Technology Administration

National Institute of

Standards and Tectinology

November 1995

FIRE SUPPRESSION

SYSTEM PERFORMANCE

O F

ALTERNATIVE AGENTS

IN AIRCRAFT ENGINE

AND

DRY BAY

LABORATORY

SIMULATIONS

SP 890: VOL. I

RICHARD G. GANN, EDITOR

Certain connmercial products

are identified in this report in

order to specify adequately the equipment used. Such identification does not imply recommendation by the National

Institute of Standards and

Technology, nor does it imply

that this equipment is the best

available for the purpose.

Building and Fire Research Laboratory

National Institute of Standards

and Tectinology

Gaittiersburg, MD 20899

U.S. Department of Commerce

Ronald H. Brown, Secretary

Tectinology Administration

Mary L. Good,

Under Secretary for Techinology

National Institute of Standards

and Tectinology

Arati Prabtiakar, Director

November 1995
National Institute of Standards U.S. Government Printing Office

and Technology Washington: 1995

Special Publication 890, Vol. I Natl. Inst. Stand. Technol.

Spec. Publ. 890, Vol. I 781 pages (Nov. 1995)

CODEN: NSPUE2

For sale by the Superintendent

of Documents

U.S. Government Printing Office

Washington, DC 20402
EXECUTIVE SUMMARY

Richard G. Gann

Building and Fire Research Laboratory

Halon 1301 (CF3Br), one of the chemicals identified as detrimental to stratospheric ozone, had become

the choice for suppressing in-flight fires in nearly all types of aircraft. Production of new halon 1301 was stopped on January 1, 1994, and efforts are underway to identify near-term replacements for

critical applications, focussing on available or currently emerging chemicals and technologies. In

particular, the three military services and the Federal Aviation Administration (FAA) have pooled

resources to provide solutions for two applications: engine nacelles and dry (avionics) bays, while

realizing that there are other aircraft areas also in need of protection. This project was managed at Wright Patterson Air Force Base (WPAFB), with oversight provided by a Technology Transition Team

of the four sponsors. The first major objective of the program was to identify the optimal available alternative fluid(s) for use in suppressing fires in aircraft engine nacelles and dry (avionics) bays. In October, 1993,

based on extensive laboratory research and real-scale testing at WPAFB, the sponsors decided on a reduced list of candidates for each application; for engine nacelles: C2HF5 (HFC- 125), C3HF-y

(HFC-227ea), and CF3I; for dry bays: C2HF5, C3F8 (FC-218), and CF3I. Much of the laboratory-scale

research leading to that decision has been described in NIST Special Publication 861, Evaluation of

Alternative In-FUght Fire Suppressants for Full-Scale Testing in Simulated Aircraft Engine Nacelles

and Dry Bays. That report documents the comprehensive experimental program to screen the

performance of possible suppressant chemicals as a means to identify the best candidates for

subsequent full-scale aircraft fire extinguishment evaluation at Wright Laboratory, and addresses the

compatibility of these agents with flight systems, people, and the environment. In particular, apparatus

and measurement methods suited to aircraft applications are carefully described, and extensive

performance data are provided and analyzed. The reader is referred to that report as a prerequisite and

companion to the current document.

The subsequent research described in this report falls into two broad categories:

Part 1 : Knowledge to help differentiate among chemicals, leading to selection of the optimal

currently available option(s) for in-flight fire suppression

Fire Suppression Efficiency. Most of the laboratory-scale information was reported in SP 861. In new studies in a deflagration/detonation tube, simulating fire suppression in a dry bay, FC-218

provided the best overall performance; HFC- 125 was comparable under many conditions. CF3I had

the greatest positive impact at low addition levels, but showed non-monotonic behavior of flame speed

and shock pressure at higher levels. In fire suppression efficiency measurements in a turbulent spray

burner simulating engine nacelle fires, CF3I was found to be more efficient than either HFC- 125 or HFC-227ea. However, at an elevated temperature of 150 °C, the three chemicals performed equally on a mass basis.
iv EXECUTIVE SUMMARY

Stability During Storage. There should be no problems with designing long-term storage capability

for any of the four agents. The agents were tested for chemical stability in the presence of likely

storage materials for over a year under typical in-use pressures and temperatures. The stability of

