vertical downcomer geometry, the results of this study 

can be conservatively applied to a Mark II containment design, which possesses a larger 

suppression chamber. A Mark III design includes a horizontal vent pipe orientation and the 

ECCS pump inlets are located significantly further from the vent exits. Consequently, the ECCS 

pumps in a Mark III design are expected to be less affected by NCG because lower void 

fractions are expected at their inlets.

Figure 6.5-1 includes an option to perform ECCS pump specific testing if the application of the 

flow logic determined that the timing and void fraction criteria results in pump failure including 

the inability to recover. This testing option would provide the actual operability assessment of a 

particular ECCS pump.

If an ECCS pump assessment concludes that a pump would fail and not recover, the plant 

operator must determine what corrective action would be necessary. Four possible corrective 

actions could be considered: 

1. The ECCS start time could be delayed to ensure that the pump would start after

elevated gas void fraction conditions are reduced in the suppression pool at the

elevation of the pump strainer and intake.

2. The location of the ECCS pump strainer and intake could be moved to eliminate or

minimize the gas void fraction concern at the pump strainer and intake.
xx 

3. Closing off a specific pump strainer is also an option if other strainer intakes provide

sufficient flow into a common header feeding a pump.

4. As outlined below, the licensee may also choose a more detailed plant-specific analysis

that models the strainer and downcomer locations and explicitly calculates the voiding at

the ECCS pump inlet.

The method outlined in this report is limited by the assumptions used to develop the criteria. For 

instance the assumptions regarding the typical ECCS pump start times is based on the survey 

information present in Reference 4. The analyses used to develop the criteria were developed 

for a DBA LBLOCA in the suppression pool of a Mark I containment using the information 

provided in References 2. The DBA LOCA results in the most severe NCG conditions in a 

suppression pool because it provides the largest and fastest break flowrates that result in the 

largest gas flow into the suppression pool in the shortest time. Additionally, the ECCS pump 

strainers are generally located closest to the downcomers in a Mark I design. The pump 

operation acceptance criteria were obtained from the literature search and testing information 

described in Sections 3 and 4. The CFD approach used to predict the location and duration of 

the NCG phase is also limited to the simplified cases considered. Other phenomena, such as 

suppression pool structures or flow patterns that could further enhance, mix, or distribute the 

NCG region, should be considered on a case-by-case basis.

For this assessment, ECCS pump performance during operation with void fraction conditions at 

the pump inlet was considered from a general, not plant-specific perspective. A review of 

Figure 6.5-1 reveals that an ECCS pump start time greater than about 2–3 seconds may be 

sufficient to conclude that pump operation is unaffected. This assessment supports the technical 

conclusion that delayed ECCS pump flow initiation beyond 3 seconds, plus a possible safety 

factor, following a DBA LOCA initiation will adequately protect the ECCS pumps from damage. 

Similarly, consideration of the void fraction at the strainer location may be sufficient to conclude 

that pump operation is unaffected. This assessment approach assumes that the void fraction at 

the strainer location is transferred unchanged through the strainer and piping to the pump inlet. 

This is a conservative assumption with regards to the determination of acceptable pump 

operation. The assessment includes the option to refine the pump operation assessment by 

using calculations to determine the void fraction at the pump inlet. Other studies and tests have 

demonstrated that the void fraction at the strainer location could be substantially different after 

traveling a distance through a pipe to the pump inlet. Consequently, calculational methods 

outlined in Reference 44 can be used to estimate the gas transport in the pipe from the strainer 

to the pump inlet in order to determine the actual void fraction at the pump inlet. Additionally, a 

defensible method would need to be developed to calculate the transport of noncondensable 

gas from the large bubble or small bubbles in the suppression pool, through the strainer and into 

the pipe downstream of the strainer. Of course, the method used to perform this calculation (e.g. 

a thermal-hydraulic computer code) must be verified against applicable test data before it could 

be used in this application.

Another approach to determine the void fraction at the pump inlet would be to include the 

strainer inlet and piping leading to the ECCS pump in the full CFD model used to predict the 

noncondensable gas bubble behavior. This model would include details of the pump strainer 

and piping along with the time dependent pump suction flow at the end of the piping. This 

approach would not only estimate the void fraction in the piping approaching the pump inlet, but 

would also predict the interactions between the size and location of the noncondensable gas 

bubble and the calculated suppression pool flow field and void fraction distribution that includes 
xxi

the effect of the pump suction. Of course, the CFD method used to perform this calculation 

should be benchmarked against applicable test data before it is used in this application.
xxii

Is strainer location

α > 10% for > 5 s?

Yes

No

Is pump start time 

≥ 3 s?

Yes

Acceptable pump 

operation

No Yes

No

Pump failure 

Move strainer, block strainer or adjust start time

Is strainer location 

α ≤ 1%?

Yes

Will pump fail?

Yes

No

Can pump recover?

Yes

No

Yes

No

Assume pump failure. 

Move strainer, block 

strainer or adjust start time

Is strainer location

α ≤ 5% for ≤ 5 s 

Acceptable pump 

operation

Determine time dependent void fraction (α) at strainer location

Figure 6.5-1: Approach for Assessing Effects of Noncondensable Gas on ECCS Pump Operation

Determine pump start time

Acceptable pump 

operation

Determine time (tα) when α ≤ 5%

Is pump start time 

≥ tα

No

Acceptable pump 

operation

Yes

No Perform pump specific testing? AND/OR

Perform detailed plant specific analysis to determine pump 

inlet void fraction?
xxiii

ACRONYMS AND ABBREVIATIONS

BEP best efficiency point

BWR boiling water reactor

CFD computational fluid dynamics

CO condensation oscillation

CS containment spray

DBA design-basis accident

DW drywell

ECCS emergency core cooling system

GE General Electric

GI generic safety issue 

ID internal diameter

LLL Lawrence Livermore Laboratory

LOCA loss-of-coolant accident

LPCI low-pressure core injection

LPCS low-pressure core spray

MIT Massachusetts Institute of Technology

NCG noncondensable gas

NOG Nordic Owners Group

NPP nuclear power plant

NPSH net positive suction head

NPSHR net positive suction head required

NRC U.S. Nuclear Regulatory Commission

PS pressure suppression

PSI Paul Scherrer Institute

PUMA Purdue University Multi-Dimensional Integral Test Assembly

PWR pressurized water reactor

RHR residual heat removal

SRV safety relief valve

VOF volume of fluid

WW wetwell

1 

1. INTRODUCTION

This report summarizes the work and recommendations developed to support the technical 

basis for resolving U.S. Nuclear Regulatory Commission (NRC) Generic Safety Issue (GI) 193, 

“Boiling Water Reactor (BWR) Emergency Core Cooling System (ECCS) Suction Concerns.” 

The following text from Reference 1 states the safety concerns related to this activity: 

GSI-193 addresses the possible failure of the ECCS pumps due to unanticipated, 

large quantities of entrained gas in the suction piping from BWR suppression 

pools. The issue applies to Mark I, II, and III containments during large- and 

medium-break LOCAs, and could potentially cause pump failure or degraded 

performance due to gas binding, vapor locking, or cavitation.

Section 2 describes the analytical methods used to calculate post loss-of-coolant accident 

(LOCA) BWR pressure suppression containment conditions including suppression pool 

behavior. 

Section 3 lists the test facilities used to test BWR pressure suppression systems and 

summarizes the test data available from these facilities, which are applicable to the GI-193 

activity. 

Section 4 of this report describes the acceptable ECCS pump operating range necessary to 

prevent pump damage if noncondensable gas ingestion occurs or to permit recovery from 

noncondensable gas ingestion. Information regarding the typical startup and full operation 

timing of the ECCS pumps of a Mark I containment design following a design-basis accident 

(DBA) LOCA is also presented.

Section 5 of this report uses analyses to define a noncondensable gas bubble “exclusion zone” 

and also provides the time-dependent noncondensable void fraction distribution in the 

suppression pool outside the large gas bubble following a LOCA. 

Section 6 describes a method to perform a plant-specific assessment to determine whether an 

ECCS pump will operate acceptably following a DBA LOCA in a Mark I suppression pool using 

the analyses results and criteria presented in this report. The assessment approach uses the 

ECCS pump strainer location, the ECCS pump operation criteria, the ECCS pump startup and 

full operation timing information, and the pool void fraction information to make its 

determination.

3 

2. BWR POST-LOCA CONTAINMENT ANALYSIS AND

SUPPRESSION POOL BEHAVIOR

Boiling water reactors (BWRs) in the United States use containment pressure suppression 

systems to minimize the containment pressure rise following a loss-of-coolant accident (LOCA). 

The containment systems include a drywell that contains the reactor vessel and piping, a 

suppression chamber or wetwell that contains a gas atmosphere and a pool of water to provide 

the pressure suppression function, and a downcomer vent system that connects the drywell to a 

submerged position in the wetwell suppression pool. Following a LOCA, the drywell atmosphere 

and break mass flow through the downcomer vent system and enter the suppression pool. The 

pool condenses the break steam flow and removes energy from any noncondensable gas 

(NCG).

2.1 Containment Analysis

The determination of vent flow into the suppression pool following a LOCA is dependent on the 

break mass and energy flow and the resulting drywell pressure. This information is analytically 

provided by a containment analysis computer code. Reference 2 describes the containment 

analyses and provides the predictions for typical Mark I, Mark II, and Mark III (Figure 2.1-1) 

containments following a DBA LOCA. The DBA LOCA for a BWR pressure suppression (PS) 

system is typically a double-end rupture of a main steam line or a recirculation line. The 

following sections concentrate on the post-LOCA predictions for a typical Mark I plant design. 

This design is considered the most limiting for GI-193 concerns because it contains a relatively 

small suppression chamber (wetwell) volume and small downcomer vent submergence when 

compared to Mark II and III containment designs (Figure 2.1-2). The Mark I designs can also 

possess ECCS pump inlet strainer locations closer to the discharge end of the downcomer vent 

(Figure 2.1-3). Because of similar vertical downcomer geometry, the results of this study can be 

conservatively applied to a Mark II containment design, which possess a larger suppression 

chamber. A Mark III design includes a horizontal vent pipe orientation and the ECCS pump 

inlets are located significantly further from the vent exits. Consequently, the ECCS pumps in a 

Mark III design are expected to be less affected by NCG because lower void fractions are 

expected at their inlets.

Figure 2.1-1: BWR Pressure Suppression System Designs

(1-Primary Containment, 2-Drywell, 3-Wetwell, 4-Suppression Pool, 5-Vent System)
4 

Figure 2.1-2: Cross Section and Dimensions of Typical (a) Mark I, (b) Mark II and 

(c) Mark III Suppression Chambers (Dimensions in meter (foot))

Figure 2.1-3: Typical ECCS Pump Strainer Locations in (a) Mark I, (b) Mark II and 

(c) Mark III Suppression Chambers (Dimensions in meter (foot))

The DBA LOCA is considered the most severe for assessing the effects of possible NCG flow 

on the ECCS pump operation because a large double-ended DBA LOCA produces the largest 

drywell pressure rise rates and the highest containment pressures resulting in the largest vent 

flow rate into the suppression pool. The larger vent flow produces a larger NCG vent “jet” than a 

smaller break. The larger gas “jet” defines the suppression pool “exclusion zone,” which can 

possess a large NCG mass. Typically, ECCS pump suctions should not be located in or near 

the “exclusion zone” to ensure successful pump operation.
5 

As shown on Table 2.1-1, the downcomer flow following the initial clearing of the submerged 

water volume in the downcomer pipe can be divided into four phases: (1) a phase when the 

water contained in the submerged downcomer pipe is cleared, (2) a large NCG bubble or “jet”

phase, (3) a small NCG bubble phase, and (4) an all steam flow phase. Table 2.1-1 indicates 

that the largest break DBA LOCA produces the largest NCG volume in the suppression pool. 

Consequently, the DBA LOCA is assessed in this study.

Table 2.1-1: Post-LOCA Downcomer Flow Phases after Vent Water Clearing

Downcomer Flow 

Phase

Largest Break DBA 

LOCA

Medium Break Small Break

Phase 1: 

Clearing of submerged 

vent water

• Fastest water clearing • Slower water clearing • Slowest water

clearing

Phase 2: 

Large NCG Bubble

• Largest NCG bubble

• Largest bubble

penetration

• Smallest large bubble

duration

• Smaller NCG bubble

• Smaller bubble

penetration

• Longer large bubble

duration

• Smallest or negligible

large NCG bubble

• Smallest bubble

penetration

Phase 3: 

Small NCG Bubbles

• Smallest small NCG

bubble duration

• Largest small NCG

bubble penetration

• Longer small NCG

bubble duration

• Smaller small NCG

bubble penetration

• Longest small NCG

bubble duration

• Smallest small NCG

bubble penetration

Phase 4:

All steam flow

• Localized

condensation 

oscillation or chugging

• Localized

condensation 

oscillation or chugging

• Localized

condensation 

oscillation or chugging

2.1.1 Vent Discharge Flow

Figures 2.1-4 and 2.1-5 show typical Mark I pressure and temperature response plots following

a DBA double-ended recirculation line break, which is the DBA LOCA presented in Reference 2. 

The plots indicate the similarity of calculated results using either the MELCOR or CONTAIN 

computer codes.
6 

Figure 2.1-4: Predictions of Mark I Containment Drywell (DW) and 

Wetwell (WW) Pressures following a DBA LOCA

Figure 2.1-5: Predictions of Mark I Containment Drywell (DW), Wetwell (WW) and 

Suppression Pool (SP) Temperatures following a DBA LOCA
7 

The vent discharge flow rate and the resultant discharge “jet” are important considerations in the 

determination of the “exclusion zone” and successful post-LOCA operation of the ECCS pumps. 

Reference 2 describes the calculation of the downcomer vent flow from the drywell to the 

wetwell suppression pool following a Mark I DBA LOCA using the MELCOR computer code. 

Figure 2.1-6 shows steam, water droplet, and NCG flows at the entrance to the vent system 

between the drywell and wetwell. Figure 2.1-7 shows the downcomer vent system exit flow rates 

for liquid, steam, water drops, and a nitrogen-oxygen gas mixture, which typifies the NCG, 

between the drywell and suppression pool. Figure 2.1-8 plots the time-dependent

noncondensable mass fraction at the downcomer exit. The total downcomer NCG flow is limited 

by the total mass of noncondensables initially present in the drywell. The LOCA is assumed to 

occur at 0 seconds. The initial clearing of the downcomer water column is predicted to occur by

0.3 seconds following the start of the LOCA after which the steam, water droplet, and NCG flow 

begins to exit the downcomers. These figures show that the NCG flow peaks before 1 second 

and reaches very small values before 10 seconds. Consequently consideration of NCG effects 

on ECCS pump operation is not considered to be a significant concern after the first few 

seconds following the LOCA because the large NCG volume “exclusion zone” at the vent exit 

will dissipate within 10 seconds. Based on analyses presented in Section 5.2, concerns 

regarding NCG ingestion by the ECCS pumps are expected to be largest during the first 

2 seconds and drop significantly after that time because the major portion of NCG rises and 

exits the pool surface. The average suppression pool void fraction is expected to significantly 

decrease within 10 seconds. However, much lower levels of NCG void fractions may remain in 

the suppression pool for about 30–40 seconds.

Figure 2.1-6: MELCOR Predictions of Mark I Post-DBA LOCA 

Constituents of Vent System Entrance Flow
8 

Figure 2.1-7: MELCOR Predictions of Mark I Post-DBA LOCA 

Constituents of Downcomer Exit Flow

Figure 2.1-8: MELCOR Predictions of Mark I Post-DBA LOCA Noncondensable 

Gas Mass Fraction of Downcomer Exit Flow
9 

2.2 Post-LOCA Suppression Pool Behavior

As indicated in Figure 2.1-7, suppression pool behavior due to vent flow following a LOCA can 

be divided into phases: (1) the vent water clearing phase, the large NCG flow and small bubble

phases, and (2) the steam condensation oscillation and chugging phase. The following sections 

describe the suppression pool phenomena occurring in each phase and identify the phase that 

presents the major concern for NCG ingestion into the ECCS pump suction.

2.2.1 Downcomer Vent Water Clearing

A water column exists in the downcomer vent before the start of a LOCA. As the drywell 

pressure increases following a LOCA the water column is pushed into the suppression pool 

because of the increased pressure differential between the drywell and wetwell. The water 

clearing phase lasts about 0.3 seconds for a double-ended DBA LOCA, but can take longer for 

smaller break sizes. The vent clearing phenomenon can produce significant turbulence in the 

suppression pool and pressure loads on the walls of the suppression pool and on submerged 

structures.

2.2.2 Noncondensable Gas Flow and Bubble Behavior

NCG flow behavior can be divided into two phases, a large gas bubble phase followed by a 

longer small gas bubble phase. Following the completion of vent water clearing, the NCG-steam 

mixture in the drywell begins to flow out of the vent and into the suppression pool. As shown on 

Figure 2.1-7, the initial flow in this phase contains a significant amount of NCG. However, as 

previously indicated, the amount of NCG in the vent flow decreases quickly and reaches very 

small levels before 10 seconds for a DBA LOCA. The initial large NCG “jet” can possess a 

length extending to the bottom of the pool with a radially expanding and oscillating surface. The 

initial gas discharge produces large and small bubbles which causes a swelling of the 

suppression pool surface. As discussed in Reference 3, the gas bubble can undergo 

expansion-contraction oscillations that have been shown to follow a Rayleigh bubble oscillation 

behavior (Figure 2.2-1). The pool swell and bubble oscillation can create loadings on the 

suppression pool walls and on submerged structures. The size of the initial NCG “jet” can be 

used to define a volume at the vent exit, which possesses a large fraction of NCG. This volume 

can be used to define an “exclusion zone” in which an ECCS strainer and pump suction should 

not extend during ECCS pump operation. However, it should be noted that the size of the NCG 

volume will decrease as the gas flow decreases, and as the gas rises and exits the top of the 

suppression pool.
10

Figure 2.2-1: Raleigh Gas Bubble Oscillation in Liquid (Ref. 3)

2.2.3 Condensation Oscillation and Chugging

As shown in Figure 2.1-7, the NCG flow increases quickly after vent clearing and subsequently 

decreases reaching very small levels well before 10 seconds following a DBA LOCA. After vent 

clearing the vapor-liquid vent flow increases and peaks before 5 seconds following the DBA 

LOCA and subsequently decreases. Consequently, the vent flow between 0.3 and 10 seconds 

consists of a decreasing amount of noncondensable gas with an increasing amount of vapor 

and liquid. The steam condensation mode, which generally exists at later times following a DBA 

LOCA when the vent flow is primarily steam-liquid flow, is not expected to impair ECCS pump 

operation because the fraction of NCG exiting the vent and entering the suppression pool is 

small and because any steam in the suppression pool is expected to condense before entering 

the ECCS pump inlet. 

After the NCG flow reaches very low levels, the bulk of the flow exiting the vent is a steam-water 

mixture. Initially this flow rate can be significant but will decrease as the vent flow decreases 

following the LOCA. Figure 2.2-2 shows the condensation modes that have been observed 

during a LOCA discharge. High steam flow results in steady condensation. Decreasing steam 

flow can produce oscillating bubbles at the vent exit resulting from steam condensation at the 

bubble surface. This phenomenon is called condensation oscillation (CO) (Figure 2.2-3). At 

even lower steam flow conditions, the exiting steam flow can create a bubble which rapidly, fully 

condenses or collapses at the vent exit causing backflow of water into the vent pipe. This 

condition can occur in a cyclical fashion consisting of steam bubble formation and full collapse. 

This phenomenon is called chugging. Figure 2.2-4 presents a typical flow map showing the 

conditions which support chugging conditions. The CO and chugging phenomena are the 

dominant condensation modes that can cause loadings on the suppression pool walls and on 

submerged structures.
11

Figure 2.2-2: Schematic of Typical Condensation Mode Regions 

during LOCA Blowdown (Ref. 3)

Figure 2.2-3: Steam Condensation Oscillation at Vent Exit (Ref. 3)
12

Figure 2.2-4: Sampling Chugging Map for Vertical Vents (Ref. 3)
13

3. LITERATURE SURVEY

This section lists the facilities used for testing BWR pressure suppression systems and 

summarizes the test results applicable to addressing the GI-193 issue. A summary of the 

analytical techniques used to predict the post-LOCA suppression pool conditions is also 

presented.

3.1 Pressure Suppression Experimental Facilities and Test Data

Since the late 1970s, numerous test facilities have been used to provide test data for the 

behavior of a pressure suppression system. However, the primary objective of many of these 

test facilities was to assess the pressure loadings on the walls of the suppression pool and not 

to provide information regarding the thermal-hydraulic conditions in the suppression pool. A 

listing of the pressure suppression tests and facilities, and the associated reference documents 

are provided in Table 3.1-1. This table also identifies the main purpose of the test facility and 

lists associated analytical efforts to model pool behavior following air injection through a vent 

pipe.

The primary tests and facilities that provide data specific to GI-193 activities are those tests 

performed at Lappeenranta University and Purdue University. Because both these tests use 

scaled facilities, it is necessary to develop an approach to apply the scaled test results to a plant 

geometry.

3.1.1 General Electric Pressure Suppression Testing

General Electric (GE) performed many full-scale tests and tests scaled to the wetwells and 

suppression pools of Mark I, Mark II, and Mark III containments in the late 1970s and early 

1980s. The suppression pool testing modeled the post-LOCA vent flow entering the suppression 

pool and discharge flow entering the suppression pool from safety relief valve (SRV) operation. 

The primary purpose of these tests was to assess loadings on the suppression pool walls and 

submerged structures following a LOCA or SRV actuation. The test data was used to develop 

and verify calculational methods to determine the wall and structure loadings, and to support 

design modifications as required.

The GE tests did not specifically measure the thermal-hydraulic conditions in the suppression 

pool following a LOCA or SRV actuation. However some thermal-hydraulic models of the vent 

discharge phenomena were developed as part of the efforts to determine wall and structure 

loads. For example efforts related to the GE testing developed models to calculate NCG bubble 

oscillations at the vent exit using Rayleigh bubble theory. Thermal-hydraulic models of the CO 

and chugging phenomena were also developed.