HFC-125, HFC-227ea, and FC-218 was excellent. Samples of CF3I were observed to be stable at 23 °C; however, during exposure at 100 °C and especially at 150 °C, small amounts of CO2 and an unknown chemical appeared whose concentration increased with exposure time. Each of the four

chemicals is compatible with a choice of materials for the storage containers: gasket materials,

lubricants, and metals. Each agent caused only minor swelling in at least three of the elastomers and

at most moderate swelling in the greases. The long-term deformation of the seven elastomers showed

that at least two elastomers are compatible with each of the agents. The long-term stability of the seven elastomers, evaluated using compression set measurements, showed that at least two elastomers were not subject to excessive permanent deformation in each of the four agents. Long-term immersion

of coupons of container metals at storage temperatures and pressure in each of the four candidate

agents or halon 1301 produced little mass change or visual corrosion, although CF3I showed some

small interaction with three of the metals at high temperature. Slow strain rate tests of the metals in four of the agents at 20 °C and 150 °C and 5.86 MPa showed no difference from metal samples

immersed in inert argon. CF3I was not compatible with a titanium alloy at elevated temperature.

Safety Following Discharge. Predictions of the production of HF during fire suppression indicate that

fires in dry bays are suppressed sufficiently quickly that only small amounts will be formed. For

engine nacelle fires, the model developed here predicts that HFC-227ea, HFC-125 and FC-218 would

produce similar amounts, while the more efficient CF3I would produce far less. Exposure to surfaces

heated by the fire would produce more HF from CF3I than from an equal amount of the other three

chemicals. Samples of aircraft materials that might be located near or downstream of an engine

nacelle fire were immersed in 1 % or 10 % aqueous HF and then stored at 100 % relative humidity for

30 days at 22 °C. No significant degradation was seen. While accidentally-discharged CF3I will

decompose under both normal outdoor and indoor lighting, laboratory measurements and dispersion

modeling show that the concentrations of potentially toxic photolysis products (HF, COF2) are not

likely to be sufficient to hinder prompt escape.

Discharge Performance. All four of the chemicals can be expected to discharge and disperse well

from their storage and distribution systems at temperatures near 20 °C. At the lower temperatures

experienced during high altitude flight or cold weather operation, CF3I and HFC-227ea, with their

higher boiling points, would not disperse as well as the other two chemicals.

Optimal Currently Available Option. Based on the results available in October, 1994, we recom- mended the selection of HFC-125 as the optimal candidate for Phase III examination for both engine

nacelle and dry bay fire suppression. FC-218 possesses an extremely long environmental lifetime. While CF3I was the most efficient suppressant, being virtually a drop-in replacement for halon 1301 in some tests, it had three drawbacks: its inhalation toxicity in cardiac sensitization testing, inconclusive

stability and materials compatibility data, and a relatively high boiling point. HFC-227ea has a similar

boiling point, and thus would also perform less well at low temperatures. The knowledge that has

accrued in the final year of the project has not changed our perspective on this recommendation.

During the fall 1994 meetings of the Technology Transition Team, these data and the results of an

extensive and carefully constructed series of real-scale live-fire tests at Wright Patterson Air Force

Base led the Team to recommend HFC-125 as the candidate for Phase III evaluation for both applications.
EXECUTIVE SUMMARY

Part 2: Knowledge to assist in the development of engineering design criteria and

suppressant system certification

Agent Discharge Behavior. The rate at which a suppressant will emerge from its storage container

depends on the thermodynamic properties of the stored fluid and any pressurizing gas as well as the

initial conditions in the container. Effective design of the suppression hardware requires quantitative

performance measures for these chemicals. The NIST computer code PROFISSY accurately calculates

binary vapor-liquid equilibria within the storage bottle, data needed by storage system designers.

Laboratory data show that the nitrogen dissolved in the stored liquid agent significantly affects the

agent discharge whether in a direct release system for dry bays or a piping system for engine nacelles. A NIST-developed storage bottle discharge model, which incorporates nitrogen degassing, generally

predicts agent discharge times to within a factor of two, but occasionally a factor of four. The only

required inputs for the model are the initial conditions in the vessel. For plumbed engine nacelle

systems, a new, validated model can be used to facilitate calculations of transient, two-phase pipe

flow. We have developed a flow chart to organize the use of these tools into a coherent process for

optimal design of a new discharge system.

Engine Nacelle Fire Suppression Guidance. The selection of the mass of agent to be stored on an

aircraft should be based on the amount needed to quench the worst realistic fire. For engine nacelles

with ribs and other obstructions, this is a baffle-stabilized pool fire. Heating the air stream and, to a

lesser extent, increasing the pressure, increases the mass of agent needed for flame extinction, while

the fuel flow had no significant effect. A step-by-step procedure has been developed that gives

guidance in determining fire suppressant concentrations and discharge times for flame extinguishment.

It shows the relative importance of agent injection duration, air flow and velocity, agent/air mixing

mode, and fire scenario. A reasonable target concentration for an agent in the fire zone is that needed

to quench the most flammable fuel/air mixture, both ensuring flame suppression and preventing re- ignition during the period of injection.

Real-Time Concentration Measurement. Determination of the dispersion of the suppressant

following discharge requires measurements of its concentration that are well-resolved in both time and

space. We examined two instruments (based on thermal conductivity and infrared absorption) in order

to assess their ability to perform in situ measurements with the ~1 ms time resolution needed for dry

bay applications. Neither performed well. Both were improved extensively and showed potential.

However, both need further work to be useful in practical systems. A review of the sensing literature shows a number of alternative approaches, but none that could be accurately adapted to this application without a significant development and testing effort. The most promising are time-resolved mass

spectrometry and mid- and near-infrared absorption combined with fiber optics to provide easy access and the needed spatial resolution.

Certification Guidance. HFC-125 closely replicates the physical and dispersion properties of halon

1301. Thus it is an excellent simulant for hardware development and can be used to certify those

engine nacelle fire suppression systems that still rely on halon 1301. The mixing time for agent

entrainment behind an obstacle is different under non-fire conditions than for fire conditions. A

method for using the non-fire data to approximate the fire suppression concentration has been

developed.

Interaction with Metal Fires. In laboratory experiments, none of the four alternative chemicals nor

halon 1301 showed exacerbation of burning of magnesium or titanium rods. It is not explicitly known
vi EXECUTIVE SUMMARY

why the fiare-ups observed during the introduction of halon 1301 to real metal fires were not observed

here. However, it may be useful to know that in the circumstances replicated in the laboratory tests, the alternative agents did not worsen the combustion relative to that with halon 1301 present.
VII

TABLE OF CONTENTS

Page

Executive Summary

R.G. Gann iii

PART 1

1. Introduction

R.G. Gann 1

2. Suppression of High Speed Flames and Quasi-detonations

G.W. Gmurczyk, W.L Grosshandler 9

3. Photodegradation of CF3I

M.R. Nyden 77

4. Effect of Suppressants on Metal Fires

T.J. Ohlenuller. J R. Shields 97

5. Compatibility with Metals

M.R. Stoudt. J.L Fink, J.F. Dante, R.E. Richer 121

6. Interactions of Agents with Elastomers

G.B. McKenna, F. Horkay, RH. Verdier, W.K. Waldron, Jr. 201

7. Agent Stability Under Storage R.H. Harris 249

PART 2

8. Optimization of System Discharge

J.C. Yang, T.G. Cleary, I. Vazquez, C.I. Boyer, M.D. King, B.D. Breuel, C.A. Womeldorf, W.L. Grosshandler, M.L. Huber, L. Weber, G. Gmurczyk 407

(continued in Volume II)

1

1. INTRODUCTION

Richard G. Gann

Building and Fire Research Laboratory

Contents

Page

1. INTRODUCTION 1

1.1 Background 2

1.1.1 Stratospheric Ozone Depletion 2

1.1.2 Aircraft Fire Suppression 2

1.1.3 Prior Work 3

1 .2 Objectives and Task Summary 3

1.3 Acknowledgments 4

1.4 References 5
2 1. INTRODUCTION

1.1 Background

1.1.1 Stratospheric Ozone Depletion. In a remarkably short period of time, the world has identified,

responded to, and ameliorated a new threat to the global climate (WMO, 1995). Following a mechanism first proposed by Rowland and Molina in 1974, chemically stable chlorine-, bromine-, and

iodine-containing molecules rise to the stratosphere and are quantitatively photodissociated by

ultraviolet radiation. The halogen atoms then catalytically convert ozone (O3) molecules, whose

chemistry shields the earth's surface from excess ultraviolet radiation, into oxygen (O2) molecules,

which have no such filtration effect. The evidence supporting this hypothesis soon became substantial,

and the international political community produced a landmark agreement in 1987, the Montreal

Protocol on Substances That Deplete the Ozone Layer. Subsequent international amendments to this and, domestically, the U.S. Clean Air Act of 1990 have led to restrictions on both production and use of identified ozone-depleting substances (ODSs).