Videos were taken during performance of the GE Mark I and Mark II pressure suppression tests. 

The Mark I and Mark II test videos primarily recorded the chugging phenomena in the 

suppression pool. Thermal-hydraulic measurements were not taken in the suppression pool 

during these tests. Visual examination of the video cannot support or dispel with certainty the 

presence of an ECCS pump safety concern which confirms the need for additional information 

to reach a definitive conclusion on the existence of a safety concern. 
14

3.1.2 NRC 1/5 Scale Mark I Pressure Suppression Testing

The NRC-scaled Mark I pressure suppression tests were performed in the early 1980s to 

support efforts to verify the findings from the GE pressure suppression tests and to support 

rulemaking activities. Testing was performed using a one-fifth scaled test facility built at 

Lawrence Livermore Laboratory (LLL) which modeled a section of a Mark I containment. As with 

the GE tests the primary purpose of these tests was to assess the loads on the suppression 

pool walls and submerged structures resulting from a LOCA.

No specific measurements of suppression pool thermal-hydraulic conditions were made. 

However, efforts were made as part of this program to develop models for the NCG and steam 

blowdown phenomena in the suppression pool. Computational fluid dynamics (CFD) computer 

codes were developed to simulate the thermal-hydraulic behavior in the suppression pool. 

Analysis results using the PELE-IC CFD computer code were compared to test data from the 

MIT air blowdown experiments and from the air bubble data from the LLL one-fifth scale tests. 

Analyses of the LLL air bubble tests were also performed using the MAITAI CFD computer 

code. Ultimately the CHAMP computer code was developed to specifically analyze the air 

discharge experiments at the LLL one-fifth scale Mark I test facility.

3.1.3 Massachusetts Institute of Technology Mark I Gas Blowdown Tests

In the late 1970s, the NRC sponsored air blowdown testing in a scaled Mark I suppression pool 

test facility at the Massachusetts Institute of Technology (MIT). Air was injected through a 

submerged vertical 25.4 mm (1 in.) pipe into a pool of water. As with the GE and NRC tests the 

primary objective of the MIT tests was to support development of loading models for the 

suppression pool wall and submerged structures. However efforts were made to model the pool 

swell dynamics and perform computer simulations of the NCG and steam blowdown into the 

suppression pool.

3.1.4 University of California Mark I Scaled Suppression Pool Tests 

The University of California small-scale testing of a Mark I suppression pool was performed in 

the early 1980s. The experiments injected air through submerged vertical tubes varying in 

diameter from 0.9 to 9.5 cm (0.35 to 3.7 in.) in a pool of water. A hydrodynamic model of the 

liquid vent clearing and the bubble phenomenon at the vent pipe exit was developed using the 

test data. The air bubble oscillation model was based on the Rayleigh bubble equation.

3.1.5 Lappeenranta University Scaled Suppression Pool Experiments

Lappeenranta University in Finland performed experiments starting about 2002 to support the 

understanding of noncondensable gas and steam blowdown in a scaled suppression pool test 

facility. These experiments were performed to support efforts related to Scandinavian BWR 

pressure suppression containment design behavior. Tests were performed at three different test 

facilities. 

3.1.5.1 Condensation Pool Experiments with Noncondensable Gas

The initial condensation pool experiments investigated the formation, size, and distribution of 

NCG bubble in a suppression pool following a postulated LOCA. The tests were generally 
15

scaled to Olkiluoto nuclear power plant (NPP) conditions. The Olkiluoto NPP, Units 1 and 2 

pressure suppression containments (Figure 3.1-1) are ASES-Atom-type BWR containment 

designs which are similar in design to a Mark II containment. The effects of NCG on the 

performance of an ECCS pump were also examined. In some of the tests, compressed air 

flowed through the vent pipes into the water pool in which an ECCS strainer was located. Tests 

were performed using downcomer vents with inner diameters of 162.3 mm (6.39 in.) and 

213.1 mm (8.39 in.). In other tests compressed air was injected directly into the intake of the 

operating ECCS pumps to determine the effect that NCG ingestion would have on ECCS pump 

behavior.

Figure 3.1-1: Schematic of the Olkiluoto Type BWR Pressure 

Suppression Containment (Ref. 31) 

The initial air flow hit the bottom of the pool during the tests using the 162.3 mm (6.39 in.) and 

213.1 mm (8.39 in.) downcomer pipes when the vent pipe was initially filled with water. The vent 

pipe outlets were submerged approximately 2.2 m (7.2 ft) during these tests. Because of the air 

impact, small air bubbles rose from the pool bottom. Pool swell was observed during which the 

rising air lifted a large amount of water up to the surface which was splashing strongly. Water 

carrying a lot of small bubbles was observed to circulate back down close to the pool wall.

Tests using the 162.3 mm (6.4 in.) downcomer pipes were run with an ECCS strainer placed in 

the suppression pool. The vent pipe outlets were submerged approximately 2.2 m (7.2 ft) during 

these tests. The top of the strainer was located 495 mm (1.6 ft) above the vent pipe exit to 

505 mm (1.66 ft) below the pipe exit. During the tests when the vent pipe was initially filled with 

water, the initial air flow hit the bottom of the pool and contacted the ECCS pump strainer. Some 

of the large air bubbles, which appeared after the time of the initial gas “jet” bubbles, were also 

observed to reach the pump strainer. The air bubbles remained at the level of and inside the 

pump strainer for the first 30 seconds because backflow circulation in the pool carried small air 

bubbles from the upper part of the pool to the lower part.
16

Integral tests were performed with a pump strainer located in the pool and an operating ECCS 

pump. For these tests air was injected into the pool through the downcomer vents. These tests 

used the 162.3 mm (6.39 in.) downcomer pipes. The vent pipe outlets were submerged 

approximately 2.2 m (7.2 ft) during these tests. For some tests, air was sucked through the 

strainer into the pump at a constant flow rate when air was blown into the pool through the vent 

pipes. With a volumetric flow rate of 5.5 liter/second (87.2 gpm), air bubbles were not detected 

inside the transparent intake pipe. With a volumetric flow rate of 11 l/s (174 gpm) air bubbles 

were visible in the intake pipe for the first 20 seconds when air was blown through two vent 

pipes. The ECCS pump head and water was not observed to decline during this period.

In the pump tests, the ECCS pump was operating at a nominal speed of 2970 rpm with four 

different volumetric flow rates when pressurized air was injected directly into the pump intake 

pipe. With volumetric flow rates of 57 and 75 l/s (903 and 1,189 gpm) a 3- to 4-percent air 

fraction was needed before a decline in pump head and flow was observed. With smaller flows 

of 12.5 and 25 l/s (198 and 396 gpm) the pump head and flow started to decline after air 

injection was initiated and the flow totally collapsed when more than a 7-percent air fraction was 

present in the intake pipe. When the air injection was stopped for the 12.5 l/s case, the pump 

head and flow reached original values 30 seconds after the air injections was turned off. At all 

other flows the pump head and flow normalized in a few seconds after the air injection was 

turned off.

A supplemental series of tests were also performed using the 162.3 mm (6.39 in.) downcomer 

pipes to study the effects of velocity on the penetration of the air “jet” from the vent pipe and to 

study the circulation of the small air bubbles from the pool surface to the level of the ECCS 

pump strainer. The vent pipe outlets were submerged approximately 2.2 m (7.2 ft) during these 

tests. Compressed air was blown through the vent pipes during this test and the ECCS pump 

was operating at a constant volumetric flow rate of 11 l/s (174 gpm). A throttle valve was used to 

control the maximum air flow velocities. With an air flow between 15 and 32 m/s (49 and 

105 ft/s) the first air “jet” hit the bottom of the pool and broke up into small bubbles which slowly 

rose to the surface. For these high-velocity tests a large number of small air bubbles were seen 

at the level of the ECCS strainer during the first 30 seconds. The air was present because of 

impingement of the air “jet” on the pool bottom and because of the circulation of the small air 

bubbles at the pool wall. Air bubbles were observed inside the strainer during the first 

30 seconds but the amount was negligible after 30 seconds. When the maximum velocity was 

limited to 5 m/s (16 ft/s) the air “jet” didn’t reach the pool bottom and only negligible amounts of 

air bubbles were detected inside the strainer or in the pump intake pipe even though air bubbles 

were noticeable at the strainer level for the first 20 seconds. The ECCS pump head and flow did 

not decline because of air bubbles in the pump during any of the tests.

3.1.5.2 Discharge Tests in the PPOOLEX Facility 

After the previously described suppression pool experiments, which investigated NCG bubble 

behavior, the test facility was modified to include input from air and steam sources. This 

modified facility was named the POOLEX facility, which consisted of a cylinder-shaped stainless 

pool with an open top. The inner diameter of the pool was 2.4 m (7.87 ft), the cross-sectional

area was 4.5 m2 (48.4 ft2

) and the height was 5.0 m (16.4 ft). After completion of a series of 

tests, the POOLEX facility was replaced by the larger closed loop PPOOLEX test facility 

(Figure 3.1-.2). The PPOOLEX pressure vessel includes a wetwell compartment and a drywell 

compartment. The drywell and wetwell are connected by a vertical vent pipe. The main 
17

component of the PPOOLEX facility is a 31 m3 (1,095 ft3

) cylindrical test vessel, 7.45 m 

(24.44 ft) in height and 2.4 m (7.87 ft) in diameter. The test facility was designed for a 4 bar 

(58.0 psi) overpressure and a 0.5 bar (7.25 psi) under pressure. 

Figure 3.1-2: POOLEX (left) and PPOOLEX (right) Test Facilities (Ref. 31) 

Results from two air discharge tests in the PPOOLEX test facility were simulated using the 

FLUENT CFD computer code. The tests were performed using downcomer vents with an inner 

diameter of 214.1 mm (8.43 in.). The vent pipe outlet was submerged approximately 1.05 m 

(3.4 ft) during these tests. The volume-of-fluid (VOF) two-phase model was used and the 

standard k-ε turbulence model was specified. The results obtained from the CFD simulations 

(Ref. 29) are in relatively good agreement with experimental results (Figure 3.1-3). Simulated 

pressures correspond to measurements and fluctuations because of bubble formations and 

break-up are captured. Generally most temperatures were simulated and the largest 

discrepancies could be explained by wetting thermocouples. The analytical activities concluded 

that the main features of downcomer air discharge experiments can be described with CFD 

calculations.

Additional CFD modeling efforts using different computer codes were expended to simulate pool 

loadings because of bubble condensation effects by simulating direct contact condensation of a 

steam or steam-air bubble. The analytical work concerning the steam bubble condensation 

process can be considered a state-of-the-art activity that requires significant additional effort. 

Consequently, because the interfacial heat and mass transfer relations at the bubble surface 

are still in development, CFD modeling of a condensing steam or steam-air bubble was not 

attempted at this time.
18

Figure 3.1-3: Vent Air Bubble Comparisons between CFD 

Calculation and PPOOLEX Experiment (Ref. 29)

3.1.6 Nordic Owners Group Testing for Swedish BWRs

Westinghouse Atom AB performed proprietary testing at the Lappeenranta University 

PPOOLEX pressure suppression test facility for the Nordic Owners Group (NOG). The NOG 

supported testing was performed to assess the effects that post-LOCA NCG flow from the 

downcomers of European BWR PS plants would have on the performance of ECCS pumps that 

take suction from the suppression pool. NRC staff members were allowed to review the 

proprietary report and accompanying videos at the Westinghouse office in Maryland on April 16, 

2013, but were not permitted to make copies of the materials. The following text summarizes the 

document review performed by the NRC Office of Research staff. 
19

The report summarized the testing performed at the PPOOLEX facilities for the NOG to 

investigate the possible suction of NCG into the ECCS intake in a BWR suppression pool 

following a LOCA. The test facility modeled scaled conditions for the Oskarsham and Barstuk 

PS containments. The downcomer pipes connecting the simulated drywell and wetwell of a 

BWR pressure suppression containment were modeled using four 219.1 mm (8.626 in.) 

diameter pipes. The top of the tank was open to atmosphere. The test facility included a 

simulated ECCS pump, transparent piping and a strainer screen. Flow into the ECCS pump was 

pumped back into the test tank. Video recordings were made of the conditions inside the test 

tank and inside the transparent ECCS intake piping.

Several test series were performed. NCG was injected into only one downcomer pipe for the 

first two test series. The ECCS pump strainer was located near the tank wall below the 

downcomer exit. The pump did not fail during any of these tests but, depending on test

conditions, reduced flow may exist during part of the test. The pump fully recovered for all tests. 

Observations from the test videos showed a bubbly flow regime in the pump intake pipe for a 

few seconds. Additionally, for the low-strainer location, the maximum gas concentration present 

in the intake piping was always less than the maximum concentration present in the pool 

immediately outside the pump intake.

Tests were also performed during which NCG was injected into the suppression pool through all 

four downcomer pipes. These tests resulted in less severe conditions in the ECCS piping 

because the gas flow was distributed among four downcomer pipes instead of one. 

Additional tests were performed using the four downcomer pipes for gas injection and with the 

ECCS pump inlet strainer and piping located above the downcomer exit in the suppression pool. 

This geometry is not typical of Mark I and II containments and was not investigated as rigorously 

during the NRC staff review. However, testing did show that an ECCS pump intake location 

above the downcomer exit resulted in larger and prolonged gas concentration in the intake 

piping in comparison to gas concentration conditions in the pool outside the intake. Additionally, 

for an intake location above the downcomer exit, the instantaneous gas concentration in the 

intake piping may be higher at times than that in the pool outside the intake.

Tests also studied the effects of NCG injection on the ECCS pump. One test injected air at the 

pump suction location. Others tested the effects of different gas injection rates on different pump 

flow rates. In another test series, NCG was injected directly into the ECCS pump inlet.

This testing showed that pump operation fully recovered with inlet pipe void fractions below 

about 8 percent. The ECCS pump may experience reduced flow depending on inlet void 

fraction, but the pump always recovered. At inlet void fractions above about 8 percent, the pump 

exhibited a large decrease in pumping capability; however, even for these conditions the pump 

recovered because the NCG injection duration was less than a few seconds. Typically a bubbly 

flow regime was observed in the inlet pipe for void fractions below about 8 percent and a 

stratified flow regime was observed for void fractions above about 8 percent. 

A report describing a CFD analysis of the suppression pool of a NOG plant for the NCG flow 

phase following a LOCA was also reviewed. The CFD analysis was performed using the 

FLUENT code with the VOF method and a k-ε turbulence model. The general conclusion was 

that the CFD analysis produced a good representation of the noncondensable behavior in the 

suppression pool. The CFD analysis predicted a high concentration gas volume in the pool, but 
20

tests showed a large concentration of small gas bubbles in about the same volume. 

Consequently, the report recommended that test results were necessary to provide confidence 

in CFD predictions. The CFD model also included an ECCS gas intake and piping. The pump 

intake was modeled using the FLUENT porous medium model. 

Because the ECCS pump designs used in the NOG plants differ from those in Mark I and II 

plants especially with respect to pump intake location, caution should be taken in making broad 

conclusions from the NOG supported tests. The NOG tests provide trends in post-LOCA 

behavior of NCG in a suppression pool and on pump performance for the NOG design plants, 

but pump performance could differ for the pumps used in Mark I and II plants. Consequently 

specific testing of Mark I and II ECCS pumps would be necessary to fully assess the effects of 

NCG ingestion on those pump designs. Additionally, because the tested downcomer diameters 

were smaller than downcomer diameters in Mark I and II plants the NOG test results would 

need to be scaled to address Mark I and II conditions. 

Summary and Conclusions

The following observations can be drawn from the Nordic Owners Group testing at the 

PPOOLEX facility. Care should be taken in extrapolating the smaller scale test results for 

European BWR pressure suppression designs to full-scale Mark I and II geometries.

1. During testing the noncondensable bubble did not persist in the pool for longer than a

few seconds.

2. With the pump strainer intake below the downcomer exit, gas ingestion was observed in

the pump intake pipe for several seconds; however, the tested pump always recovered

after ingesting NCG although immediate pump operation and efficiency was effected by

high void fraction gas flow to the pump.

These observations are consistent with the information obtained from other sources in the 

literature search for test data and expert opinions related to the GI-193 concerns. The 

observation regarding the NCG bubble duration in the pool is consistent with findings that the 

post-LOCA gas “exclusion zone” at the vent exit will dissipate in a few seconds and that NCG

would only be present in the suppression pool for the time period necessary for the gas to rise 

and exit the pool surface. The observation regarding the pump recovery is consistent with the 

literature search finding that a centrifugal ECCS pump which operates with small inlet NCG void 

fraction for a few seconds can be considered capable of recovery without damage, and can 

operate at normal pump head and flow conditions.

In conclusion, this testing supports the current activities to assess post-LOCA noncondensable 

suppression pool gas concentration on ECCS pumps and suggests that the limited duration 

post-LOCA NCG concentration in a Mark I and II suppression pool would not have a major 

effect on ECCS pump operation.

3.1.7 Purdue University Mark I Scaled Air and Air/Steam Suppression Pool Tests

The NRC supported suppression pool tests at the Purdue University Multi-Dimensional Integral 

Test Assembly for ESBWR applications (PUMA-E) test facility to study the void fraction 

distribution and void penetration in a scaled Mark I suppression pool (Ref. 33). The test 
21

apparatus was scaled for a typical Mark I BWR containment design with considerations for 

downcomer size, suppression pool water level, and downcomer water submergence. 

Instrumentation was installed to measure gas volumetric flow, void fraction, pressure and 

temperature. Two sets of experiments were conducted using the PUMA-E facility. Videos of the 

water pool were taken during the performance of these tests.

In one set of 16 tests air was injected through a 0.102 m (4 in.) downcomer pipe into the 

suppression pool at various air volumetric flow rates, downcomer void conditions and air velocity 

ramp rates. The initial air injection phase resulted in the maximum void penetration depth. 

During this phase the air volumetric flow rate had a minor effect on void fraction distribution and 

void penetration for high volumetric flow rate conditions. The later quasi-steady state phase 

produced less void penetration, but possessed oscillations in the void penetration. The air 

volumetric flow rate resulted in a larger effect on void fraction distribution and void penetration 

for the entire range of air flow rates. The initial downcomer void conditions were found to 

strongly affect the void fraction distribution and void penetration during the initial period. The air 

velocity ramp rates were found to have a minor effect on void fraction distribution and 

penetration in both phases.

In the second test set, sequential flows of air, steam-air mixtures and pure steam at various flow 

rates were injected in the drywell and passed though the downcomer vent into the suppression 

pool. Eight tests with two different downcomer sizes (0.078 m (3 in.) and 0.102 m (4 in.)) with 

various downcomer gas volumetric fluxes and two different initial drywell air concentrations were 

conducted. The initial phase of these tests produced the maximum void penetration depth. 

Chugging was observed at the end phase of these tests. It was observed that the void 

distribution and void penetration area was governed by the downcomer gas volumetric flux and 

the downcomer air concentration.

3.1.8 GKSS Pressure Suppression Research Program

Large-scale tests of pressure suppression systems were also performed at the GKSS test 

facility in Germany (Refs. 21 and 22). The German tests were performed in the 1980s and were 

primarily intended to study the dynamic loadings on the suppression pool boundary and on 

submerged structures. The German tests concentrated on testing for the Mark II and German 

KWU Type-69 PS systems. The tests were performed in a time frame coincident with the GE 

test program and one of the authors was involved with the Mark I tests funded by the NRC in 

the early 1980s. Consequently, most of the information from the German testing should be 

reflected in the GE and NRC test results. However, information regarding flow behavior in the 

water pool was not specifically measured.

The NRC funded testing at the GKSS facility to study the condensation behavior in a Mark II 

suppression pool behavior following a DBA LOCA. One of the videos from the NRC funded 

full-scale Mark II tests recorded the vent clearing and NCG bubble behavior in a suppression 

pool. Figure 3.1-4 provides video frames of the suppression pool conditions at different times 

following the start of the test. The video frames show two of the vent pipes submerged in the 

water pool. The video frames illustrate the different phases of the NCG and steam flow behavior 

in the water pool. The video frames depict a reflection of the initial NCG jet off the bottom 

surface of the pool. Because the distance of approximately 0.98 m (3.2 ft) from the vent exit and 

pool bottom in the test facility is smaller than the similar distances of approximately 2.44 m 
22

(8.0 ft) and 3.7 m (12.4 ft) in a Mark I or Mark II suppression pool, the gas jet reflection may not 

be present in the suppression pool of a Mark I and Mark II plant. 

Thermal-hydraulic measurements were not taken in the suppression pool during these tests; 

however, visual examination of the test video provided meaningful insights into the NCG “jet” 

dimensions and duration to qualitatively identify the “exclusion zone” that defines the volume of 

high NCG mass associated with the discharge vents. Visual examination of the video cannot 

support or dispel with certainty the presence of an ECCS pump safety concern, and confirmed

the need for additional information to reach a definitive conclusion on the existence of a safety 

concern. However, information obtained from the visual examination was useful in identifying an 

appropriate calculational model for GI-193 assessment, and in qualitatively assessing its results. 

0.0 seconds 1.19 seconds

Before Test Start Near Start of Noncondensable Gas Flow

1.29 seconds 1.78 seconds

Large Gas Jet Flow Period Large Gas Jet Flow Period

Figure 3.1-4: Video Frames from a GKSS Full-Scale Test of Mark II 

Suppression Pool Behavior following a LOCA (1 of 3)
23

2.08 seconds 2.18 seconds

Large Gas Jet Flow Completed Large Gas Jet Reflection off Pool Bottom

2.68 seconds 2.98 seconds

Rise of Reflected Gas Jet Near Completion of Reflected Gas Jet Rise

Figure 3.1-4: Video Frames from a GKSS Full-Scale Test of Mark II 

Suppression Pool Behavior following a LOCA (2 of 3)
24

3.98 seconds 14.38 seconds

Small Gas Bubble Distribution During Steam condensation Period

Figure 3.1-4: Video Frames from a GKSS Full-Scale Test of Mark II 

Suppression Pool Behavior following a LOCA (3 of 3)

3.1.9 Marviken Test Facility Testing

A reference to recent pressure suppression test in the Marviken test facility in Sweden in 2007 

(Ref. 23) was discovered. Unfortunately this recent Swedish referenced material is not publicly 

available; however it appears that these tests were primarily performed to verify the operation of 

PS systems.

3.1.10 Paul Scherrer Institute Tests

The large PANDA and LINX experimental facilities were used in the ALPHA project of the Paul 

Scherrer Institute (PSI) to study pressure suppression containments. The PANDA facility is a 

large integral scale test facility whose primary goal was to study passive containment cooling 

systems and the interaction between the various system components during long-term cooling. 

The facility has 1:1 height, 1:25 volume, and 1:25 power scaling to the General Electric 

Simplified BWR design. The facility was later modified to simulate the GE Economic Simplified 

Boiling Water Reactor (ESBWR) and the Siemens SWR-1000 design. CFD calculations were 

undertaken to analyze the three-dimensional drywell flows (Ref. 34). 

The LINX facility has been used for separate effects tests of condensation, mixing, effect of 

NCG and stratification in suppression pools (Ref. 35). CFD analyses have been performed to 

study the behavior of bubble plumes in water pools using other published plume experiments 

because of the unavailability of the LINUX test data at the time the CFD analyses were being 

performed.

Also at PSI, Meier and colleagues (Refs. 36 and 37) conducted small-scale air and steam-air 

blowdown experiments with a transparent tank as part of a thesis at the Federal Institute of 

Technology Zurich. They observed that the length of pipe and compressibility of the gas in the 

piping affects the bubble size, shape and frequency at low flow rates. With high flow rates the 

effect of compressibility became insignificant. Two-dimensional CFD analyses modeled the joint 
25

motion of gas and liquid in the water pool, including the steam bubble interface heat and mass 

transfer. The CFD analyses used the VOF method, which included a newly developed surface 

tension calculation method. Information from this PSI activity appears to have application to 

GI-193 and test data and videos could be used to verify developed calculation techniques for 

addressing the GI-193 concerns. 

3.1.11 Alden Research Laboratory Testing

Alden Research Laboratory performed testing on a 1/2.4 scale model of a segment of a 

Mark I BWR suppression pool to study debris transport. The water volume of the tank per 

downcomer was scaled to the volume per downcomer pipe of a typical BWR Mark I 

suppression pool. Turbulent chugging associated with post-LOCA steam condensation 

oscillations was simulated in these tests by four 0.254 m (10 in.) diameter downcomer pipes 

fitted with pistons. Parametric tests were conducted to examine the type, form, and quantity 

of insulation debris tested; the quantity of sludge tested; and the period and amplitude of the 

downcomer piston chugging. The NRC has videos of some of this project’s tests. This 

testing does not provide information useful to GI-193 assessment.
Table 3.1-1: Pressure Suppression Test Facilities

Organization/Test Facility Downcomer ID

Reference

Test Objectiv

Test T/H

Analysis

General Electric (GE)

• Mark I full-scale test

program

• Mark I 1

/4 scale pressure

suppression test program

• Mark II full-scale test

program

• Mark III full-scale

confirmatory test program

0.6096 m

(24 in.) 

NA

0.5476 m

(21.56 in.)

NA

5 

7, 8

9 

6

-- The primary purpose of the experim

the loading on the suppression poo

LOCA.

Associated assessments used test 

computational methods for predicti

behavior in the suppression pool.

NRC/Lawrence Livermore 

Laboratory (LLL)

• Mark I 1/5 scale BWR

pressure suppression 

tests

0.1194 m 

(4.7 in.)

10, 11,

12

13, 14, 15, 

16, 17

The primary purpose of the experim

the loading on the suppression poo

LOCA. Scaling methods were also 

Associated assessments used test 

models for predicting suppression 

MIT

• Air and noncondensable

Mark I gas blowdown 

experiments

0.0254, 0.0508 and 

0.0983 m

(1, 2, and 3.87 in.)

18, 19 13

Tests for air discharge in a water p

including scaling considerations. A 

developed to analyze pool behavio

compressed air injection from a dry

submerged downcomer pipe.

University of California

• Mark I scaled

suppression pool tests

0.009 to 0.095 m 

(0.35 to 3.7 in.)

20 20

These tests model bubble behavio

pool of water. A theoretical model 

growth at the vent pipe exit was de

GKSS Pressure 

Suppression Research 

Program

0.6 m 

(23.6 in.) 

21, 22 -- These full-scale tests were primari

the dynamic loadings for the GE M

KWU Type-69 pressure suppressio

Marviken Test Facility 

Testing

NA 23 -- References for these recent tests a

are not publicly available; however 

tests were primarily performed to v

pressure suppression systems.

26
Table 3.1-1: Pressure Suppression Test Facilities (cont.) 

Organization/Test 

Facility Downcomer ID

Reference

Test Objective Test T/H

Analysis

Lappeenranta University

• Scaled suppression

pool experiments

0.1623 and 0.2131 m 

(6.39 and 8.39 in.)

24, 25 --

These tests measured gas bubble siz

and without an ECCS pump and stra

also run with direct gas injection into 

• Scaled PPOOLEX test

facility air tests

• Scaled PPOOLEX test

facility air-steam tests

0.2141 m

(8.43 in.) 

0.2141 m 

(8.43 in.)

26, 49

27, 28, 29, 

30, 31, 49

CFD analyses predicted air bubble s

the PPOOLEX tests.

CFD analyses for air-steam bubble in 

Lappeenranta University

• Nordic Owners Group

(NOG) testing for 

Swedish BWRs

0.2191 m 

(8.626 in.) 

-- --

These proprietary tests were perform

facility to address the effect of air ing

pump behavior. NRC staff reviewed t

describing these tests.

Purdue University

• Mark I scaled air and

air/steam 

0.078 and 0.102 m 

(3 and 4 in.)

33 --

These test studied air and air/steam 

submerged vent pipe that exits into a 

fraction is measured in the water poo

Paul Scherrer Institute

• GE and Siemens

pressure suppression 

system design concept 

tests using the PANDA 

test facility

• Small tank tests

NA

5 cm 

(1.97 in.) 

34, 35

36, 37

--

36, 37

The goal of the project was to study p

cooling systems and the interaction a

system components during long-term 

This work involved CFD modeling of 

steam-air blowdown experiments with 

A request has been made to gain acc

and videos.

Alden Research 

Laboratory

• Mark I post-LOCA

debris transport tests

0.2499 m 

(9.84 in.)

38 -- A 1

/2.4 scale model of a segment of a 

suppression pool was used to study d

testing does not provide information 

assessment.

27

29

4. ECCS PUMP OPERATION

An important consideration in assessing the acceptable post-LOCA operation of an ECCS pump

is the determination of the maximum acceptable pump inlet void fraction that will not affect 

pump performance and not cause damage. An additional consideration is the determination of 

the minimum void fraction range and related operation time from which an ECCS pump will 

return to acceptable operation after liquid flow is reestablished. This section presents the current 

knowledge regarding these topics resulting from a literature review.

4.1 ECCS Pump Timing Considerations

Figure 2.1-7 shows that ECCS pump operation could primarily be affected if the pump operates 

within the first few seconds following a DBA LOCA. Concerns regarding NCG ingestion by the 

ECCS pumps would be expected to last until the bulk of the NCG rises and exits the pool 

surface. 

For a DBA LOCA, the ECCS low-pressure core spray (LPCS) and low-pressure core injection 

(LPCI) pump startup signal is typically generated at the high drywell pressure set point. A survey 

by the BWR Owners Group for Mark I containment designs (Ref. 4) reveals that the 

high-containment pressure signal can occur as early as 1 second following the start of a DBA 

LBLOCA. An additional 3 to 10 second pump startup delay is to be expected if a loss of offsite 

power is considered concurrent with the DBA LBLOCA to account for emergency diesel 

generator startup and ECCS pump initiation sequencing. Therefore, the earliest start time of the 

ECCS pumps with the availability of offsite power is about 1 second, but is about 4 seconds with 

an associated loss of offsite power.

DBA LBLOCA analyses are typically performed assuming the unavailability of offsite power.

Consequently, it can be concluded that ECCS pump operation for DBA LBLOCA calculations 

will not be affected by the presence of NCG in the suppression pool coming from downcomer 

flow following a DBA LBLOCA for a Mark I containment because the earliest pump start time of 

4 seconds is greater than the 3 second duration of the large gas bubble and large void fractions 

in the suppression pool.

During the initial period after receipt of a startup signal, the LPCS and LPCI pumps are operated 

to reach full speed. Additionally, the initial pump flow is recirculated to the suppression pool and 

is not injected into the core. Regardless of the availability of offsite power, the LPCS and LPCI 

systems are expected to pump flow into the core at about 40 and 50 seconds after appropriate 

valve reconfiguration. Consequently, during the first 40 to 50seconds following a LOCA, when

the ECCS pumps circulate flow within the suppression pool, the effects of NCG ingestion do not 

affect the ability of the ECCS pumps to supply flow to the core. The primary concerns in this 

early time frame pertains to the possibility of incurring pump damage or, if pump damage does 

not occur, the ability of the ECCS pumps to recover to full operation after possible NCG

ingestion during the early period. 

4.2 Acceptable ECCS Pump Operation

An ECCS pump can sustain damage by over speed and vibration if excessive amounts of NCG

are ingested. The following sections discuss the results of a literature search addressing 
30

acceptable pump operation and culminate with recommended criteria for determining 

acceptable BWR ECCS pump operation in the presence of NCG. 

4.2.1 ECCS Pump Performance in BWRs (NUREG/CR-2772) 

Testing reported in NUREG/CR-2772 (Ref. 39) finds that air concentrations greater than about 3 

to 5 percent by volume in a suction line of an ECCS pump can considerably lower the 

head-discharge curves of centrifugal pumps causing lower pump capacity at a given head.

4.2.2 Pump Performance under Air Ingesting Conditions (NUREG/CR-2792)

NUREG/CR-2792 (Ref. 40) assesses the performance of residual heat removal (RHR) and 

containment spray (CS) pumps during recirculation operation with sump suction following a 

postulated LOCA in a pressurized water reactor (PWR). The results of air/water tests on 

centrifugal pumps from several experimental programs were assessed to quantify the effects of 

air on head degradation. To apply the information in this report to BWR ECCS pumps, one must 

assume that conclusions developed for the assessed centrifugal pump operation tests can be 

applied to ECCS centrifugal pumps in BWR PS containments. Results for only three pump tests 

were considered applicable because the tested pump needed to meet the following 

requirements.

a. A pump-specific speed of 800–2000 was required. The criterion eliminated air/water

performance tests for primary coolant pumps and axial flow pumps. (Note, specific pump

speed is defined as N Q1⁄2 / H 3⁄4 where N is the shaft speed in rpm, Q is the volumetric

flow in U.S. gpm, and H is the pump differential head in feet.)

b. The test needed to be well documented.

c. The RHR and CS pumps surveyed had impeller pump diameter up to about 0.508 m

(20 in.). Data on test pumps with impeller diameters less than one-third this size were

not included.

d. The test fluids were required to be air and water.

Specifically, the report notes that the performance of centrifugal pumps is known to degrade 

with increasing gas content. It further notes that the amount of degradation is a function of 

various parameters; the important ones being pump design, specific speed, flow rate, inlet 

pressure and fluid properties. The report states that a general guideline adhered to by the pump 

industry is that degradation is not a concern at normal flow rates if air ingestion levels are less 

than 2 percent by volume; that acceptable pump performance with air ingestion levels between 

2 percent and 15 percent is dependent on pump design, and that centrifugal pump performance 

is fully degraded for air ingestion volumes greater than 15 percent. At low net positive suction 

head (NPSH) values close to the net positive suction head required by the pump (NPSHR), air 

ingestion will increase the degradation in performance in comparison to operation in the 

absence of air.

Because of a review of available test data, it is concluded that no case of severe pump 

degradation was observed for void fractions less than 2 percent and that degradation in 

performance does not occur until the pump inlet void fraction exceeds about 3 percent. Limited 
31

test data indicated that deterioration in performance will occur until the test pumps loses prime 

at air volume fractions between 7 percent and 15 percent. 

Discussions with pump specialists resulted in the opinion that pump inlet void fractions between 

1 and 3 percent would result in negligible degradation, that air binding might occur for flows less 

than 50 percent of best efficiency, that pump degradation between 3 and 15 percent is 

dependent on individual pump design and operating conditions, and that air quantities greater 

than 15 percent would result in full degradation in most pumps.

The following general conclusions were reached from the studies mentioned in the report: 

1. For a wide range of operating flow rates, RHR and CS pumps should handle volumetric

air quantities up to 2 percent with negligible degradation in performance.

2. For air quantities greater than 2 percent, pump performance degradation varies

substantially depending on design and operating conditions.

3. For very low flow rates (less than 50 percent of the best efficiency point), the presence of

air may cause air binding or loss of prime in the pump even at low-ingestion levels (less

than 2 percent by volume). However, for low flow rates, evaluations show that the

likelihood of air ingestion is low.

4.2.3 Pump Operation and Acceptable Gas Quantity (Generic Letter 2007-02) 

Generic Letter 2007-02 (Ref. 41) concludes from the data in NUREG/CR-2792 (Ref. 40) that the 

commonly used limit of 5 percent gas into ECCS pumps may be reasonable only if a substantial 

flow rate can be insured. At reduced flow rates gas can accumulate with time and the pump can 

eventually become gas bound with gas ingestion rates that are not normally a problem at higher 

flow rates.

4.2.4 Lappeenranta University Pump Experiments with Noncondensable Gas

As discussed in Section 3.1.5 Lappeenranta University performed scaled suppression pool 

experiments with NCG. This testing includes assessments of an ECCS pump representative of 

the Olkiluoto BWR plant PS system, which was considered typical of the Scandinavian designs. 

The general observations regarding ECCS pump performance tests in the presence of NCG are 

listed below.

1. In the pump tests, the ECCS pump was operating at a nominal speed of 2,970 rpm with

four different volumetric flow rates when pressurized air was injected directly into the

pump intake pipe. With volumetric flow rates of 57 and 75 l/s (903 and 1,189 gpm), a 3

to 4-percent air fraction was needed before a decline in pump head and flow was

observed.

2. With smaller flows of 12.5 and 25 l/s (198 and 396 gpm) the pump head and flow started

to decline after air injection was initiated and the flow totally collapsed when more than a

7-percent air fraction was present in the intake pipe.

3. When the air injection was stopped for the 12.5 l/s case, the pump head and flow

reached original values after 30 seconds after the air injection was turned off. At all the
32

other flows the pump head and flow normalized in a few seconds after the air injection 

was turned off.

These integral tests demonstrated that there is a relationship between the air discharge 

velocities through the downcomer vents and the amount of pump air ingestion. Consequently, 

an assessment of NCG effects on ECCS pump behavior must include the effects of the 

downcomer vent air flow and the flow behavior of the gas and water in the suppression pool. 

As discussed in Section 3.1.6, the Nordic Owners Group performed proprietary testing at the 

Lappeenranta University test facility when NCG was injected directly into the ECCS pump inlet.

The testing showed that the ECCS pump with inlet pipe void fractions below about 8 percent

may experience reduced flow depending on inlet void fraction, but the pump operation always 

fully recovered. The pump experienced a large decrease in pumping capability at inlet void 

fractions above 8 percent for durations of a few seconds; however, the pump always fully 

recovered.

It should be emphasized that these observations are typical of the Scandinavian BWR PS

system pump designs and may not be totally applicable to ECCS pumps used in US Mark I, 

Mark II and Mark III designs.

4.2.5 BWR Owners’ Group ECCS Pump Suction Void Fraction Study

Reference 42 states that “most BWR ECCS pumps are from three suppliers: Byron Jackson 

(now Flowserve), Sulzer, and Ingersoll (now part of Flowserve).” The reference further states 

that the pump types vary from single-stage centrifugal to 15-stage centrifugal pumps. This 

reference makes the following recommendations regarding ECCS pump operation.

1. Based on the review of pump gas intrusion data an upper limit bound of 2 percent by

volume continuous suction gas void fraction is acceptable. The report notes that the

acceptability of the 2 percent value is based on the test data and research compiled in

NUREG/CR-2792 (Ref. 40). The report further notes that test data suggest that little to

no pump degradation in pump performance occurs in the range of 0.5 percent to

5 percent with the most recurring values being 2 percent or less.

2. The report notes that NUREG/CR-2792 suggests that limited testing for multistage

pumps indicates that at a given air volume pump performance degradation is less

pronounced at higher inlet pressures and in multistage pumps.

3. As a guideline the report recommends that an ECCS pump will not incur damage and

will recover if a 10 percent average void fraction for no greater than 5 seconds exists. It

justifies this approach by referencing plant data that concluded that 3 to 5 percent air

entrainment by volume over a 20 second period would not be expected to cause any

significant problems.

4.2.6 University of Minnesota Centrifugal Pump Performance Tests with Air Ingestion

Reference 43 described full-scale testing to examine the effects of suspended air bubbles on 

the performance of a sea-water circulating centrifugal pump. The pump was a single suction, 

vertical discharge, horizontal suction pump. Its single-stage impeller was capable of delivering 

11.356 m3

/m (3,000 gpm) at a total dynamic head or 6.895 × 104 Pa (10 psi) at 1,150 rpm. The 
33

pump used a 0.2032 m (8 in.) diameter suction leg and 0.254 m (10 in.) diameter discharge leg. 

The suction pressure of about 6.895 × 103 Pa (1 psig) was used during testing. Air was injected 

into a vertical leg of the pump suction line through on injector nozzle.

The report states the following, “the tests established that a sudden loss of head developed 

across the pump which occurs between 4 and 5 percent volume concentration of air; however, 

the pump performance will reliably recover if the void fraction is reduced below 0.04.” It should 

be emphasized that these observations are typical of the tested centrifugal pump and may not 

be totally applicable to ECCS pumps used in US Mark I, Mark II and Mark III designs

4.2.7 Expert Panel Study of Gas Accumulation on Pumps – “The Pump Roadmap”

References 44 and 45 present the results of a study performed by panel of pump experts to 

develop acceptance criteria for pump operation with NCG ingestion. The study resulted in the 

BWR pump operational criteria summarized in Table 4.1-1. The results of this activity represent 

the latest and best information addressing acceptable pump operation in the presence of NCG. 

Consequently, the criteria provided by this panel will be used as the basis for assessing 

acceptable BWR ECCS pump operation. 

Table 4.1-1: Allowable Average Noncondensable Void Fractions to 

Preclude Damage to Typical BWR Pumps

Operating Condition % Q / Q(BEP) a, b Void Fraction c

Steady State Operation > 20 s 40%–120% 2%

Steady State Operation > 20 s < 40% or > 120% 1%

Transient Operation 70%–120 % 10% for ≤ 5 s

Transient Operation < 70% or > 120% 5% for ≤ 5 s

a Q is the water volumetric flow.

b BEP is best efficiency point.

c Transient void fraction is averaged over the specified time span.

The BWR ECCS operational criteria developed from the information in References 44 and 45

are summarized below.

1. Low transient flow rates (less than 70 percent of the best efficiency point) with an inlet

void fraction greater than 5 percent for more than 5 seconds will result in pump damage

unless pump-specific testing can show otherwise.

2. Higher transient flow rates (between 70 percent and 120 percent of the best efficiency

point) with an inlet void fraction greater than 10 percent for more than 5 seconds will

result in pump damage unless pump-specific testing can show otherwise.

3. Low flow operation (less than 40 percent of the best efficiency point) at greater than

1 percent by volume continuous suction NCG void fraction results in pump damage

unless pump-specific testing can show otherwise.

4. Full flow operation (40 percent to 120 percent of the best efficiency point) at greater than

2 percent by volume continuous suction NCG void fraction results in pump damage

unless pump-specific testing can show otherwise.
34

ECCS pumps that operate within the void fraction conditions specified in Table 4.1-1 and within 

the gas void fractions indicated above in items 1 to 4 can be considered capable of recovery 

without damage, and will operate at normal pump head and flow conditions if the inlet void 

fraction is reduced to a value below 1 percent. 

This criteria will be used as the basis for assessing acceptable BWR ECCS pump operation. 

Of course, an ECCS pump assessment would be permitted to use other operation criteria if 

acceptable pump-specific tests are performed that provide better operational guidance.
35

5. POST-LOCA SUPPRESSION POOL ANALYSIS

Two approaches were considered for determining suppression pool void fraction distribution 

following a LOCA with the intent to compare the results from both approaches for verification 

purposes or even to pursue a synergistic analysis that combines both approaches. The first

approach scaled the test results from the Lappeenranta University tests and the PUMA tests to 

full-scale plant geometry. The second approach used CFD calculations to define the NCG and 

fluid behavior in the suppression pool after verification of modeling methods using the scaled 

Lappeenranta University tests, the PUMA tests and qualitative verification against the Mark II 

full-scale test video.

5.1 Scaling of Test Data to Plant Conditions and Geometry

Scaling laws may be used to determine full-scale plant conditions from small-scale test results. 

Of particular interest here are test results from the Lappeenranta and Purdue Universities

scaled tests because both test programs provided results related to air and air/steam bubble 

size and behavior in a water pool following air and air/steam flow from a submerged pipe. The 

PUMA tests also measured void fractions in the water pool. As previously mentioned, the 

Lappeenranta University scaled tests were performed using downcomer IDs of about 0.1623 m 

and 0.213 m (6.39 in. and 8.4 in.), and Purdue University scaled tests were performed using 

0.078-m and 0.102-m (3-in. and 4-in.) ID downcomers. Because the PUMA tests provided 

measurements of air and air/steam bubble size and behavior as well as void fraction in the 

water pool, attempts were made to scale those small-scale test results to full-scale plant 

conditions. Data from other tests—such as the MIT tests, which used 0.0254 m, 0.0508 m and 

0.0983 m (1 in., 2 in., and 3.87 in.) ID vents, the University of California tests, which used 0.009

m to 0.095 m (0.35 in. to 3.7 in.) ID vents, and the PSI tests which used 0.05 m (1.97 in.) ID

vents—were not considered because these tests measured only the transient bubble size in the 

water pool; measurements of the void fraction distribution in the water pool are not available. 