Halon 1301 (CF3Br) is one of the detrimental compounds identified in this process. It is used

principally as a fire suppressant, and had become the choice for many applications where effective,

efficient, and clean fire control is needed. Ideally, a typical year's production of halon 1301 would be

loaded into fire suppression systems, where it would reside for many years until called upon to quench

the infrequent fire. In practice, most of the chemical was released sooner, in training exercises, during

system testing, from discharge of the suppression system when no fire was present, or from negligent

handling. In the evolving regulatory process, future uses of halon 1301 were restricted, and its production was stopped on January 1, 1994 with limited allowances made for certain developing

countries.

Many of the systems being protected by halon 1301 are essential to military readiness; and in 1991, the Deputy Director, Defense Research and Engineering in the Department of Defense initiated an urgent research program to identify near-term alternatives for weapons systems use, mostly by

September 1996. This research focusses on commercially available or currently emerging chemicals

and technologies.

1.1.2 Aircraft Fire Suppression. One of the most important uses of halon 1301 is the suppression of

in-flight fires in nearly all types of aircraft. The three military services and the Federal Aviation

Administration (FAA) have pooled resources to provide solutions for a number of problems as the

aviation community reduces and eventually eliminates its reliance on halon 1301. The research and

engineering projects have been carried out at laboratories of the Army, Navy, Air Force and FAA,

and at the National Institute of Standards and Technology (NIST). The research program focussed on two applications: engine nacelles and dry (avionics) bays, while realizing that there are other aircraft areas also in need of protection.

The requirements for engine nacelle fire protection in commercial and military aircraft are similar. The interior of the nacelle is nominally tubular in shape, often with ribs and multiple other obstructions. Two types of fires are possible: spray fires, such as from a ruptured fuel line or hydraulic fluid

line, and pool fires, resulting from either liquid settling on the bottom surface of the nacelle. Either

fire type can be stabilized behind an obstruction, posing an additional difficulty for the fire suppressant. Currently, halon 1301 is stored in cylinders at pressures of ca. 2 MPa - 4 MPa. When a thermal

sensor detects an abnormal, overheated condition, the flight of the airplane is leveled and the

appropriate halon 1301 bottle is discharged. The part gas, part liquid agent flows through up to several meters of tubing to the engine nacelle. Once released, it rapidly fills the nacelle volume. The

certification process requires that enough agent be available to maintain a minimum concentration {ca.

6 % by volume) throughout the nacelle for a minimum time interval (0.5 s) to ensure that the fire will
1. INTRODUCTION 3

be extinguished and not re-ignite. Many systems have a duplicate bottle to be used as a back-up

should the first shot be unsuccessful.

The dry bay fire is specific to military aircraft. These bays are cluttered compartments, typically

0.2 m"^ - 3 m"* in volume, located along the wings and fuselage. An incident incendiary shell could

penetrate both the bay wall and that of an adjacent fuel storage volume, leading to a deflagration and

loss of the aircraft. Such fires must be quenched automatically within a few tens of milliseconds.

1.1.3 Prior Work. The first major objective of the four-agency program was to identify the optimal

available alternative fluid(s) for use in suppressing fires in aircraft engine nacelles and dry (avionics)

bays. This project was managed at Wright Patterson Air Force Base, with oversight provided by a Technology Transition Team of the four sponsors. In October, 1993, based on extensive laboratory

research and real-scale testing at Wright-Patterson Air Force Base, the sponsors decided on a reduced

list of candidates for each application:

Engine Nacelle Dry Bay

C2HF5 (HFC- 125, pentafluoroethane) C2HF5 (HFC- 125, pentafluoroethane)

C3HF7 (HFC-227ea, C3F8 (FC-218, octafluoropropane)