For comparison, full-scale Mark I and Mark II downcomer inner diameter range from 0.5334 to 

0.6096 m (21 to 24 in.). 

5.1.1 Scaling Approach using Experiment Modeling

The small-scale modeling method rests on knowledge of the modeling laws which ensure that 

dynamic similarity exists between the model and the full-scale system. This modeling method 

should provide the scaling laws for predicting full-scale conditions from data derived from the 

small-scale model. This scaling approach was used for the pool swell process as part of the 

pressure suppression testing that was performed in the 1970s and 1980s. Because of the 

complexity of the suppression pool swell process, it would be almost impossible to achieve 

exact dynamic similarity. However, F.J. Moody (Ref. 3) developed a set of modeling laws that

were sufficiently simple to allow convenient small-scale modeling to be applied to a large-scale 

geometry. These laws became the basis for the industry’s small-scale tests.

The scaling method developed by A.A. Sonin and P.W. Huber (Ref. 46) for air clearing in 

suppression systems follows Moody’s approach but includes a more rigorous consideration of 

the vent system’s hydraulic resistance. An approach similar to the one used by used by Sonin 

and Huber for the pool swell problem was considered for the GI-193 scaling analysis. 
36

The scaling analysis focused heavily on the PUMA tests because a detailed description of the 

scaling approach was provided in the PUMA report. Unfortunately, the primary tests of interest,

the Lappeenranta University and Purdue University scaled tests, were based on much simpler 

scaling approaches. For example, the Purdue University tests were based on a scaling method 

that depended primarily on simple geometrical (e.g., length, diameter, area) ratios between the 

prototype and the model, and lacked the depth associated with the more rigorous and 

well-established approaches that are based on non-dimensional groups such as the Reynolds 

number and Froude number. Therefore, a more rigorous approach, such as Moody’s approach, 

cannot be used to directly scale those test results to full-scale plant conditions. Attempts to 

develop a specific scaling approach that follows Moody’s approach while still taking advantage 

of those small-scale test results were unsuccessful.

5.1.2 Scaling Approach using Downward Gas Jet Behavior

In a study of downward gas jets into a liquid by A. Emami and C. Briens (Ref. 47), the 

penetration length of downward nonreacting, subsonic gas jets in liquid was measured by optical 

methods. The effects of nozzle diameter, gas density, gas jet velocity, liquid viscosity, and liquid 

surface tension on jet penetration were examined. A generalized correlation to predict the 

penetration depth of downward gas jets was obtained. Also a simple model, using a shell 

momentum balance and assuming downward gas jet as a downward cone with constant angle, 

was introduced and validated with experimental data. All experiments were performed at room 

temperature and in a rectangular glass tank with dimension of 0.6 m (1.97 ft) length, 0.3 m 

(0.98 ft) width, and 0.4 m (1.31 ft) height which was filled up to 75 percent of its height with 

different liquids, such as distilled water, ethanol (95 percent), or sucrose solutions, to vary the 

liquid viscosity and surface tension. Helium, argon, and oil-free air were used as the gas phase. 

The gas mass flow rate for helium was 0.15 g/s (0.00033 lbm/s) and for air and argon were 

1.1 and 1.3 g/s (0.0024 and 0.00287 lbm/s) , respectively. The gas was introduced into the liquid 

through a submerged downward stainless steel tube. Four nozzle diameters were used. The 

inner diameters of the nozzles were 1.07, 2.03, 2.85, and 4 mm (0.042, 0.08, 0.112, and 

0.1575 in.). The length of stainless steel tube was 0.7 m (2.3 ft). Preliminary measurements 

were taken with a fast video camera. A photodiode cell and a horizontal laser beam were used 

to measure the jet penetration length. The average jet penetration length (L50%) was taken to 

correspond to a 50 percent probability of jet presence and a four parameter sigmoid curve, fitted 

to the experimental data, was used for interpolation. The span of the jet fluctuation range was 

taken as the differences between the distances corresponding to 10 and 90 percent jet presence 

probabilities (L10%–L90%).

When gas is injected to the liquid through a downward nozzle, the whole process can be 

described as follows: The gas flow produces a bubble, which keeps growing until the buoyancy 

force gets strong enough to overcome inertial forces. At this point, the bubble starts moving 

upward. Then the bubble detaches from the nozzle tip moving upward while covering the nozzle 

tube. At low gas flow rates a bubbling regime is dominant, and with increasing gas flow rate a 

fluctuation gas jet grows and the bubble size increases. The high-speed videos taken at different 

gas flow rates suggest that the cavity can be assumed as a downward cone.

This model, which is based on a momentum balance, assumes a conical shape for the 

downward gas jet. Momentum input comes from the gas entering the control volume at the 

nozzle tip. Momentum output is negligible because of the low bubble velocity leaving the control 

volume. The force acting on the control volume is the buoyancy force. The frictional force of the 

liquid on the control volume is negligible as the jet is assumed to be stationary (Figure 5.1-1).
37

This approach seems to be rigorous enough and is based on the appropriate nondimensional 

groups such as a modified Froude number. However, there are two points of concern: The first 

is that the nozzle diameters used in the tests were much smaller than the downcomer diameter, 

which may result in a different flow regime, and the second is that although the results may 

have accurately described the behavior of a gas jet into liquid, they may have missed the 

inertial effect of the initial clearing of the water column on the subsequent behavior of the gas 

jet. For these reasons, this alternate approach was considered inadequate for scaling the test 

results up to actual plant geometry and conditions.

Figure 5.1-1: Model for Downward Gas Jet in Liquid

5.1.3 Assessment of Scaling Approaches Using CFD Analysis

The evaluation of the two scaling approaches described above indicated that neither approach 

is adequate for accurately scaling test results to full-scale plant geometry and conditions. The 

PUMA approach was deemed inadequate because it was overly simplified. It is based on simple 

geometrical (e.g. length, diameter, area) ratios and not on basic nondimensional groups such as 

Froude number. The scaling approach by Emami and Briens, although more rigorous, was also 

deemed inadequate because the nozzle diameters used in the tests were much smaller than the 

downcomer diameter, which may result in a different flow regime, and also because the analysis 

ignores the important inertial effect of the initial clearing of the water column and its effect on the 

initial depth of the gas jet. 
38

The adequacy of both approaches was further investigated using CFD analysis. The 

investigation was performed as follows:

• One of the PUMA tests was used as a benchmark test to determine the adequacy of the

CFD approach. The analysis was based on the test geometry and dimensions, as well

as initial and boundary conditions. The test was simulated successfully and the CFD

model was considered adequate for these type of predictions.

• An idealized full-scale CFD case was developed from the PUMA test using the PUMA

scaling method. CFD analysis was performed using the idealized full-scale geometry

and conditions. The scaled results and CFD results did not match. This suggested that

the PUMA scaling approach was not appropriate.

• An idealized full-scale CFD case was developed from the same PUMA test using the

Emami and Briens scaling method, and a CFD analysis was performed using the

idealized full-scale geometry and conditions. The scaled results and CFD results did not

match. This again suggested that the Emami and Briens scaling method was not

appropriate for predicting the large scale behavior of the wetwell PS system.

A matrix of CFD cases was developed to further investigate the possibility of scaling the PUMA 

test results to actual plant conditions. The CFD runs included in this matrix were performed for 

various downcomer diameters (0.1 m, 0.2 m, and 0.6 m) and for various flowrates. The CFD 

sensitivity analysis results showed that because of distortion in the PUMA scaling approach, the 

PUMA test results may have fallen into a condition where the behavior of the jet was not

representative of the actual jet at full-scale conditions. Therefore, it was concluded that it was 

not possible to scale the PUMA test results to actual plant conditions. However, it is considered

appropriate to use the PUMA tests to generically benchmark the CFD methods used for 

performing the analysis of a downcomer jet in a full-scale plant geometry.

5.1.4 Conclusions Regarding Attempt to Scale Test Data to Plant Geometry

It was concluded that an approach to scale currently available test data for NCG behavior in a

water pool to full-size geometry was not feasible. This conclusion also eliminated the option to 

pursue a synergistic scaling/CFD analysis effort and the ability to compare both independent 

results for verification purposes.

5.2 CFD Analysis of Suppression Pool

CFD analysis techniques have been used to simulate the vent clearing and NCG flow exiting a 

downcomer vent. The CFD technique has been used to simulate the LLL one-fifth Mark I test 

facility tests, the MIT scaled Mark I NCG blowdown tests, the University of California scaled 

Mark I tests, the scaled test results performed at Lappeenranta University, and the PSI 

small-scale tests. The CFD analyses for all but the Lappeenranta University and PSI analyses 

were performed using older CFD codes or using a CFD code specifically developed to analyze 

the Mark I NCG phenomenon. The CFD methods have improved since the older codes were 

used about twenty years ago. Recently Lappeenranta University has successfully used the CFD 

Fluent code to analyze the water clearing and NCG phenomena in the suppression pool. 

Similarly, PSI has used current CFD calculational methods to analyze its small-scale test data.
39

PSI has performed two-dimensional CFD analyses of air and steam-air flow into a water pool. 

The Lappeenranta University staff is also performing work to extend the CFD analysis to 

simulate the steam-noncondensable gas behavior in the suppression pool. This effort involves 

developing the direct condensation behavior on the bubble surface in contact with the 

suppression pool water. This new analysis is state-of-the-art work and is not recommended for 

general use at this time.

Considering the experience with CFD analyses for analyzing the vent clearing and NCG portion 

of the downcomer vent flow into the suppression pool following a LOCA, it is reasonable to 

assume that a CFD analysis could be used to analyze the initial vent clearing and NCG flow 

transient in a full-scale plant geometry.

CFD is being used to study the initial blast of NCG that is blown into the suppression pool 

following a LOCA in the containment of a Mark I BWR. This rapid blowdown of gas through a 

vertical downcomer pipe submerged in the suppression pool creates a large NCG region that 

spreads down and outward into the suppression pool. A safety concern relates to the possibility 

of the NCGs being pulled into the ECCS pump intakes. This could cause a significant 

degradation in pump performance or even pump failure. The pump intake system is submerged 

in the suppression pool and in some cases may be in close proximity to the downcomer pipes. 

The size, depth, and transient nature of the NCG region must be considered along with the 

intake locations, pump startup times, and pump performance characteristics to assess the 

potential effect of these scenarios. The present CFD analysis is focused on the issue of 

predicting the size, depth, and transient nature of representative gas bubbles that are blown 

down through a single downcomer pipe into the surrounding region of a suppression pool.

To build confidence in the modeling approach, the CFD methods are benchmarked using some 

related test data. The first test considered comes from the PPOOLEX test facility in Finland, 

which is described in Reference 49. Prior CFD analyses of these tests in Reference 26 are used 

as a foundation to build the current modeling approach. These prior analyses demonstrated the 

applicability of the VOF method for predicting the noncondensable bubble dynamics in a scaled 

suppression pool environment.

The second set of test data comes from the NRC-sponsored testing at the PUMA facility 

described in Reference 33. Three tests using air injected through a submerged 0.102 m (4 in.)

downcomer pipe in a suppression pool are considered. The initial air injection period results in 

the maximum void penetration depth and spread. This initial phase of the blowdown is the focus 

for the analysis.

Full-scale test data and videos are considered from the steam-driven GKSS PS test facility in 

Reference 48. The facility setup and test data, however, are more focused on later stages of the 

blowdown event where steam condensation induced oscillation and chugging issues are of 

interest. Quantitative measurements of the size and depth of the initial large noncondensable 

bubble formed during the blowdown are not available from these tests. The videos do provide a 

qualitative understanding of the complexity of the initial blowdown event at full-scale conditions 

including the formation of many smaller bubbles that break away from the main bubble.

The verified CFD techniques were used to develop models of actual plant geometries and 

simulate the NCG behavior in a plant’s suppression pool. This analysis could be used to define 

the NCG “exclusion zone” for plant geometries and to help define the void fraction at ECCS 
40

pump strainer locations. The analysis would also help to define the time duration during which 

local void fraction effects at the pump strainer would be important.

5.2.1 PPOOLEX Tests

5.2.1.1 Background

The PPOOLEX facility consists of a 31 m3 (1,095 ft3

) cylindrical tank with rounded end caps. 

The height and diameter are approximately 7.5 and 2.4 m (24.6 and 7.87 ft) respectively. The 

facility is divided into two regions by a plate that separates an upper (drywell) region from the 

lower (wetwell) region. A vertical blowdown pipe connects the upper and lower regions. This 

pipe has an inner diameter of 0.2141 m (8.43 in.) and is submerged in water that partially fills 

the lower wetwell region. During the tests, pressurized air is forced into the upper drywell region 

through an inlet plenum. The drywell region pressurization forces air down through the vertical 

blowdown pipe. After the water is pushed out of the lower end of the blowdown pipe, the air 

begins to exit and gas bubbles are formed in the wetwell region. The prediction of the size, 

shape, and duration of this gas region is the focus of the benchmark study. A more complete 

description of the test is contained in the Finnish report documenting some earlier CFD work on 

these tests (Ref. 26). An overview of the facility is provided in Figure 5.2.1-1 which is copied 

from the test report.

Figure 5.2.1-1: Overview of PPOOLEX Test Facility (Ref. 26) 

Two tests from the PPOOLEX facility are considered for a CFD benchmark study. These tests 

are identified as CHAR-09-01 and test CHAR-09-03. The tests differed by the inlet plenum mass 

flow rates with test CHAR-09-03 having nearly three times the mass flow of test CHAR-09-01.
41

5.2.1.2 CFD Approach

A CFD model is created for the PPOOLEX test facility using the FLUENT v15.0 code. The 

model is setup using an approach similar to the one outlined in the 2008 Finnish report. That 

prior study also used the FLUENT code. The VOF method is used to track the interface 

between the air and the water as the bubble exits the blowdown pipe and rises through the 

wetwell. The VOF method solves a single set of momentum equations and tracks the interface 

between two non-mixing fluids. Each control volume in the model is filled with either one of the 

two fluids or is considered part of the interface. It is assumed that the interface dimensions are 

larger than an individual control volume and bubbles smaller than the CFD control volumes are 

not tracked. The method is appropriate for tracking large bubbles or interfaces in a liquid and 

can support compressible gases. The gas bubble is assumed to be noncondensable in this 

case. Table 5.2.1-1 lists the basic modeling options used.

Table 5.2.1-1: CFD Model Basic Solver Settings

Code ANSYS/FLUENT v15.0

General pressure based transient solver, gravity

Models VOF Method, Implicit volume fraction scheme, energy equation

Turbulence standard k-epsilon, standard wall functions, compressibility option

Materials water (primary phase)

air, ideal gas assumption (secondary phase)

Conditions floating operating pressure, specified operating density = 0.0

Boundary 

Conditions

inlet – specified mass flow, constant temperature

walls – adiabatic, no slip condition

P-V coupling SIMPLE method

Discretization density, momentum, energy

turbulence – first order

pressure – PRESTO

time – second order (time step = 1/300 s)

5.2.1.3 Geometry/Mesh

The geometry is obtained from the Finnish report detailing the PPOOLEX experiments (Ref. 8). 

Small geometric details are omitted for this study because they are not expected to affect the 

gas bubble dynamics in the wetwell. The components included in the model are the main outer 

pressure vessel, the intermediate floor dividing the wetwell and drywell, the inlet plenum pipe, 

and the blowdown pipe. The dimensions and locations of these components are taken directly 

from the test report. All walls and surfaces are modeled with zero thickness. No solid structures 

are considered. Figure 5.2.1-2 illustrates the geometry used to define the CFD modeling 

domain. The cylindrical section of the large tank has an inner diameter of 2.38 m (93.7 in.) and 

the blowdown pipe has an inner diameter of 0.2141 m (8.43 in.). The blowdown pipe is slightly 

off the vertical axis of the cylindrical tank. Overall height of the facility is roughly 7.4 m (291 in.). 

The meshing refinement regions illustrated in Figure 5.2-2 are not part of the geometry. These 

regions are simply defined to help specify the mesh sizing in regions of interest.

Figure 5.2.1-3 illustrates the computational mesh design on a vertical plane. The mesh is 

created using the ANSYS meshing platform “cut cell” option which uses cubes to mesh the 
42

entire domain. The dimensions of the cubes are successively cut in half for grid refinement. At 

walls, the cubes are skewed to maintain the outer wall shape. During the creation of the model 

several mesh designs were used and refined on an iterative basis to ensure the grid refinement 

regions covered the areas of interest for the air to water interface. The bulk regions of the model 

used cell dimensions of 0.08 m (3.15 in.) as a maximum size. The upper portion of the 

blowdown (downcomer) pipe was meshed using a cell size of 0.04 m (1.575 in.). The refinement 

regions which covered the initial water surface region, the submerged portion of the blowdown 

pipe and a region around this pipe were meshed with 0.01 m (0.394 in.) cells. This mesh design 

results in a total of 2.5 million computational volumes with the majority of these located in the 

mesh refinement regions.

Figure 5.2.1-2: CFD PPOOLEX Geometry Figure 5.2.1-3: PPOOLEX Computational Mesh

5.2.1.4 Boundary and Initial Conditions

The boundary and initial conditions are obtained from the PPOOLEX test report and the 

associated Finnish report documenting the prior CFD analyses. Tests 09-01 and 09-03 are 

considered because they span a wide range of mass flows. The interest in this study is the 

downward and lateral penetration of the initial gas bubble exiting the blowdown pipe. For this 

reason, only a short period of time is modeled and longer term effects such as wall heat transfer 

are not considered. Table 5.2.1-2 lists the boundary and initial conditions applied for each test. 

Only the mass flow of air is varied during the simulation and this is illustrated on Figure 5.2.1-4.
43

Table 5.2.1-2: Boundary and Initial Conditions for PPOOLEX Predictions

Condition Test 09-01 Test 09-03

wall heat transfer adiabatic adiabatic

initial water temperature 297.1 °K (75.11 °F) 296.3 °K (73.67 °F)

initial air temperature 297.1 °K (75.11 °F) 296.3 °K (73.67 °F)

initial pressure 103,390 Pa (14.995 psi) 103,890 Pa (15.068 psi)

initial velocities 0.0 0.0

inlet air temperature 297.1 °K (75.11 °F) 297.1 °K (75.11 °F)

inlet mass flow time-dependent table time-dependent table

inlet turbulence estimation

5% intensity

Hydraulic diameter 0.2 m 

(0.656 ft)

5% intensity

Hydraulic diameter 0.2 m 

(0.656 ft)

Figure 5.2.1-4: Mass Flow at Inlet for PPOOLEX Tests

5.2.1.5 Results

Only limited data are available from the PPOOLEX tests. Pressures and temperatures were 

measured, and there is video of the bubbles around the downcomer exit. The focus of this study 

is the penetration of the gas bubble into the wetwell liquid. Both depth and width of the gas 

bubble are of interest. Comparisons are made between the CFD predictions and the video to 

qualitatively compare the CFD predictions of bubble size and duration. Quantitative 

comparisons of predicted system pressures are used to demonstrate that the overall CFD 

boundary conditions and predicted system response are in line with the test. Temperature 

comparisons are not included because the duration of the first bubble is relatively short and 
44

temperature variations are minimal. It is noted that initial temperatures are established to match 

the test data. 

Figure 5.2.1-5 shows contour plots that highlight the predicted bubble size and location 

compared to test video frames at the same time intervals for test 09-01. The first frame is 

selected to correspond to the time when the bubble first starts to leave the blowdown pipe which 

is 10 seconds for test 09-01. Six video frames are shown alongside the CFD predictions. The 

original video frame rate is approximately 1

/30th of a second and every fourth video frame is 

shown. Simulation time is indicated on each set of figures. The qualitative results indicate that 

the CFD predictions adequately predict the bubble size, shape and duration. These results 

illustrate the first bubble that exits the blowdown pipe after the initiation of the mass flow. The 

simulation is carried out further to simulate the release of several more large bubbles into the 

wetwell region. The predictions matched the frequency of bubble formation for the duration of 

the simulation considered. This is a good indication that the boundary conditions for the CFD 

model are a good representation of the test conditions.

Figure 5.2.1-5: Comparison of Predicted Bubble Size and Location for Test 09-01

Pressures are monitored in the CFD prediction and compared to the test data for points in the 

drywell and wetwell. Figure 5.2.1-6 shows a comparison with three of the measured test 

pressures during the first 20 seconds of test 09-01. Pressures are considered in the upper 

drywell, the airspace in the wetwell, and a point submerged in the wetwell. The pressurization in 

the upper drywell (P1101) is well predicted by the CFD code. The initiation of the pressure rise 

and the total pressure rise are a good indication that the inlet boundary mass flow is properly 
45

accounted for. The wetwell pressure (P04) is also predicted well. The initiation of the pressure 

rise around 10.0 seconds corresponds to the time when the gas bubble begins to emerge from 

the blowdown pipe. The rate of pressure rise after 10 seconds is a good indication that the mass 

flow of air exiting the blowdown pipe is accurately predicted. The predictions show a small 

upward bias in the pressure which remains consistent throughout the test. This could be the 

result of a reduced pressure at the start of the test (slight vacuum), which was not accounted for 

or a simple bias in the measurement. Overall the predictions do a good job of following the 

pressure from the experiment. A third comparison is made for the pressure measured near the 

exit of the blowdown pipe (P05). This pressure rise is similar to the dry measurement (P04) in 

the wetwell region but includes the hydrostatic pressure of the water above it. The predictions 

match the test data. The pressure response predictions indicate that the CFD model is setup 

correctly to model the incoming mass flow and that the timing and mass flow rates for gas 

entering the wetwell region is accurately represented.

Figure 5.2.1-6: Comparisons of Predicted and Measured Pressures, Test 09-01

Test 09-03 has a mass flow rate that is approximately 300 percent higher than the mass flow in 

test 09-01. A similar CFD modeling approach is used for test 09-03 and similar results are 

obtained. Figure 5.2.1-7 shows contour plots that highlight the surface of the gas bubble exiting 

from the blowdown pipe compared to images taken from the test video. Once again, the CFD 

model does a good job of predicting the size, shape, and timing of the gas bubble blown down 
46

into the wetwell region. There appears to be a slight tendency for the CFD model to predict the 

bubble to rise faster than the test video indicates but the difference is minor. The parameters of 

interest, depth and width of the bubble, are qualitatively well predicted as observed from the 

comparison with the video.

Figure 5.2.1-7: Comparison of Predicted Bubble Size and Location for Test 09-03

Pressures are monitored in the CFD prediction and compared to the measured test data for 

points in the drywell and wetwell. Figure 5.2.1-8 shows a comparison with three of the 

measured test pressures during the first 8 seconds of test 09-03. Pressures are considered in 

the upper drywell, the airspace in the wetwell, and a point submerged in the wetwell. The 

prediction of the pressurization in the upper drywell (P1101) matches the data very well. The 

trend for the wetwell pressure (P04) is also predicted well but there appears to be a slight shift 

in the initial pressure value and the timing for the start of the pressure rise. It is suspected that 

there is uncertainty in the timing of the rapidly changing boundary condition for the inlet mass 

flow during the initial phase of the transient and this could account for the difference in pressure 

rise time. The rate of the pressure increase is similar between the test and the predictions. The 

initiation of the wetwell pressure rise at around 4 seconds corresponds to the time when the gas 

bubble emerges from the blowdown pipe. A third comparison is made for the pressure 

measured near the exit of the blowdown pipe (P05). This pressure prediction also shows a slight 

delay in the pressure response compared to the test data but the rate of pressure rise is similar. 

Overall, the pressure response predictions indicate that the CFD model is doing a reasonable 

job of accounting for the incoming mass flow for the purposes of determining the bubble 

dynamics in the wetwell.
47

Figure 5.2.1-8: Comparisons of Predicted and Measured Pressures, Test 09-03

5.2.1.6 Conclusions

These PPOOLEX tests provide a valuable data set for benchmarking the capabilities of a CFD 

code to predict the general size and shape of NCG bubbles exiting from a submerged blowdown 

pipe. The VOF method used in the FLUENT code does an adequate job of predicting the overall 

bubble size, shape, and rise time based on the qualitative comparisons made with test videos. 

Pressures within the system are well predicted, which is a good indication that the CFD code is 

accurately representing the mass flows into and through the system. 

One limitation of these tests is the minimal penetration depth of the gas bubble. The NCG exits 

the pipe and quickly turns upward towards the surface. Penetration depths are not measured 

directly but are estimated at 0.2 L/d (length to pipe diameter) and 0.5 L/d below the pipe outlet 

from the video for tests 09-01 and 09-03 respectively. This bubble depth is small compared to 

prototypical penetration depths at full-scale conditions where bubble penetration depths of 

several pipe diameters are expected. In addition, the NCG bubble remains well defined with little 

or no bubble break up. The limited penetration depth and bubble break up is not prototypical for 

a large-break reactor accident scenario.
48

5.2.2 PUMA Tests

5.2.2.1 Background

Building upon the lessons learned from the PPOOLEX modeling effort, a second test series 

from the PUMA test facility is modeled. The NRC sponsored this test program, and multiple 

blowdown tests are available for consideration. A few select tests from the steady-state test 

series are chosen for modeling to provide a wide range of conditions. Some scaling is included 

in the test report as the basis for the pipe sizing and water depth. Details of the test program, 

facility, and scaling analysis are found in the NUREG/CR-7186 (Ref. 33) documenting the 

testing.

Figure 5.2.2-1, which comes from the test report, illustrates some of the key dimensions of the 

test vessel. The 2.74 m (107.9 in.) ID vessel is a little over 3 m (118.1 in.) high. The initial water 

level for the tests is 1.05 m (41.3 in.) as measured from the inside base of the tank. The 

downcomer pipe is 0.102 m (4 in.) in diameter and is submerged 0.37 m (14.6 in.) (roughly 

3.6 pipe diameters). A larger vent pipe, which is unused for these tests, is in place on the center 

line of the facility. The downcomer pipe is offset 0.68 m (26.8 in.) from the center of the tank to 

provide space for the bubble to expand. 

Figure 5.2.2-1: Overview of PUMA Vessel Used for Tests

There is a support structure built around the pipe exit to hold instrumentation for measuring the 

void fraction. Although the instrumentation probes are small, the support structure is relatively 

large, and there is a concern that this structure will break up the incoming air bubble as it 

expands out of the downcomer pipe. This is a qualitative observation only. Figure 5.2.2-2 shows 
49

three frames of a test video before, during, and after the first bubble emerges from the pipe. The 

support structure is clearly shown in the figure, and it clearly interacts with the bubbles at 

various elevations. 

Figure 5.2.2-2: Video Frames Showing Downcomer Pipe and Instrumentation Supports

Support rods are placed below the downcomer exit directly in the path of the gas bubble as 

shown in the middle image of Figure 5.2.2-2. Dimensions and precise locations are not available 

for the support rods but the size of each rod appears to be around 8–10 percent of the 

downcomer diameter. The two rods crossing directly under the downcomer exit create a flow 

obstruction that is expected to add turbulence to the incoming flow and potentially affect the 

bubble dynamics. The PUMA tests showed a significant amount of small bubbles breaking away 

from the main bubble when compared to the PPOOLEX tests, which showed no significant void 

fraction away from the main bubble (see Figures 5.2.1-5 and 5.2.1-7). The right side of 

Figure 5.2.2-2 shows the region below the downcomer exit after the initial large bubble has 

risen. A cloud of small bubbles is left behind by the large rising bubble. These bubbles are 

observed to slowly rise back to the surface.

The observations of the small bubbles in the PUMA test are only qualitative. No direct 

measurements of these bubbles are included in the test report. They are noted here because 

the origin of the small bubbles is not clear. Although no small bubbles are observed in the 

PPOOLEX tests, the PUMA tests appear to be more dynamic based on the significantly larger 

bubble penetration depth. The deeper penetration of the main bubble (several pipe diameters) 

and the more dynamic nature (higher turbulence) of the PUMA tests is expected to result in 

more small bubbles. The support rods are noted as a possible source of additional turbulence 

and bubble break up. This observation is made to help understand the results and potential 

uncertainties in the results.

5.2.2.2 CFD Approach

The CFD model is created for the PUMA test facility using the FLUENT v15.0 code. The model 

is setup using the same VOF modeling approach applied successfully to the PPOOLEX tests 

outlined above. As before, the gas is noncondensable and only the initial bubbles are 

considered. See Table 5.2.1-1 for a list of the basic modeling options. 
50

5.2.2.3 Geometry/Mesh

The geometry is obtained from the NUREG/CR-7186 report (Ref. 33), which details the PUMA 

facility experiments. Small geometric features are omitted to simplify the mesh design. The

components included in the model are the main outer pressure vessel, the central vertical vent 

pipe (which is not used for these tests), and the downcomer (blowdown) pipe. The dimensions 

and locations of these components are taken from the NUREG/CR-7186 report. All walls and 

surfaces are modeled as zero thickness walls. No solid structures are considered. Small 

flanges, and other supporting structures, such as the instrumentation support rods near the 

downcomer exit, are not included in the model. The piping upstream of the vertical downcomer 

pipe is also simplified.

Figure 5.2.2-3 illustrates the geometry used for the CFD model. The vertical downcomer 

(blowdown) pipe is modeled in the position identified by the drawings. The test report identifies a 

0.0508 m (2 in.) line that connects the 0.102 m (4 in.) downcomer pipe to the pressure vessel 

that supplies the gas for the test. The inlet boundary condition is established at the approximate 

location of the mass flow measurement on the 2 in. line. A horizontal length of 0.3048 m (12 in.)

is used for the 0.0508 m (2 in.) line in the model to represent the flow path and volume between 

the mass flow measurement location and the entrance to the vertical downcomer pipe. Valves 

and elbows are not included in the CFD model of the inlet piping. The mesh is created using the 

same cubic mesh design (ANSYS cut-cell option) used for the PPOOLEX models. This 

approach uses a cubic mesh array with mesh refinements made by reducing the cube 

dimensions successively by one half. Hanging nodes (non-conformal mesh design) on the plane 

of the transition are a result of this approach. Transition regions are setup away from areas of 

interest to create a uniform mesh in the areas of interest. The main area of interest is the region 

below and around the downcomer exit. Other regions for refinement include piping and the 

water surface. Figure 5.2.2-4 shows a mesh cross section on a vertical plane of the model 

which includes the central vertical vent as well as the downcomer pipe. The mesh consists of 

2.96 million cells.

Figure 5.2.2-3: CFD PUMA Geometry Figure 5.2.2-4: PUMA Computational Mesh
51

5.2.2.4 Boundary and Initial Conditions

Boundary and initial conditions are obtained from the PUMA test report. PUMA tests labeled as 

“steady state” test A1, A2, and A3 are considered because they provide a range of mass flow 

rates. The interest in this study is the downward and lateral penetration of the initial gas bubble 

exiting the blowdown pipe. For this reason, only a short period of time is modeled and longer 

term effects, such as wall heat transfer, are not considered important. Table 5.2.2-1 lists the 

boundary and initial conditions for the CFD simulations. Only the mass flow of air into the inlet 

pipe is variable. These flow rates are plotted in Figure 5.2.2-5.

Table 5.2.2-1: Boundary and Initial Conditions for PUMA Models

Condition A1 A2 A3

wall adiabatic adiabatic adiabatic

initial water temperature 296.4 K (73.85 °F) 296.2 K (73.49 °F) 297.3 K (75.47 °F)

initial air temperature 296.4 K (73.85 °F) 296.2 K (73.49 °F) 297.3 K (75.47 °F)

initial pressure 102,440 Pa

(14.858 psi)

102,440 Pa

(14.858 psi)

102,470 Pa

(14.862 psi)

initial velocities 0.0 0.0 0.0

inlet air temperature 288 °K (58.73 °F) 288 °K (58.73 °F) 288 °K (58.73 °F)

inlet mass flow time-dependent time-dependent time-dependent

inlet turbulence 

estimation

5% intensity;

hydraulic diameter 

0.05 m (0.164 ft)

5% intensity;

hydraulic diameter 

0.05 m (0.164 ft)

5% intensity;

hydraulic diameter 

0.05 m (0.164 ft)

Figure 5.2.2-5: Mass Flows at the Inlet Pipe Boundary for the PUMA Tests

5.2.2.5 Results

The PUMA testing involved limited measurement locations and a high-speed video that 

captured bubbles emerging from the downcomer pipe. A qualitative comparison is made for 

each of the tests with the video. Figures 5.2.2-6, 5.2.2-7, and 5.2.2-8 show selected snapshots 

in time for tests A1, A2, and A3 respectively. Each test is compared to contour plots of void 

fraction from the CFD predictions at equivalent times. The contour plots show the void fraction 
52

in a range from 0.4 to 0.6 which highlights the interface region between liquid and gas in the 

prediction. Times shown on the figures are from the CFD simulation. The test video times did 

not correspond to the actual test times. The time scale on the video was shifted so that the 

timing of the first emergence of gas from the downcomer pipe was identical for the CFD 

prediction and the test video. 

Figure 5.2.2-6: Comparison of Test Video and CFD Predictions for Test A1
53

Figure 5.2.2-7: Comparison of Test Video and CFD Predictions for Test A2

Figure 5.2.2-8: Comparison of Test Video and CFD Predictions for Test A3
54

The video comparisons show the dynamic bubble response and demonstrate how quickly the 

first large bubbles expand down into the pool and then rises again towards the surface. In just a 

little over 0.3 seconds, the main bubble exits the downcomer pipe, plunges nearly three pipe 

diameters into the pool, and then rises up past the downcomer pipe exit. As illustrated by the 

video images, when the main bubble rises up towards the surface, it leaves behind a trail of 

smaller bubbles which rise more slowly. These smaller bubbles are also affected by any overall 

fluid motion in the pool.

The tests generally show an initial large bubble forming and then, after a slight pause, another 

large bubble forms above the first. The entire region is generally connected and begins to rise 

quickly with the upper bubble still expanding as gas exits the downcomer pipe. The initial bubble 

follows the second bubble upward and is broken down as it joins the first leaving behind a 

smaller set of bubbles in its wake.

The CFD predictions qualitatively match the test video. The same general behavior is observed 

with an initial bubble penetrating down approximately 3 pipe diameters and then another large 

bubble forming above it just below the downcomer exit. As the bubbles rise, there is a small 

amount of gas left in the wake. The CFD predictions using the VOF method are not able to 

predict the cloud of small bubbles observed in the tests because of the resolution of the 

computational grid. It is encouraging, however, that an overall similar behavior is predicted.

The depth, spread, general structure, and timing of the bubbles are well predicted based on the

comparisons with test videos. Details of the bubbles outer structure are slightly different, and 

this is not surprising given the uncertainty in the rapidly changing blowdown boundary 

conditions and the absence of the support structures in the CFD model. Another important 

factor in the bubble dynamic behavior is the piping upstream of the downcomer exit. Sensitivity 

studies show that changing the pipe length has an effect on the bubble dynamics. It appears 

that as the bubble begins to form and exit the pipe, the pressure drops in the piping and the 

mass flow exiting the pipe is dependent upon the total volume of pipe upstream of the exit as 

well as any flow restrictions. The final predictions are made with a simplified piping system that 

approximates the upstream length and volume to the mass flow measurement location, which 

serves as the inlet boundary condition.

The video comparisons show qualitatively that the CFD predictions do a reasonable job of 

predicting the bubble dynamics and structure. Quantitative comparisons of bubble dynamics are 

also made using the available data. The test included void fraction measurements at a few 

selected locations. These measurements, however, do not provide the resolution needed to 

quantify the bubble. The test report also included a quantitative reading of the bubble 

penetration depth obtained by scaling results off of the video frame. These measurements are 

compared to similar measurements made from the CFD predictions. Figure 5.2.2-9 shows the 

penetration depth for tests A1, A2, and A3 as a function of time. The CFD predictions are read 

from contour plots similar to the example shown from test A2 in the middle plot on Figure 5.2.2-

9. A grid array is overlaid on the image showing L/d lines. All data are reported in the

normalized L/d scale.
55

Figure 5.2.2-9: Penetration Depth vs Time Comparisons for PUMA Tests

The results in Figure 5.2.2-9 indicate that the CFD model does a good job of predicting the main 

bubble penetration depth and rise time. Several observations should be made to help qualify the 

results. The PUMA test data readings from the video only focused on the main bubble region. 

The small bubbles that appear in the wake behind the rising large bubble are not considered in 

the test report. Another consideration is that scaling depth from video frame images can result in 

some measurement uncertainty. A complication is the identification of the lower edge of the 

main bubble when there is an attached wake of smaller bubbles. During this time, there is 

uncertainty in the precise location of the lower interface of the main bubble. The video field of 
56

view is also an issue. The vantage point is above the lowest point of the main bubble and a 

direct view of the lower edge of the main bubble may be obstructed by the bulging sides of the 

bubble region. In addition, the facility view port window blocks off visibility for anything below 

approximately -3.2 L/d. It appears from the videos of these tests that the main bubble does not 

penetrate further below this depth. Observations of smaller bubbles below this depth, however, 

are not possible. There is also no instrumentation at these lower levels.

The CFD predictions do a good job of predicting the bubble penetration depth. This is especially 

true while the bubble is descending and only the main bubble is present. During the period 

when the main bubble is rising, the location of the lower edge is difficult to determine from the 

CFD predictions. This is because of the difficulty in identifying the interface between the main 

bubble and the attached wake. The observations made when the bubble penetration depth is 

growing are generally more accurate than the predictions associated with the period of time 

when the bubble is rising with an attached wake.

The CFD method used does not resolve small bubbles that break away from the main bubble 

but does predict a small contiguous region that detaches from the main bubble in a manner 

similar to the test observations. This region not only forms in the wake of the large bubble but 

also continues downward during the initial penetration forming a small amount of NCG that 

penetrates towards the floor of the pool (at L/d = -6). This region is tracked in the CFD results 

and plotted in Figure 5.2.2-9 as the “small-bubbles.” These small bubbles do not emerge from 

the main bubble until it begins to rise. The smaller bubbles do not rise quickly and are affected 

by the global liquid flow patterns in the pool.

An interesting result is the relative consistency of the bubble penetration depth for all three 

tests. The different mass flow rates for each test do not affect the penetration depth, which is 

just over 3 pipe diameters for each test. The bubble duration is also similar for each test. Within 

0.4 to 0.5 seconds, the bubble exits the downcomer pipe, penetrates downward 3 L/d, and rises 

up past the downcomer exit. In all three tests considered the predicted small bubble region 

extended to the floor of the facility at L/d = -6.

The width of the NCG region is similar for all three tests. Tests A1 and A2 extend outward 

approximately 2.25 L/d at the maximum point below the downcomer exit. Test A3 extends 

outward approximately 2 L/d at the maximum spread while the main bubble is below the 

downcomer exit. The test results include bubble spread measured at one fixed elevation with 

limited instrumentation. These limited instruments do not capture the maximum bubble spread 

which is observed at a different elevation. Direct comparisons with the test data are not 

attempted. 

5.2.2.6 Conclusions

These PUMA tests provide benchmark test data for assessing the capabilities of a CFD code to 

predict the general size and shape of NCG bubbles emerging from a submerged blowdown 

pipe. The VOF method used in the FLUENT code does an adequate job of predicting the overall 

bubble size, shape, and rise time based on the comparisons made with test data. The good 

comparisons suggest that the applied CFD approach is adequate for predicting bubble 

penetration depths and rise times.

One limitation of these tests is the structural blockages associated with the instrument support 

rods. These rods potentially influence the bubble shape and break up characteristics. Another 
57

limitation is the lack of measurements for the small bubbles that break away from the main 

bubble and the lack of visibility for bubbles approximately 3.2 L/d (and deeper) below the 

downcomer exit. Penetration depths for all three tests (three different mass flow rates) are 

similar at just over -3 L/d. 

5.2.3 Full-Scale Simulations

5.2.3.1 Background

The methods used to simulate the two small-scale test programs are applied to a full-scale 

scenario based upon MELCOR predictions of a postulated large-break LOCA (recirculation line 

break scenario considered) in a BWR Mark I design discussed in Section 2 and Reference 2. 

The MELCOR analysis provides the drywell and wetwell initial conditions, the predicted drywell 

and wetwell transient pressures and temperatures, as well as the predicted mass flow rates 

through the downcomer vent system into the wetwell region (see Figures 2.1-4, 2.1-5, 2.1-6 and 

2.1-7). These values are used, along with a simplified representation of a BWR suppression 

chamber geometry to make predictions of the NCG bubble penetration depth and duration after 

exiting the blowdown pipe. The BWR Mark I geometry is used because it represents one of the 

smaller wetwell (suppression chamber) regions with some designs having pump strainer 

screens in close proximity to the downcomer pipe.

The large-break LOCA scenario is considered for this analysis because it is predicted to 

develop the highest drywell pressures. The boundary conditions used as well as flow paths, flow 

areas, and volumes are representative of a typical BWR design. These values could be different 

for other Mark I BWR designs. Figure 5.2.3-1 shows an overview of a Mark 1 BWR containment. 

The area of interest for this study is the suppression chamber.

Figure 5.2.3-1: Generic BWR Mark I Containment Layout 
58

5.2.3.2 CFD Approach

An initial CFD model is created for the full-scale simulations using the modeling approach 

demonstrated for the scaled test facilities above. The basic solver settings are outlined in 

Table 5.2.1-1. It is assumed that the downcomers and drywell to torus vent piping will be initially 

filled with liquid below the suppression pool water level and with NCG above the water column. 

Consequently, after the liquid is expelled from the downcomer, the subsequent vent flow is 

assumed to be 100 percent NCG. The assumption of 100 percent NCG is expected to provide a 

conservative prediction of the size and duration of the NCG bubble in the suppression pool. 

In comparison, for the MELCOR predictions shown in Figure 2.1-7, the initial flow from the 

drywell to wetwell is a mixture of steam, water droplets, and NCG. As shown if Figure 2.1-8, the 

initial vent flow is calculated to contain a maximum NCG mass fraction of about 73 percent. The 

mass fraction of NCG in the vent flow is predicted to rapidly drop below 10 percent after 

2 seconds following the start of the LOCA. 

It is suspected that many small bubbles may break away from the main bubble during the initial 

blowdown phase under full-scale conditions. This is based on qualitative insights gained from 

video recordings of full-scale tests. Bubbles smaller than the CFD control volumes are not 

tracked with the VOF method. This modeling limitation poses some additional uncertainties for 

these full-scale predictions. In addition to the VOF method outlined above, a Eulerian-based 

two-phase CFD approach is applied to add further insights to the analysis. This approach allows 

interpenetrating fluids and has the capability to track the void fractions associated with any small 

bubbles that break away from the main bubble. The Eulerian approach, however, is far less 

robust than the VOF methods, and no data are readily available to benchmark this approach. 

The Eulerian results serve as only an indication of the potential bubble behavior and significant 

work is needed in this area to build further confidence in the Eulerian approach under these 

conditions.

5.2.3.3 Geometry/Mesh

A simplified geometry is developed based upon information obtained from several sources. 

Volumes and flow areas from the MELCOR model are used as a basis. The model is based on 

a torus geometry with a major diameter of 32.61 m (107 ft) and a cross-sectional diameter of 

8.53 m (28 ft). The total number of downcomer pipes is 80 and each downcomer pipe has an 

inner diameter of 0.59 m (23.2 in.). The downcomers are assumed to be arranged in 40 pairs 

and spread evenly around the torus.

For the CFD model, a section of the torus equal to 1

/40th of the total is assumed to be cylindrical 

(straight) and this section contains a pair of downcomer pipes. The small cylindrical section is 

2.56 m (8.4 ft) in length and its volume is equal to 1

/40th of the full torus volume. To further 

reduce the size of the computational domain, a symmetry plane is assumed between the two 

downcomer pipes so that only 1⁄2 of this region (or about 1

/80th of the torus region) is modeled. 

Figure 5.2.3-2 illustrates the CFD model domain for the full-scale simulation. The model 

consists of the symmetrical half section of an 8.52 m (27.95 ft) inner diameter cylinder 

(representing the torus) with a length of 2.56 m (8.4 ft). The downcomer pipe has an inner 

diameter of 0.59 m (23.2 in.) and a total length, from inlet boundary to exit, of 20 m (65.62 ft). 

The 20-m length is selected to create a volume upstream of the downcomer exit that is 

consistent with 1

/80th of the volume of the drywell to torus vent piping system based upon the 

MELCOR model. The geometry is broken up into a few volumes to make creating initial 
59

conditions easier. The lower volumes, with an upper level 3.86 m (12.66 ft) from the base of the 

domain, are initially filled with water. The downcomer pipe exit is 1.0 m (3.28 ft) below the water 

surface (or 2.86 m (9.38 ft) above the base of the domain). 

The mesh design is similar to that used for the benchmark studies above. The ANSYS Fluent 

cut-cell mesh design uses cubic meshes with refinement regions created by successively 

dividing each edge of the cubes in half for each level of refinement. These refinement transition 

regions are created away from the expected bubble domain. Figure 5.2.3-3 illustrates a cross 

section of the computational mesh on a vertical plane through the downcomer pipe. Mesh sizes 

range from 0.02 m (0.787 in.) in the most refined region up to 0.32 m (12.6 in.) for the largest 

cells above the water surface. The total number of cells is 3.3 million.

Figure 5.2.3-2: CFD Torus Section Model Figure 5.2.3-3: CFD Torus Section Mesh

5.2.3.4 Boundary and Initial Conditions

The boundary and initial conditions are based on predictions from a MELCOR code simulation 

of a recirculation line break. The MELCOR predictions provide the basis for the initial 

temperatures of the liquid and gas spaces, the initial system pressure, and the time-dependent 

mass flow from the drywell to the vent pipes. The mass flow prediction from MELCOR includes

the total mass flow from the drywell into the vent piping system which includes 80 downcomer 

pipes. Because only one downcomer pipe is modeled with CFD, the inlet mass flow condition is 

set to 1

/80th of the MELCOR value (Figure 2.1-6). For the CFD analysis, the initial gas in the 

drywell, vent piping system, and wetwell gas space is assumed to be NCG. The drywell and 

vent system gas is assumed to make up the gas blown down into the suppression pool. 

Because the focus is on the initial gas bubble, only noncondensables (air in this case) are 

considered. Table 5.2.3-1 lists the boundary and initial conditions applied to the CFD model. 

Only the mass flow of air into the inlet pipe is varied with time during the simulation. The

assumption that all of the initial flow is composed entirely of air is expected to lead to a 
60

conservative prediction of the depth and size of the initial NCG bubbles in the suppression pool. 

Figure 5.2.3-4 illustrates the time-dependent mass flow into the inlet pipe boundary condition. 

MELCOR predictions (Figure 2.1-6) have provided the total mass flow of gases, including 

steam, and the CFD boundary condition assumes this entire flow is composed of air which will 

be pushed out of the pipe in front of the steam. All walls are considered to be adiabatic because

potential heat losses are considered minimal during the short duration of the initial blowdown. 

Table 5.2.3-1: Boundary and Initial Conditions for Torus Model

Boundary Condition Value

wall heat transfer adiabatic

initial water temperature 330.4 °K (135.05 °F)

initial air temperature 330.4 °K (135.05 °F)

initial pressure 111690 Pa (16.2 psi)

initial velocities 0.0

inlet air temperature 330.4 °K (135.05 °F)

inlet mass flow time-dependent table

inlet turbulence estimation 5% intensity, Hydraulic diameter 0.6 m (1.97 ft)

Figure 5.2.3-4: Mass Flows at Inlet for Full-Scale Simulation

5.2.3.5 Results

The focus of this study is the penetration of the gas bubble into the wetwell liquid. Both depth 

and width of the gas bubble are of interest. Before considering the bubble dynamics, a series of 

sensitivity studies are completed to ensure that the modeling is consistent with our expectations 

based upon the MELCOR predictions. All indications are that the CFD predictions are consistent 

with expectations of torus pressure and vent mass flow when compared to the MELCOR 
61

predictions. The fluid clears from the downcomer pipe in approximately 0.3 seconds in both 

codes. The pressure rise in the wetwell (suppression chamber) is also consistent with the 

MELCOR analysis. These qualitative comparisons build confidence in the overall setup of the 

CFD model domain and boundary conditions. Matching the pressure rise in the suppression 

pool helps ensure that the mass flow out of the downcomer pipe is consistent with the MELCOR 

model. Several sensitivity studies are completed on the mesh size, turbulence models, 

interfacial tracking methods, time step size, and downcomer pipe pressure drop to ensure that 

the final predictions are not sensitive to these parameters. Temperature comparisons are not 

included because the duration of the first bubble is relatively short and temperature variations 

are minimal. It is noted that initial temperatures are established to match the MELCOR 

modeling. 

Figure 5.2.3-5 shows contour plots of void fraction in the suppression pool at various times for 

the VOF method predictions. The contour plots are clipped between 0.1 and 0.9 to remove all of 

the pure gas or liquid regions. This view shows only the interface transition between water and 

gas and highlights these surfaces. The water surface and the expanding gas bubble surface are 

clearly highlighted. Grid lines are overlaid in the region around the downcomer exit to help 

quantify the size and depth of the gas bubble. These grid lines are centered on the pipe exit and 

spaced at a distance of 1⁄2 the downcomer pipe diameter (1

/2 D = 0.295 m (0.968 ft)). 

Figure 5.2.3-5: Contour Plots of Time-Dependent Bubble Interface (VOF Method)
62

The results in Figure 5.2.3-5 show that the water is cleared from the downcomer pipe 

approximately 0.3 s after the initiation of the large-break scenario. This is consistent with the 

MELCOR predictions. The gas bubble penetrates deep into the pool as it follows the path of the 

initial water slug that shoots out of the downcomer pipe. The peak penetration depth of nearly 

3.5 diameters is reached about 0.5 seconds later at a time of 0.8 s. From there, it takes less 

than 0.5 seconds more for the main bubble to rise back up toward the pipe outlet. The mass 

flow exiting from the pipe outlet beyond this time only reaches a depth below the exit of 

approximately 0.5 to 1 diameter. This penetration depth could decrease further as the gas 

concentration changes over to steam and condensation begins. The effect of steam and 

condensation is not considered in these cases.

It should be emphasized that these predictions for gas bubble behavior are conservative 

because the CFD analysis assumes that the downcomer vent flow is composed entirely of NCG. 

As previously discussed and shown of Figures 2.1-7 and 2.1-8, the NCG fraction of the total 

downcomer flow during the first 2 seconds following a LOCA is substantially less than the steam 

and water flow exiting the downcomer. 

The Eulerian approach is applied to consider the effect of small bubbles that are expected to 

break away from the main bubble. This two-phase CFD approach relies on a series of interfacial 

terms that can affect the final results. Validation of these terms for this project is beyond the 

scope of the current effort and the ANSYS/FLUENT default values are used to show the 

methods potential. Results are reported as predicted but it is important to note that the 

uncertainty in this method has not been established for this problem. Results should be treated 

as a qualitative estimation of the small bubble behavior. Table 5.2.3-2 lists the basis solver 

settings used for the Euler based simulations. Second order differencing methods were found to 

be unstable so a first order approach was adopted.

Table 5.2.3-2: Eulerian Multi-Phase CFD Model Basic Solver Settings

Code ANSYS/FLUENT v15.0

General pressure based transient solver, gravity

Models Eulerian Multiphase with VOF option, explicit volume fraction scheme, 

energy equation, assumed small bubble diameter = 0.005 m (0.197 

in.)

Turbulence standard k-epsilon, standard wall functions, compressibility option, 

Mixture turbulence option

Materials Water

air, ideal gas assumption 

Conditions floating operating pressure, specified operating density = 1.1777 kg/m3

(0.0735 lbm/ft3)

Boundary 

Conditions

inlet – specified mass flow- time-dependent, constant temperature

walls – adiabatic, no slip condition

P-V coupling Phase Coupled SIMPLE

Discretization density, momentum, turbulence, energy, time – first order

volume fraction - compressive

Eulerian multiphase predictions are illustrated in Figure 5.2.3-6. The results are plotted at times 

that correspond to the times used in Figure 5.2.3-5 for the VOF approach. The figures show 
63

contour plots of void fraction clipped between 0.01 and 0.99 to remove regions where there is 

less than 1 percent air or water. This view clearly shows the extent of the penetration of gas into 

the suppression pool. Grid lines are overlaid in the region around the downcomer exit to help 

quantify the size and depth of the gas bubbles. These grid lines are centered on the pipe exit 

and spaced at a distance of 1⁄2 the downcomer pipe diameter (1

/2 D = 0.295 m (0.968 ft)). The 

color bar on the left of the figure shows the scale used to map void fraction to color.

The void fraction contours in Figure 5.2.3-6 show that the Eulerian method predictions of the 

main bubble region (void > 0.99) are similar to those from the VOF method. The main bubble 

exits the blowdown pipe at around 0.3 seconds and extends to a maximum depth of around 

3.5 pipe diameters 0.5 seconds later at 0.8 s. From there, it takes less than 0.5 seconds for the 

main bubble to rise back up towards the pipe outlet. The gas flow exiting from the pipe outlet 

beyond this time penetrates downward less than 1 diameter. The Euler results do not show 

significant void away from the main bubble until after 0.8 seconds when the main bubble stops 

expanding downward and begins to rise quickly back towards the surface. At this time, a small 

void fraction representing bubbles that break away from the main bubble remains at the full 

penetration depth. This gas region rises much more slowly than the large bubble.

Figure 5.2.3-6: Contour Plots of Time-Dependent Void Fraction (Eulerian Method)

To quantify the penetration depth predictions, a graph of the maximum penetration depth as a 

function of time is created for both the VOF and Eulerian multiphase predictions. Figure 5.2.3-7 
64

shows the maximum penetration as a function of time for the VOF and Eulerian multiphase 

approaches. The VOF results are taken from results such as those in Figure 5.2.3-5 and 

represent the maximum penetration of gas into the water below the downcomer exit. For the 

Eulerian multiphase predictions, several specific void fractions are considered. The results in 

Figure 5.2.3-7 show the depth of the Eulerian predictions with a minimum void fraction of 

1 percent. Values of 2 percent and 15 percent are also considered in Figure 5.2.3-7 to provide 

additional insights. 

The results for the penetration depth show that maximum penetration is reached around 

0.8 seconds after the initiation of the transient. After that time, the main bubble rises quickly and 

recedes back to a depth of less than 1 diameter below the pipe exit at around 1.3 seconds. This 

represents the maximum penetration depth of the gases exiting the downcomer pipe at that 

time. In a prototypical scenario, steam flow is expected to follow shortly after the initial burst of 

NCGs and the bubble dynamics would change as the steam condenses.

Figure 5.2.3-7: Time-Dependent Maximum Penetration Depth

The Eulerian-based predictions provide insights into how the smaller bubbles respond. As 

expected, these do not rise as fast as the main bubble and they will be affected by global flow 

patterns in the liquid region. A region with a void fraction of 1 percent is predicted to remain at a 

depth of 3.5 diameters for a short time and then rises to the level just below the downcomer exit 

by 2.5 seconds. The 15 percent and 2 percent regions are predicted to rise more quickly and 

reach the upper location at approximately 1.5 and 2.1 seconds respectively. The Eulerian 

predictions are very similar to the VOF predictions while the main bubble is expanding down into 

the suppression pool (up until 0.8 seconds). There is a slight amount of additional penetration in 

the Eulerian results during this initial blowdown time which may be partially explained by the use 

of first order differencing in the Eulerian model (first order methods are used for stability in the 

Eulerian case). Although the VOF and Eulerian models used the same boundary and initial 
65

conditions, these are different approaches with different modeling options. After the initial 

downward penetration, the Eulerian results become more interesting. As the main bubble rises 

up, there is a wake of smaller bubbles left behind. These predictions provide some insights into 

the magnitude and void fraction of this wake region.

The lateral spread of the main bubble is another consideration. The width or spread is different 

at various depths and changes with time. To quantify this spread, fixed depths of 1, 2, and 3 L/d 

below the downcomer exit are considered. The maximum radial spread in the plane of the 

images considered (see Figure 5.2.3-6) for the three depths considered are plotted as a function 

of time in Figure 5.2.3-8. The distance is normalized in terms of downcomer diameter and 

measured from the center of the downcomer pipe. When the bubble is nonsymmetric, the larger 

radial measurement is recorded on the plot. Grid lines overlaid on the contour plots showing the 

bubble interface are used to determine the bubble width as illustrated in the example on the 

right side of Figure 5.2.3-8. The maximum radius of the main bubble in this figure is 

approximately 1.9 pipe diameters and occurs at a depth below the exit of 1 pipe diameter. It is 

noted that the bubble is wider above this level but these higher elevations are not considered.

Figure 5.2.3-8: Time-Dependent Radial Width of Main Bubble

The predictions indicate that the initial blowdown and rise of the main bubble occurs relatively 

quickly. Results indicate that within 2.5 seconds, the void fraction below the downcomer exit 

region is reduced to less than 1 percent. The results for the small bubble residence times 

inferred from Figure 5.2.3-7 have an unknown degree of uncertainty, however, because of the 

modeling approach. The physics of the bubble breakup and bubble formation in the wake of the 

main bubble has not been validated. In addition, the smaller bubbles are expected to be 

affected by larger flow patterns in the suppression pool which are likely to be plant-specific and 

not accounted for in this simplified model domain. Condensation events, which are also 

expected to occur once the NCGs clear, will also cause flow disturbances resulting in additional 
66

mixing of the smaller bubbles. These and other limitations of the approach should be considered 

before applying these results in a quantitative manner. 

As previously discussed, small gas bubbles can break away from the large gas bubble when it 

is present in the suppression pool. The small gas bubbles can remain in the suppression pool 

for a longer time than the large bubble because they are more affected by turbulence and 

circulation currents which can lengthen the time to rise and exit the pool. After the exit of the 

large gas bubble, the void fraction in the suppression pool is primarily the result of small 

bubbles. 

Figure 5.2.3-9: Average Pool Void Fraction below the Downcomer Exit 

Using the VOF and Euler Methods 

The average void fraction in the volume below the downcomer exit is of interest after the 

departure of the large gas bubble. This void fraction, which primarily results from the presence 

of small bubbles, provides an estimate of the void fraction that can enter the pump inlet strainers 

which are located below the downcomer exit in most plants. The VOF method only tracks gas 

bubbles that are larger than the computational grid size. To assess the magnitude of the 

suppression pool void fraction after the large bubble exit, the average void fraction for the 

volume below the downcomer exit is calculated for the VOF analysis. This average void fraction 

is plotted on Figure 5.2.3-9. At 2 seconds, the predicted void fraction value of about 0.02 

(2 percent) is mainly attributed to the large bubble that is formed at the exit of the downcomer 

pipe and is not associated with any distributed small bubbles. Consequently, the average void 

fraction in the suppression pool below the downcomer exit is expected to be less than 0.02 

(2 percent) after 2 seconds after the LOCA start. 
67

In contrast to the VOF method, the Eulerian method tracks void fractions associated with small 

gas bubbles that break away from the large bubble. Predictions using the Eulerian method 

provide an approximation of the void fraction in the suppression pool outside of the large gas 

bubble and in the suppression pool after the rise and exit of the large gas bubble. Figure 5.2.3-9 

includes the average void fraction in the suppression pool below the downcomer exit predicted 

by the Eulerian method. At 2 seconds, the predictions are approximately 0.03 (3 percent). Most 

of this void fraction results from the remaining large gas bubble that continues to exit from the 

downcomer. 

After 2 seconds the void fraction in the volume below the downcomer exit is expected to 

continually fall because the small NCG bubbles will rise and exit the pool and the NCG flow 

entering the pool from the downcomer drops off quickly. As shown on Figures 2.1-7 and 2.1-8, 

the MELCOR analysis predicts that the NCG flow rapidly drops and by 2 seconds consists of 

only about 8 percent of the total downcomer flow. However, as previously stated, in the CFD 

calculation the downcomer vent flow is conservatively assumed to be composed entirely of 

NCG. Additionally, using the NCG flow assumed for the CFD analysis (Figure 5.2.3-4), all the 

NCG initially present in the drywell is deposited in the wetwell by about 1.4 seconds. However, 

the CFD analyses conservatively continues gas flow after that time. Consequently, the values 

for void fraction below the downcomer exit shown in Figure 5.2.3-9 represent conservative 

estimates of the expected values.

The VOF plot on Figure 5.2.3-9 only uses the large bubble volume to calculate the average void 

fraction below the downcomer exit because the VOF method only calculates the behavior of the 

large bubble in the suppression pool. In contrast, the Euler plot uses both the large and small 

bubbles volumes to calculate the average void fraction below the downcomer exit because the 

Euler method calculates the behavior of the large gas bubble as well as any small bubbles.

Because the large bubble volume is consistent between the Euler and VOF approaches, the 

average small bubble void fraction in the fluid region outside the large bubble is approximately 

the difference between the average void fractions for the region below the downcomer exit 

calculated by the Euler and VOF analyses. Consequently, the average void fraction at a strainer 

located below the downcomer exit and outside of the large bubble will be 0.01 (1 percent) or 

less for times greater than about 2 seconds. This calculation demonstrates that the void fraction 

in the suppression pool water outside the large bubble at 2 seconds is smaller than indicated on 

Figure 5.2.3-9. A void fraction of 1 percent would result in acceptable pump operation using the 

BWR ECCS pump operation criteria described in Section 4.2.7.

5.2.3.6 Conclusions

A full-scale simulation using a simplified representation of a single downcomer pipe and 

conservative boundary conditions from a MELCOR analysis provides valuable insights into the 

bubble dynamics associated with a full-scale blowdown event. The main bubble penetrates 

approximately 3.5 pipe diameters below the downcomer exit and then rises back past the 

downcomer pipe exit within 1.5 seconds. Smaller bubbles, which are predicted to break away 

from the large main bubble, rise more slowly. These smaller bubbles are predicted to rise up 

past the downcomer exit approximately 2.5 seconds after the start of the scenario. Both a VOF 

and Eulerian approach are used to predict the bubble penetration depth. The Eulerian 

predictions provide results for a range of void fractions (i.e. for the small bubbles) but are not 

benchmarked and only serve as an indication of the small bubble behavior. The width of the 

main bubble is also considered and is limited to about 1.9 pipe diameters for depths more than 
68

1 pipe diameter below the downcomer exit. Other phenomena which could further enhance, mix, 

or distribute the NCG region, such as plant-specific features, flow fields caused by pump 

suction, or strong condensation events, should be considered on a case-by-case basis.
69

6. ASSESSMENT APPROACH FOR ECCS PUMP OPERATION IN

BWR PRESSURE SUPPRESSION CONTAINMENT SYSTEMS 

This section summarizes the recommendations for addressing the GI-193 concerns resulting 

from the review of available, applicable references and the results of the analyses presented in 

this report. The work to address a GI-193 concern for a particular plant can be divided into two 

activities. The first activity involves the determination of a time-dependent “exclusion zone” and 

a time-dependent void fraction distribution in the suppression pool following a LOCA. The 

analyses presented in Section 5.2.3 provide the information necessary to define a suppression 

pool “exclusion zone” within which a pump strainer and inlet should not be located, and a 

noncondensable void fraction distribution in the suppression pool. The second activity involves 

the determination of the ECCS pump start times and operation, and the void fraction criteria for 

acceptable ECCS pump operation. The void fraction criteria is used in conjunction with the 

information developed from the first activity to determine whether a safety concern exists in a 

particular plant.

Assuming the void fraction at the strainer location is representative of the void fraction at the 

pump inlet, the assessment of ECCS pump operation following a LOCA consists of several 

steps:

1. Determine the post-LOCA NCG “exclusion zone” in the suppression pool.

2. Determine the void fraction in the pool water outside the “exclusion zone.”

3. Determine the location of the ECCS pump strainer in relation to the “exclusion zone.”

4. Determine the acceptability of ECCS pump performance if NCG is present at the pump

strainers.

5. If the assessment determines that the ECCS pump operation would be affected, the final

step would be the determination of the course of action to correct the condition.

This approach does not include several phenomena that could affect the determination of pump 

behavior such as gas penetration into the pump strainers, gas transport through the piping to 

the pump inlet, NCGs transport in the pool toward the pump strainers caused by suction flows, 

and possible NCGs coming out of solution at the pump location in the suppression pool. These 

items will be discussed later in this section.

Using the results of the analysis results presented in Section 5.2.3 and the results of the 

literature search information presented in Section 3, the following sections discuss the method 

to define the “exclusion zone” and void fractions in a Mark I suppression pool following a DBA 

LOCA. The following sections will also provide information on the ECCS pump timing and 

acceptable operation as related to the time-dependent suppression pool void fraction conditions. 

This information will be used in a procedural method to determine the acceptability of ECCS 

pump operation following a DBA LOCA.
70

6.1 Post-LOCA Suppression Pool Behavior

As indicated in Table 2.1-1 and Figure 2.1-7, the vent flow and suppression pool conditions 

following a LOCA can be divided into four phases: (1) the downcomer vent solid water clearing 

phase, (2) a large NCG bubble or “jet” phase, (3) a small NCG bubble phase, and (4) an all 

steam flow phase that results in suppression pool steam condensation oscillation and chugging. 

The following sections describe the suppression pool phenomena occurring in each phase and 

identify the phase that presents the major concern for NCG ingestion into the ECCS pump 

suction. Figure 6.1-1 illustrates the timing of the suppression pool conditions and pump 

operation using information discussed in Sections 2, 4 and 5. The high-containment pressure 

signal occurs about 1 second following the start of a DBA LOCA. An additional minimum 

3-second pump startup delay is applied with loss of offsite power to account for emergency 

diesel generator startup and ECCS pump initiation sequencing.

Figure 6.1-1: Downcomer Flow/Suppression Pool Conditions

and ECCS Pump Operation Comparison

6.1.1 Downcomer Vent Water Clearing

A water column exists in the downcomer vent before the start of a LOCA. As the drywell 

pressure increases following a LOCA the water column is pushed into the suppression pool 

because of the increased pressure differential between the drywell and wetwell. As discussed in 

Sections 2 and 5.2, the water clearing phase lasts about 0.3 seconds for a DBA LOCA, but can 

take longer for smaller break sizes.

6.1.2 Noncondensable Gas Flow and Bubble Behavior

The NCG flow behavior can be divided into two phases, a large gas bubble phase followed by a 

longer small gas bubble phase. After the completion of vent water clearing, the noncondensable 

gas-steam mixture in the drywell begins to flow out of the vent and into the suppression pool. As 

shown on Figure 2.1-7, the initial flow in this phase contains a significant amount of NCG. This 

initial large gas “jet” can penetrate a significant depth into the pool as it follows the initial water 

“jet” downward. The amount of NCG in the vent flow decreases quickly and reaches very small 

levels after 2–3 seconds for a DBA LOCA. The size of the initial large NCG “jet” can define an

Pump Start Signal ≥1 to 20s

Vent Water 

Clearing≈0.3s Small NCG bubble Void Fraction ~0.3 to ~20s Downcomer steam Flow after ~20s

~0.3 s 20 s 40 s

Start Pump Core Injection at ~40s

0 s

Large NCG jet

~0.3 to <~2-3s

Largest NCG 

Concentration

~2-3 s

Pump Startup & Pool Recirculation ≥4 to ~40s

With offsite power

With loss of offsite power Start Pump Core Injection at ~40s

Pump Startup & Pool Recirculation ≥1 to ~40s

Pump Start Signal ≥4 to 20s
71

“exclusion zone” volume at the vent exit which possesses a large fraction of NCG. The ECCS 

strainer and pump suction should not extend into the “exclusion zone” during ECCS pump 

operation. After 2–3 seconds the size of the NCG volume decreases as the initial large gas 

bubble rises and exits the top of the suppression pool. After the large NCG bubble exits, small 

gas bubbles can still flow into the suppression pool water; however the volume and penetration 

of this gas flow is significantly less than the initial large gas “jet.” The average suppression pool 

void fraction is expected to significantly decrease to low levels by about 10 seconds, but, much 

lower levels of NCG void fractions levels may last in the suppression pool to about 

30–40 seconds.

6.1.3 Condensation Oscillation and Chugging

After the NCG flow reaches very low levels, the bulk of the flow exiting the vent is a steam-water 

mixture. Initially this flow rate can be significant but will decrease as the vent flow decreases 

following the LOCA. High steam flow results in steady condensation. Decreasing steam flow can 

produce oscillating bubbles at the vent exit resulting from steam condensation at the bubble 

surface. This phenomenon is called condensation oscillation. At even lower steam flow 

conditions, the exiting steam flow can create a bubble that fully condenses rapidly or collapses 

at the vent exit causing backflow of water into the vent pipe. This “chugging” phenomenon can 

occur in a cyclical fashion consisting of steam bubble formation and full collapse.

The steam condensation mode, which generally exists at times later than 10 seconds after a 

DBA LOCA, is not expected to impair ECCS pump operation because the fraction of NCG

exiting the vent and entering the suppression pool is small and because any steam in the 

suppression pool is expected to condense before entering the ECCS pump inlet. 

6.2 Other Considerations Regarding Suppression Pool Conditions

Additional concerns were considered in the development of the logic for assessing acceptable 

ECCS pump performance following a LOCA. The effect of initial dissolved NCG in the 

suppression pool was determined to have a secondary or negligible effect on pool void fraction. 

Additionally, it was considered conservative to use the pool void fraction outside the pump inlet 

strainer in this assessment because consideration of NCG behavior in the intake line to the 

ECCS pump would generally result in a lower void fraction at the pump inlet.

6.2.1 Consideration of Dissolved Noncondensable Gas in the Suppression Pool

The suppression pool water contains dissolved NCG. This section assesses whether significant 

levels of NCG in the suppression pool can come out of solution following a LOCA and affect the 

pool void fraction. 

The maximum saturated equilibrium amount of dissolved gas that a water pool can contain is 

dependent on the pool temperature and pressure. Following downcomer vent discharge after a 

large double-ended guillotine break LOCA, local heat up of the suppression pool could result in 

some of the dissolved gas coming out of solution as bubbles if the equilibrium NCG 

concentration at the elevated temperature and pressure is less than the equilibrium gas 

concentration at the initial steady-state conditions before the LOCA. Because of the following 

calculations, the inclusion of the initial dissolved gas in the suppression pool is not considered 

an important criterion for determining void fraction distribution in the suppression pool. 
72

Additionally, it was concluded that the dissolved air present in the suppression pool water is not 

expected to significantly affect the void fraction distribution calculated by a CFD analysis in the 

suppression pool for the initial 1- to 2-second period following a large-break LOCA.

Information from a typical Mark I plant can be used as the basis for determining increase in pool 

void fraction resulting from a LOCA. Typical Mark I containment design parameters are as 

follows:

Drywell net free volume 4,785.5 m3

(169,000 ft3) 

Max. suppression pool water volume at high water level 3,454.7 m3

(122,000 ft3) 

Min. suppression chamber free volume at high water level 3,780.3 m3

(133,500 ft3) 

Drywell and suppression chamber pressure 1.034×104 Pag

(1.5 psig) 

Drywell temperature 330.37 °K

(135 oF) 

Suppression chamber atmosphere and pool temperature 308.15 °K

(95 oF) 

Peak short term suppression chamber pressure following a 

LOCA

1.896×105 Pag

(27.5 psig) 

Peak short term suppression pool temperature following a 

LOCA

329.26 °K

(133 oF) 

Using information for equilibrium air concentration in water from References 52 and 53, the 

equilibrium concentration of air in water can be estimated. The equilibrium air concentration in 

water 

at 308.15 °K (95 oF) and 1.034×104 Pag (1.5 psig) 

is about 0.0192 kgair/m3

mixture (1.2×10-3 lbmair/ft3mixture) 

at 288.71 °K (60 oF) and 1.034×104 Pag (1.5 psig)

is about 0.0256 kgair/m3

mixture (1.6×10-3 lbmair/ft3mixture) 

at 329.26 °K (133 oF) and 1.896×105 Pag (27.5 psig)

is about 0.0216 kgair/m3

mixture (1.35×10-3 lbmair/ft3mixture) 

Lower Limit Calculation

With the suppression pool initially at its maximum steady-state operating temperature of 

308.15 degrees K (95 degrees F), it is concluded that the dissolved air in the suppression pool 

will remain dissolved because the equilibrium air concentration at the peak short-term postLOCA pressure and temperature condition (329.26 degrees K (133 degrees F) and 

1.896×105 Pag (27.5 psig)) is larger than the steady-state conditions before the LOCA.
73

Upper Limit Calculation

If the suppression pool is initially assumed to be at 288.71 °K (60 oF) and the air dissolved in the 

suppression pool water comes out of solution at the peak short term conditions following a 

LOCA, the resulting air mass released by the suppression pool water can be estimated as:

(1.6x10-3 – 1.35x10-3) lbmair/ft3

mixture × 122,000 ft3 = 30.5 lbmair (13.8 kgair) 

The density of air at 313.15 °K (104 oF) is 1.089 kg/m3

 (0.068 lbm/ft3) and at 333.15 °K (140 oF) 

is 1.025 kg/m3

 (0.064 lbm/ft3). 

Consequently, the average void fraction in the suppression pool at 333.15 degrees K 

(140 degrees F) resulting from the release of dissolved air can be estimated as:

30.5 lbm / 0.064 lbm/ft3 / 122,000 ft3 = 0.0039 = 0.39%

This value is the upper limit on void fraction increase in the suppression pool caused by the 

instantaneous release of dissolved air in the pool water for the short-term period following a 

LOCA. This upper limit void fraction increase is small compared to the expected large void 

fraction distribution in the suppression pool during the first 2 seconds following a large-break

LOCA. A pool void fraction increase of 0.39 percent would not affect the conclusions regarding 

ECCS pump operation resulting from the larger pool void fraction increase caused by 

downcomer air flow following a LBLOCA. Additionally, even for this upper limit case scenario, 

the void fraction in the suppression pool is not expected to rise immediately to the maximum 

value because dissolved air is not released instantaneously from a water mixture but takes 

some time to be released (Ref. 50).

Therefore, it is concluded that the dissolved air present in the suppression pool water is not 

expected to significantly affect the void fraction distribution calculated by a CFD analysis in the 

suppression pool for the initial 1- to 2-second period following a large-break LOCA.

6.2.2 Consideration of Noncondensable Gas Transport to and Accumulation at the 

ECCS Pump Inlet before Pump Startup

This concern relates to the potential for the NCG to flow through the ECCS pump inlet strainer 

before pump startup, remain in the dead-ended pipe region downstream of the strainer, and 

enter the pump as it starts. This concern, which is dependent on the design of the pump inlet 

strainer and the geometry of the pump intake line, was judged to be insignificant for the reasons

discussed in the following paragraphs. 

As previously indicated NCG from the drywell can enter the suppression pool only after the 

downcomer vents are cleared of water. As shown on Figure 6.1-1, vent clearing is completed at 

about 0.3 seconds following the start of the DBA LOCA. Calculations indicate that NCG starts 

flowing into the suppression pool at about 0.3 seconds and forms a large gas bubble that lasts 

to about 2–3 seconds when the large gas bubble exits the pool. As indicated by the plant survey 

performed by the BWR Owners Group, the earliest time for the start of the ECCS pumps is 

about 1 second assuming the availability of offsite power and at about 4 seconds assuming the 

loss of offsite power. Consequently, the earliest time frame during which NCG can flow from a 

large gas bubble into a dead-ended intake pipe is the time period between 0.3 seconds and 1 or

4 seconds when the pump starts following a DBA LOCA. NCG flow from small bubbles in the 

suppression pool will exist from about 0.3 seconds to about 20 seconds following a DBA LOCA. 
74

The possibility for gas to enter the pipe intake before pump startup is dependent on the local 

pool void fraction conditions at the pipe intake strainer. Consequently, the time and location 

dependent void fraction at a specific strainer location in the suppression pool can be determined 

from Figures 5.2.3-5 and 5.2.3-6. The postulated NCG flow from the suppression pool through 

the intake strainer could be greater if the strainer is located in a region with a larger pool void 

fraction.

Before pump startup, NCG flow from the suppression pool to the intake pipe is unlikely because 

the gas must flow from the pool through an intake strainer, which has a large flow resistance to 

enter the intake pipe. Additionally, the incompressible water in the pipe must flow backwards 

from the pipe through the high flow resistance strainer into the suppression pool to make room 

for the incoming gas flow. Consequently countercurrent gas-liquid flow must be present in the 

strainer and intake pipe for gas to accumulate in the intake pipe. The presence of 

countercurrent flow in the intake strainer and intake pipe would further increase the flow 

resistance for additional gas to enter the intake pipe. 

Additional resistance to gas flow into the intake pipe can exist because of the surface tension 

effects in the small holes that may be present in the intake strainer. To permit gas flow through 

the small holes, a sufficient pressure differential must exist across the small holes in the intake 

strainer to overcome the capillary surface tension force represented by the following equation.

Δp = 2 σ / r

where p is the pressure force required to overcome the capillary surface tension force

σ is the surface tension, and

r is the radius of curvature of the intake strainer holes, typically the equivalent hole 

radius. 

It is highly unlikely that the flow conditions described above would exist to allow gas to flow 

through the inlet trainer and into the inlet pipe, and for countercurrent gas-liquid flow to occur 

across an inlet strainer connected to a dead ended pipe.

In the unlikely case that gas does flow through the inlet strainer, the small holes in the inlet 

strainer will break up any large pool gas bubble into smaller gas bubbles because the bubble 

size in the intake pipe will be limited to the size of the strainer holes. This behavior will result in 

a smaller pipe void fraction after mixing with the water volume in the intake pipe. Additionally, 

the smaller bubbles in the intake pipe would need more time than available to coalesce into a 

larger bubble before the ECCS pumps start.

Additionally, each ECCS pump is connected to a header which is fed by several strainer pipes. 

Consequently, any two-phase, low void fraction flow in an individual intake pipe will mix with 

water coming from other intake lines in the header that feeds the pump reducing the void 

fraction before entering the pump.
75

6.2.3 Consideration of Noncondensable Gas Transport in the Inlet Piping during Pump 

Operation

The developed assessment procedure conservatively uses the void fraction at the strainer inlet 

for its assessment and does not account for the void fraction decrease resulting from the 

pressure drop that can exist between the strainer inlet and pump inlet during ECCS pump 

operation. The void fraction decrease from the strainer inlet to the pump inlet is dependent upon 

the piping geometry (e.g. length and diameter), the pump operation characteristics (Refs. 50 

and 51), and the flow resistances and fluid characteristics of the inlet line as described in 

Section 6.2.2. A time-dependent pool void fraction at the strainer inlet can be estimated by 

determining whether the strainer inlet is located inside or outside the “exclusion zone”; however, 

the void fraction change from the strainer to the pump inlet can only be calculated using 

plant-specific design information. Consequently, the determination of the void fraction at the 

pump inlet is beyond the scope of the GI-193 activities, but may be included in a calculation 

using plant-specific design information. Alternatively, it is conservative to assume that the void 

fraction at the strainer inlet is representative of the void fraction at the pump inlet.

6.3 Determination of ECCS Pump Strainer Location in a Suppression Pool

The location of the ECCS pump strainer relative to the downcomer vents in the suppression 

pool of a BWR PS containment is plant dependent. Consequently the assessment of the 

acceptability of an ECCS pump strainer location must be determined on a plant-specific basis.

6.4 ECCS Pump Operation

The assessment of ECCS pump behavior must consider the NCG distribution in the 

suppression pool. The CFD analysis results can be used to define the “exclusion zone” and the 

gas void fraction in the suppression pool outside the “exclusion zone.” The pump strainer and 

intake should be located outside the gas “exclusion zone” during ECCS pump operation. The 

gas void fraction concentration outside the “exclusion zone” is important to define the gas void 

fraction entering the pump strainer and intake. 

ECCS pump operation should be considered to be acceptable if the gas void fraction at the 

pump intake is determined to be less than the void fraction at which pump degradation would be 

expected. If the local pool void fraction falls above the void fraction at which pump degradation 

would be expected, an assessment must be made as to whether the pump operation would 

recover after the NCG exits the pool. This assessment should consider the pump start time, 

pump speed, and flow rate, operating duration and the duration of the elevated void fraction at 

the pump inlet.

6.4.1 ECCS Pump Timing Considerations

Section 4.1 discusses the post-LOCA startup timing considerations for the ECCS pumps. Each 

plant should perform a plant-specific assessment to determine the ECCS pump startup times 

with and without the availability of offsite power. The plant-specific assessment should also 

specify the pump speed during the period when flow is recirculated in the suppression pool and 

the time when pump flow is reconfigured to provide core injection.
76

As mentioned in Section 4.1, a typical minimum start-up time for the LPCS and LPCI pumps is 

about 1 second with offsite power available and about 4 seconds without the availability of 

offsite power. During the initial period the LPCS and LPCI pumps are operated to reach full 

speed. Additionally, between startup and 40 to 50 seconds the ECCS pump flow is recirculated 

to the suppression pool and is not injected into the core. Regardless of the availability of offsite 

power, the LPCS and LPCI systems are expected to start pump flow into the core at about 40 to

50 seconds after the appropriate valve reconfiguration.

As discussed in Section 5.2.3 and illustrated in Figure 6.1-1, ECCS pump operation could 

primarily be affected if a pump operates within the first 2 seconds following a DBA LOCA. 

However, lower levels of NCG can last to about 30–40 seconds because of the time necessary 

for the bulk of the gas to rise and exit the pool surface. Figure 6.1-1 shows the relationship 

between the downcomer flows described in Section 5.2.3 and the ECCS pump operations as 

described in Section 4. 

6.4.2 ECCS Pump Acceptable Operation Criteria

An ECCS pump can sustain damage caused by over speed and vibration if gas is ingested. The 

acceptable ECCS pump operating range for NCG inlet void fraction was discussed in Section 4. 

Section 4 also discussed the ability of the ECCS pump to recover especially if only a small 

amount of gas is ingested. As discussed in Section 4, the BWR ECCS pump operational criteria 

were developed and provided in References 44 and 45. The BWR ECCS operational criteria are

summarized in Table 4.1-1.

The BWR ECCS operational criteria presented in Section 4 are summarized below.

1. Low transient flow rates (less than 70 percent of the best efficiency point) with an inlet

void fraction greater than 5 percent for more than 5 seconds will result in pump damage

unless pump-specific testing can show otherwise.

2. Higher transient flow rates (between 70 percent and 120 percent of the best efficiency

point) with an inlet void fraction greater than 10 percent for more than 5 seconds will

result in pump damage unless pump-specific testing can show otherwise.

3. Low flow operation (less than 40 percent of the best efficiency point) at greater than

1 percent by volume continuous suction NCG void fraction results in pump damage

unless pump-specific testing can show otherwise.

4. Full flow operation (40 percent to 120 percent of the best efficiency point) at greater than

2 percent by volume continuous suction NCG void fraction results in pump damage

unless pump-specific testing can show otherwise.

ECCS pumps that operate within the void fraction conditions specified in Table 4.1-1 and within 

the gas void fractions indicated above in items 1 to 4 can be considered capable of recovery 

without damage, and will operate at normal pump head and flow conditions if the inlet void 

fraction is reduced to a value below 1 percent. 

An ECCS pump assessment would be permitted to use other operation criteria if acceptable 

pump-specific tests are performed that provide better operational guidance.
77

6.5 Procedure for Performing a GI-193 Plant Assessment

The assessment of the ECCS pump behavior must consider the NCG distribution in the 

suppression pool. If a pump starts before the large gas bubble exits the suppression pool, the 

pump strainer and intake should be located outside the transient gas “exclusion zone” defined 

by the large gas bubble. However, the gas void fraction concentration outside of the “exclusion 

zone” is also important to define the gas void fraction entering a pump strainer and intake 

located outside the “exclusion zone.” The results from CFD methods described in Section 5.2.3

could be used to define the gas void fraction in the suppression pool outside the “exclusion 

zone” defined by the large NCG bubble. If the gas void fraction at the pump intake is determined 

to be less than the void fraction at which pump degradation would be expected, the ECCS pump 

operation should be considered to be acceptable. (This assumes that pump suction does not 

move the gas bubble if the pumps starts outside of but in close proximity to the gas region.)

If the local pool void fraction falls above the void fraction at which pump degradation would be 

expected, an assessment must be made as to whether the pump operation would recover 

without damage after the NCG exits the pool. This assessment should consider the pump start 

time, pump speed and flow rate, operating duration and the duration of the elevated void 

fraction at the pump inlet.

Section 6.4.1 discusses the post-LOCA startup timing considerations for the ECCS pumps. 

Each plant should perform a plant-specific assessment to determine the ECCS pump startup 

times with and without the availability of offsite power. The plant-specific assessment should 

also specify the pump speed during the period when flow is recirculated in the suppression pool 

and the time when pump flow is reconfigured to provide core injection. 

After pump startup, the criteria for acceptable pump operation will be determined using the flow 

logic presented in Figure 6.5-1 which provides a logic stream for the procedure to assess the 

post-LOCA operation of the ECCS pumps. The procedure proposes that the ECCS pump 

operation be assessed using the maximum and transient NCG void fraction at the strainer 

location using the graphs contained in Figures 5.2.3-5, 5.2.3-6, 5.2.3-7 and 5.2.3-8. Additionally, 

Section 5.2.3.5 and Figure 5.2.3-9 suggest that the void fraction at a strainer located below the 

downcomer exit and outside of the large bubble will be 0.01 (1 percent) or less for times greater 

than about 2 seconds. This assessment approach is proposed because of the short duration of 

the large NCG large bubble which is less than 3 seconds. For this assessment it is assumed 

that the “exclusion zone” for the large NCG bubble is defined by the boundary where the pool 

void fraction exceeds 10 percent. This assumption is consistent with the ECCS pump 

acceptable operation criteria listed in Section 6.4.1. Other void fraction criteria are also 

consistent with the information listed in Section 6.4.2. 

The assessment procedure presented in Figure 6.5-1 was developed using information for a 

Mark I containment. Because of similar vertical downcomer geometry, the results of this study 

can be conservatively applied to a Mark II containment design, which possesses a larger 

suppression chamber. A Mark III design includes a horizontal vent pipe orientation and the 

ECCS pump inlets are located significantly further from the vent exits. Consequently, the ECCS 

pumps in a Mark III design are expected to be less affected by NCG because lower void 

fractions are expected at their inlets.
78

Figure 6.5-1 includes an option to perform ECCS pump-specific testing if the application of the 

flow logic determines that the timing and void fraction criteria results in pump failure including 

the inability to recover. This testing option would provide the actual operability assessment of a 

particular ECCS pump.

For this assessment, ECCS pump performance during operation with void fraction conditions at 

the pump inlet was considered from a general, not plant-specific perspective. A review of 

Figure 6.5-1 reveals that an ECCS pump start time greater than about 2–3 seconds may be 

sufficient to conclude that pump operation is unaffected. This assessment supports the technical 

conclusion that delayed ECCS pump flow initiation beyond 3 seconds, plus a possible safety 

factor, following a DBA LOCA initiation will adequately protect the ECCS pumps from damage. 

Similarly, consideration of the void fraction at the strainer location may be sufficient to conclude 

that pump operation is unaffected. This assessment approach assumes that the void fraction at 

the strainer location is transferred unchanged through the strainer and piping to the pump inlet. 

This is a conservative assumption with regards to the determination of acceptable pump 

operation. The assessment includes the option to refine the pump operation assessment by 

using calculations to determine the void fraction at the pump inlet. Other studies and tests have 

demonstrated that the void fraction at the strainer location could be substantially different after 

traveling a distance through a pipe to the pump inlet. Consequently, calculational methods 

outlined in Reference 44 can be used to estimate the gas transport in the pipe from the strainer 

to the pump inlet to determine the actual void fraction at the pump inlet. Additionally, a 

defensible method would need to be developed to calculate the transport of NCG from the large

bubble or small bubbles in the suppression pool, through the strainer and into the pipe 

downstream of the strainer. Of course, the method used to perform this calculation (e.g. a 

thermal-hydraulic computer code) must be verified against applicable test data before it could 

be used in this application.

Another approach to determine the void fraction at the pump inlet would be to include the 

strainer inlet and piping leading to the ECCS pump in the full CFD model used to predict the 

NCG bubble behavior. This model would include details of the pump strainer and piping along 

with the time-dependent pump suction flow at the end of the piping. This approach would not 

only estimate the void fraction in the piping approaching the pump inlet, but would also predict 

the interactions between the size and location of the NCG bubble and the calculated 

suppression pool flow field and void fraction distribution that includes the effect of the pump 

suction. Of course, the CFD method used to perform this calculation must be verified against 

applicable test data before it could be used in this application.

6.6 Possible Corrective Actions

If an ECCS pump assessment concludes that a pump would fail and not recover, the plant 

operator must determine what corrective action would be necessary. Four possible corrective 

actions could be considered.

1. The ECCS start time could be delayed to ensure that the pump would start after

elevated gas void fraction conditions are reduced in the suppression pool at the

elevation of the pump strainer and intake.

2. The location of the ECCS pump strainer and intake could be moved to eliminate or

minimize the gas void fraction concern at the pump strainer and intake.
79

3. Closing off a specific pump strainer is also an option if other strainer intakes provide

sufficient flow into a common header feeding a pump.

4. As previously mentioned, the licensee may also choose to perform a more detailed

plant-specific analysis which models the strainer and downcomer locations, and explicitly

calculates the voiding at the ECCS pump inlet.
80

Is strainer location

α > 10% for > 5 s?

Yes

No

Is pump start time 

≥ 3 s?

Yes

Acceptable pump 

operation

No Yes

No

Pump failure 

Move strainer, block strainer or adjust start time

Is strainer location 

α ≤ 1%?

Yes

Will pump fail?

Yes

No

Can pump recover?

Yes

No

Yes

No

Assume pump failure. 

Move strainer, block 

strainer or adjust start time

Is strainer location

α ≤ 5% for ≤ 5 s 

Acceptable pump 

operation

Determine time dependent void fraction (α) at strainer location

Figure 6.5-1: Approach for Assessing Effects of Noncondensable Gas on ECCS Pump Operation

Determine pump start time

Acceptable pump 

operation

Determine time (tα) when α ≤ 5%

Is pump start time 

≥ tα

No

Acceptable pump 

operation

Yes

No Perform pump specific testing? AND/OR

Perform detailed plant specific analysis to determine 

pump inlet void fraction?
81

6.7 Sample Calculation for Performing a GI-193 Plant Assessment

To illustrate the application of the plant assessment procedure, assessments are made for 

strainer locations at four points. The depth position is measured as the ratio of length over pipe 

diameter (L/d) from the bottom of the downcomer. The radial position is measured as L/d 

lengths measured from the downcomer centerline. The suppression pool void fractions are 

measured off Figure 5.2.3-6. Table 6.7-1 provides the void fraction measurements for these four 

locations. The maximum void fraction is highlighted in bold characters. Figures 6.7-1 and 6.7-2 

plot the measured void fractions.

Table 6.7-1: Suppression Pool Void Fraction Sample Cases for GI-193 Assessment

Time (s)

Case 1 Case 2 Case 3 Case 4

x (L/d) r (L/d) x (L/d) r (L/d) x (L/d) r (L/d) x (L/d) r (L/d)

3 0 3 ±1 2 ±1 2 ±1.5

0.0 0.0 0.0 0.0 0.0

0.3 0.0 0.0 0.0 0.0

0.4 0.0 0.0 0.0 0.0

0.6 0.15 0.0 1.0 0.0

0.8 1.0 0.25 0.75 0.02

1.0 0.25 ≤0.01 0.35 0.02

1.2 0.06 ≤0.01 0.06 0.02

1.4 0.01 ≤0.01 0.02 ≤0.01

1.6 0.01 ≤0.01 ≤0.01 ≤0.01

1.8 0.01 ≤0.01 ≤0.01 ≤0.01

2.0 ≤0.01 ≤0.01 ≤0.01 ≤0.01
82

Figure 6.7-1: Plot of Suppression Pool Void Fraction 

Sample Cases for GI-193 Assessment

Figure 6.7-2: Log Plot of Suppression Pool Void Fraction 

Sample Cases for GI-193 Assessment
83

Case 1—This case reaches a maximum void fraction of 100 percent at the strainer location and 

remains above 5 percent for about 0.7 seconds. The assessment procedure indicates that this 

case would result in pump failure if the pump start time is less than about 1.3 seconds because 

the void fraction drops below 5 percent between 1.2 and 1.3 seconds. The void fraction is less 

than 5 percent after about 1.4 seconds; therefore the ECCS pump would be expected to 

operate acceptably if it starts after 1.3 seconds. Pump-specific test can be performed to 

ascertain if pump performance is acceptable if the start time is less than 1.3 seconds.

Case 2—This case reaches a maximum void fraction of about 25 percent at the strainer location 

and remains above 5 percent for about 0.2 seconds. The assessment procedure indicates that 

this case would result in pump failure if the pump start time is less than about 0.9 seconds. 

However, because no plant is expected to have an ECCS pump start time less than 

1.0 seconds, this strainer location would result in acceptable ECCS pump performance. 

Case 3—This case reaches a maximum void fraction of 100 percent at the strainer location and 

remains above 5 percent for about 0.8 seconds. The assessment procedure indicates that this 

case would result in ECCS pump failure if the pump start time is less than about 1.3 seconds. 

The void fraction falls to about 5 percent at less than 1.3 seconds; therefore the ECCS pump 

would operate acceptably if it starts after about 1.3 seconds. Pump-specific test can be 

performed to ascertain if pump performance is acceptable if the start time is less than 

1.3 seconds.

Case 4—This case reaches a maximum void fraction of about 2 percent at the strainer location 

and remains at that value for about 0.6 seconds between 0.6 and 1.4 seconds. The void fraction 

falls below 2 percent between 1.2 and 1.3 seconds and falls below 1 percent at less than 

1.3 seconds. The ECCS pump is expected to operate acceptably or recover from failure

because the void fraction at the strainer does not exceed 5 percent for more than 5 seconds 

and quickly falls below 1 percent. 

85

7. SUMMARY AND CONCLUSIONS

This report describes a method to perform a plant-specific assessment to determine whether an 

ECCS pump will operate acceptably following a DBA LOCA in a Mark I pressure suppression 

pool using the analyses results and criteria presented in this report. The development of an 

assessment method is summarized in Section 6. The assessment approach uses the ECCS 

pump strainer location, the ECCS pump operation criteria, the post-LOCA ECCS pump startup 

and full operation timing information, and the pool void fraction information at the pump strainer 

to make its determination. It should be noted that the plant operators will be responsible for 

applying the methodology and criteria to their specific plant geometry to complete an 

assessment for a particular plant.

The work to address GI-193 for a particular plant can be divided into several activities. The first 

activity presented in this report defined the transient penetration and duration of the large 

noncondensable gas bubble “exclusion zone” in a Mark I suppression pool following a DBA 

LOCA. The analyses also provide the time and location dependent void fraction in the 

suppression pool water outside of the large gas bubble. The second activity defined the 

acceptable void fraction criteria for ECCS pump operation obtained from literature review. The 

void fraction criteria would be used in conjunction with the information developed from the first 

activity to determine whether a safety concern exists in a particular plant.

7.1 Determination of Suppression Pool Exclusion Zone and Void Fraction 

Distribution Assessment Method

As mentioned in Section 5.1, all attempts to develop a scaling approach using small-scale test 

data to determine a NCG “exclusion zone” and void fraction distribution for post-LOCA 

downcomer venting in a full-scale suppression pool proved unattainable. However, a CFD 

analysis approach was used successfully to address the GI-193 concerns. CFD analyses of the 

water pool for two test facilities were performed to obtain the large gas bubble size and the 

surrounding void fraction distribution using data at different downcomer pipe diameters. Data 

from the following facilities provided the necessary test data:

1. the Lappeenranta University PPOOLEX tests

2. the Purdue University (PUMA) tests

A CFD model and analysis was applied to a full-scale plant geometry using the modeling 

techniques developed to analyze the test facilities. The full plant CFD analysis predictions were

qualitatively compared to the videos and data from the GE full-scale Mark II tests.

7.2 Acceptable ECCS Pump Operation Criteria

The acceptable BWR ECCS pump operational criteria presented in Section 4 are summarized 

below.

1. Low transient flow rates (less than 70 percent of the best efficiency point) with an inlet 

void fraction greater than 5 percent for more than 5 seconds will result in pump damage 

unless pump-specific testing can show otherwise. 
86

2. Higher transient flow rates (between 70 percent and 120 percent of the best efficiency 

point) with an inlet void fraction greater than 10 percent for more than 5 seconds will 

result in pump damage unless pump-specific testing can show otherwise.

3. Low flow operation (less than 40 percent of the best efficiency point) at greater than 

1 percent by volume continuous suction NCG void fraction results in pump damage 

unless pump-specific testing can show otherwise.

4. Full flow operation (40 percent to 120 percent of the best efficiency point) at greater than 

2 percent by volume continuous suction NCG void fraction results in pump damage 

unless pump-specific testing can show otherwise.

ECCS pumps which operate within the void fraction conditions specified in Table 4.1-1 and 

within the gas void fractions indicated above in items 1 to 4 can be considered capable of 

recovery without damage, and will operate at normal pump head and flow conditions if the inlet 

void fraction is reduced to a value below 1 percent.

An ECCS pump assessment would be permitted to use other operation criteria if acceptable 

pump-specific tests are performed that provide better operational guidance.

The method outlined in this report is limited by the assumptions used to develop the criteria. For 

instance the assumptions regarding the ECCS pump start times is based on the survey 

information present in Reference 4. The pump operation acceptance criteria were obtained from 

the literature search and testing information described in Reference 44. The CFD approach 

used to predict the location and duration of the NCG phase is also limited to the simplified cases 

considered. Other phenomena such as suppression pool structures or flow patterns which could 

further enhance, mix, or distribute the NCG region should be considered on a case-by-case 

basis.

The analyses used to develop the criteria were developed for a DBA LBLOCA in the 

suppression pool of a Mark I containment using the information provided in Reference 2. The 

DBA LOCA results in the most severe NCG conditions in a suppression pool because it 

provides the largest and fastest break flowrates that result in the largest gas flow into the 

suppression pool in the shortest time. A Mark I containment results in the most limiting NCG

suppression pool conditions because of its smaller size and smallest downcomer vent 

submergence when compared to the larger Mark II and Mark III suppression pool designs. 

Additionally, the ECCS pump strainers are generally located closest to the downcomers in a 

Mark I design. Because of similar vertical downcomer geometry, the results of this study can be 

conservatively applied to a Mark II containment design, which possesses a larger suppression 

chamber. A Mark III design includes a horizontal vent pipe orientation and the ECCS pump 

inlets are located significantly further from the vent exits. Consequently, the ECCS pumps in a 

Mark III design are expected to be less affected by NCG because lower void fractions are 

expected at their inlets.
87

8. REFERENCES

1. U. S. Nuclear Regulatory Commission, Memo from Farouk Eltawila to Carl Paperiello, 

“Task Action Plan for Resolving Generic Safety Issue 193, BWR ECCS Suction 

Concerns,” May 24, 2004 (Agencywide Documents Access and Management System 

(ADAMS) Accession No. ML041450208).

2. Notafrancesco, A., H. Esmalli, B. Lee and J.L. Tills, “Application of the MELCOR Code to 

Design Basis BWR Containment Analysis,” U.S. Nuclear Regulatory Commission, Office 

of Nuclear Regulatory Research, RES/FSTB 2011-01, May 2011 (ADAMS Accession 

No. ML111570060).

3. Lahey, R.T. and Moody, F.J., The Thermal Hydraulics of a Boiling Water Reactor –

Second Edition, American Nuclear Society, 1993.

4. Fleischer, L., “Assessment of NEC Generic Issue #193 Prepared for Generic Issue #193 

Committee of the BWR Owners’ Group,” General Electric Hitachi Nuclear Energy, 

NEDC-33526-P Rev. 1, October 2009.

5. “Mark-I Containment Program, Full-Scale Test Program, Final Report,” General Electric, 

NEDE-24539-P, April 1979.

6. “Mark-III Confirmatory Test Program, Full Scale Condensation and Stratification 

Phenomena Test Series 5707,” General Electric, NEDE-21853-P, August 1978.

7. “1/4 Scale Test Report, Loads on Submerged Structures due to LOCA Air Bubbles and 

Water Jets,” General Electric, NEDE-23817-P, September 1978.

8. “Mark-I Containment Program, 1⁄4 Scale Pressure Suppression Pool Swell Test Program: 

LDR Load Tests,” General Electric, NEDE-23545-P, December 1978.

9. “Condensation Oscillation Test Program Final Test Report,” General Electric, NEDE24811-P, May 1980.

10. McCauley, E.W, et al., “Extended Analysis of Data from the 1/5-Scale Mark I Boiling

Water Reactor Pressure Suppression Experiment,” U.S. Nuclear Regulatory 

Commission, NUREG/CR-0761, September 1979 (ADAMS Accession 

No. ML072710074).

11. Pitts, J.H. and E.W. McCauley, “A 1/5 Scale Experiment of a Mark I Boiling Water 

Reactor Pressure-Suppression System under Hypothetical LOCA Loads,” Lawrence 

Livermore Laboratory, UCRL-79143, July 1977.

12. Lai, W. and E.W. McCauley, “Air Scaling and Modeling Studies for the 1/5-Scale Mark I 

Boiling Water Reactor Pressure-Suppression Experiment,” Lawrence Livermore 

Laboratory, UCRL-52383, January 1978.

13. Norris, D.M. et al., “Computer Calculations of Air and Steam Blowdown Suppression,” 

Nuclear Engineering and Design, Vol. 59, pp. 301-313, 1980.
88

14. McCauley, E.W. et al, “The Analysis of BWR Pressure Suppression Pool Dynamics,” 

Lawrence Livermore Laboratory, UCRL-78694, August 1976.

15. Glenn, L.A. and W.H. McMaster, “BWR Pool Dynamics in a Loss of Coolant Accident of 

the Air Venting Phase,” Lawrence Livermore Laboratory, UCRL-52435, March 1978.

16. Edwards, L.L. “Computer Simulations of a 1/5-Scale Experiment of a Mark I Boiling

Water Reactor Pressure-Suppression System under Hypothetical LOCA Conditions,” 

Lawrence Livermore Laboratory, UCID-17782, May 1978.

17. Edwards, L.L., et al., “CHAMP: A Coupled HEMP and Multifield Eulerian Program for 

Fluid Flow Simulations,” Lawrence Livermore Laboratory, UCRL-52444, April 1978.

18. Anderson, W.G., P.W. Huber and A.A. Sonin, “Modeling of Pool Swell Dynamics,” 

U.S. Nuclear Regulatory Commission, NUREG-0288, March 1977.

19. Huber, P.W., et al., “Considerations in Small-Scale Modeling of Pool Swell in BWR 

Containments,” U.S. Nuclear Regulatory Commission, NUREG/CR-1143, 

November 1979 (ADAMS Accession No. ML073520012).

20. Chen, C. and Dhir, V.K., “Hydrodynamics of a Bubble Formed at a Vent Pipe Exit,” 

International Journal of Multiphase Flow, Vol. 8 No. 2, pp.147-163, 1982.

21. Aust, E., G.F. Schultheiss, D. Seeliger, and E.W. McCauley, “Experimental Results 

About Dynamic Load Mitigation for BWR Pressure Suppression Containments under 

LOCA Conditions,” 7th International Conference on SMiRT, Chicago, Illinois, 1983.

22. Aust, E. and D. Seeliger, “Pool Dynamics and Dynamic Loads in Pressure Suppression 

Containment Systems,” Transactions of the American Nuclear Society, Vol. 41, 

pp. 696–699, 1982.

23. Wikdahl, C.E., “Marvikenreaktorn - ett industripolitiskt utvecklingsprojekt i otakt med 

tiden,” SKI Rapport Technical Report, 2007.

24. Laine, J., “Condensation Pool Experiments with Non-Condensable Gas,” Lappeenranta 

University of Technology, TOKE-2/2002, December 2002 (ADAMS Accession 

No. ML101410625).

25. Laine, J. “Supplementing Condensation Pool Experiments with Non-Condensable Gas,” 

Lappeenranta University of Technology, YTY-01/2002, October 2002.

26. Tanskanen, V. and M. Puustinen, “CFD Simulation of Air Discharge Tests in the 

PPOOLEX Facility,” Lappeenranta University of Technology, NKS-169, July 2008.

27. Timperi, A, et al., “Numerical Analyses of a Water Pool Under Loadings Caused by a 

Condensation Induced Water Hammer,” VTT Industrial Systems, Finland, NKS-96, 

March 2004.

28. Timperi, A., et al., “Analysis of Loads and Fluid-Structure Interactions in a Condensation 

Pool,” VTT Technical Research Center of Finland, NKS-154, April 2007.
89

29. Pattikangas, T., et al., “CFD and FEM Modeling of PPOOLEX Experiments,” VTT 

Technical Research Center of Finland, NKS-236, January 2011.

30. Timperi, A., et al., “Modeling of Interaction of Multiple Vent Pipes in a Pressure 

Suppression Pool,” VTT Technical Research Center of Finland, NKS-264, April 2012.

31. Tanskanen, V., “CFD Modeling of Direct Contact Condensation in Suppression Pools by 

Applying Condensation Models of Separated Flow,” Lappeenranta University of 

Technology, Thesis for the Degree of Doctor of Science (Technology), March 2012.

32. Reference Deleted

33. Ishii, M., et al., “Experimental Measurement of Suppression Pool Void Distribution during 

Blowdown in Support of Generic Issue 193,” U.S. Nuclear Regulatory Commission, 

NUREG/CR-7186, September 2014 (ADAMS Accession No. ML14255A123).

34. Smith, B.L., “A Numerical Investigation of Three-Dimensional Flows in Large Volumes in 

the Context of Passive Containment Cooling in BWRs,” Nuclear Engineering and 

Design, Vol. 237, No. 11, pp. 1175–1184, 2007.

35. Hart, J., et al., “TEPSS–Technology Enhancement for Passive Safety Systems,” Nuclear 

Engineering and Design, Vol. 209, No. 1-3, pp. 243–252, 2001.

36. Meier, M., “Numerical and Experimental Study of Large Steam-Air Bubbles Injected in a 

Water Pool,” Swiss Federal Institute of Technology Zurich, PhD Thesis No. 13091, 1999.

37. Meier, M., M. Andreani, and G. Yadigaroglu, “Experimental Study of Large Steam-Air 

Bubbles Condensing in a Suppression Pool,” International Mechanical Engineering 

Congress and Exposition ’98, ASME, Anaheim, 1998.

38. “Experience and Qualifications for Nuclear Power Plant Studies,” Alden Research 

Laboratories, Inc., January 2009.

39. Padmamabhan, M., “Hydraulic Performance of Pump Suction Inlets for Emergency Core 

Cooling Systems in Boiling Water Reactors,” U.S. Nuclear Regulatory Commission, 

NUREG/CR-2772, June 1982 (ADAMS Accession No. ML112440074).

40. Kamath, P.S. et al., “An Assessment of Residual Heat Removal and Containment Spray 

Pump Performance Under Air and Debris Ingesting Conditions,” U.S. Nuclear Regulatory 

Commission, NUREG/CR-2792, September 1982 (ADAMS Accession 

No. ML100110155)

41. “NRC Generic Letter 2008-01: Managing Gas Accumulation in Emergency Core Cooling 

Decay Heat Removal and Containment Spray Systems,” U.S. Nuclear Regulatory 

Commission, December 2007 (ADAMS Accession No. ML072910759).

42. Rogers, R.M., “BWR Owners’ Group Technical Report ECCS Pumps Suction Void 

Fraction Study,” General Electric Hitachi Nuclear Energy, 0000-0101-3794-R0-NP, April 

2009.
90

43. Killen, J.M and J.M. Wetzel, “Effect of Air Ingestion on Performance of a Centrifugal 

Pump,” University of Minnesota, Final Report for Contract N00024-80-C-5944, Project 

Report No. 201, July 1981.

44. “Report of the Expert Panel on the Effect of Gas Accumulation on Pumps—‘The Pump 

Roadmap,’” 1026498 Final Report, Electric Power Research Institute, August 2012.

45. “Guidelines for Effective Prevention and Management of System Gas Accumulation,” 

NEI 09-10 (Rev1a-A), Nuclear Energy Institute, April 2013 (ADAMS Accession 

No. ML13136A129).

46. Sonin, A.A. and P.W. Huber, “On the Scaling Laws for Air Clearing in Water-Type 

Pressure Suppression Systems,” Journal of Heat Transfer, Vol. 100, pp. 601-604, 

November 1978.

47. Emami, A. and C. Briens, “Study of Downward Gas Jets into a Liquid,” AIChE Journal, 

Vol. 54, No. 9, September 2008.

48. McCauley, E., et al., “Multivent Effects in a Large Scale Boiling Water Reactor Pressure 

Suppression System,” Lawrence Livermore Laboratory, UCRL-91369, Fifth International

Meeting on Thermal Nuclear Reactor Safety, September 1984.

49. Puustinen, M and J. Laine, “Characterizing Experiments of the PPOOLEX Test Facility”, 

Lappeenranta University of Technology, Finland, NKS-167, July 2008.

50. Zielke, W. and H-D Perko, “Gas Release in Transient Pipe Flow,” Pressure Surges—

Proceedings of the 6th International Conference, BHRA Fluid Engineering, Cranfield, 

Bedford, UK, pp. 3-14, 1990.

51. Kalra, A., “Task 5—Effects of Non-Condensable Gases on Seals (CVDS Pumps),” BWR 

Owners Group, BWROG-TP-12-013, Revision 0, August 2012 (ADAMS Accession 

No. ML12300A222).

52. Hutcherson, M.N. “Contribution to the Theory of Two-Phase Blowdown Phenomenon,” 

Argonne National Laboratory, ANL-75-82, December 1975.

53. Kreith, F, Principles of Heat Transfer, International Textbook Company, Scranton, PA, 

1965.




NUREG-2196 BWR ECCS Pump Suction Concerns following a LOCA