1 , 1 ,1 ,2,3,3,3-heptafluoropropane)

CF3I (trifluoroiodomethane) CF3I (trifluoroiodomethane)

Much of the laboratory-scale research leading to that decision was performed at NIST and has

been described in Evaluation of Alternative In-Flight Fire Suppressants for Full-Scale Testing in Simulated Aircraft Engine Nacelles and Dry Bays (Grosshandler et al., 1994). That report documents

the comprehensive experimental program to screen the performance of possible suppressant chemicals

as a means to identify the best candidates for subsequent full-scale aircraft fire extinguishment

evaluation at Wright Laboratory, and addresses the compatibility of these agents with flight systems,

people, and the environment. In particular, apparatus and measurement methods suited to aircraft applications are carefully described, and extensive performance data are provided and analyzed. The

reader is referred to that report as a prerequisite and companion to the current document.

1.2 Objectives and Task Summary

The research described in this report has multiple origins, but falls into two broad categories:

Part 1: Knowledge to help differentiate among chemicals, leading to selection of the optimal

currently available option(s) for in-flight fire suppression. These projects are described in Part 1. Section 2 of Part 1 (which follows this introduction) provides further data and explanation of

previously identified phenomena regarding potential suppression of dry bay fires. Section 3 addresses

the potential for degradation of photosensitive CF3I following an accidental discharge. Research on

the effectiveness of halon 1301 alternatives on metal fires is discussed in Section 4. Sections 5 and 6

continue to longer exposure times (from Grosshandler et al., 1994) the measurements and analysis of

the compatibility of the candidate agents with metals and organic materials that might be used in storage vessels. Section 7 similarly extends the prior study of the stability of the agents themselves

during long-term storage.
4 1. INTRODUCTION

Part 2: Knowledge to assist in the development of engineering design criteria and suppressant system certification. Section 8 develops understanding of and equations for calculating the

dispersion of the agent from the storage bottle and subsequent plumbing. Section 9 provides guidance on the agent concentration requirements for flame suppression in engine nacelles. Section 10 develops

a predictive capability for the formation of toxic and corrosive hydrogen fluoride (HF) during fire suppression. Section 11 reviews approaches to making the high-speed, real-time measurements of

suppressant concentration needed both for research and certification, and describes our research on two

of the approaches. Section 12 documents our search for and identification of a surrogate chemical for

certification testing of aircraft fire suppression systems in which halon 1301 is still used.

Section 13 provides a summary of the results and puts them in the context of the fire suppression

problems under consideration.

1.3 Acknowledgments

The work described in Sections 2 and 5-11

Government:

Wright Laboratory

Flight Dynamics Directorate

Vehicle Subsystems Division

Survivability Enhancement Branch

Mr. Michael Bennett

Lt. Gretchen Brockfeld

Lt. Gregg Caggianelli

Army Aviation and Troop Command

Mr. Michael Clauson

co-sponsored by four agencies of the U.S.

Naval Air Systems Command

Subsystems Branch

Fuel and Fluid Power Systems Section

Mr. James Homan

Department of Transportation

Federal Aviation Administration

Technical Center

Ms. Louise Speitel

The work in Sections 4 and 12 were sponsored by the Naval Air Systems Command, under the

active guidance of Mr. David Thurston and Mr. William Leach, respectively. The work in Section 3 was performed under NIST internal funding.

The overall Program Manager was Mr. Michael Bennett of the Wright Laboratory. His energy

and engineering perception were essential to the successful completion of the research reported here.

Lt. Gregg Caggianelli and Lt. Brett Poole provided expert assistance in carrying out the full-scale instrumentation tests reported in Section 1 1

.

Numerous staff from industrial firms provided advice, materials, and their time. They are credited

in the individual sections.

Finally, full credit is due to Ms. Paula Garrett for her extensive efforts in bringing this manuscript

to reality.
1. INTRODUCTION

1.4 References

Grosshandler, W.L., Gann, R.G., and Pitts, W.M., Eds., Evaluation of Alternative In-Flight Fire

Suppressants for Full-Scale Testing in Simulated Aircraft Engine Nacelles and Dry Bays, Special

Publication 861, National Institute of Standards and Technology, Gaithersburg, MD, 1994.

Molina, M.J. and Rowland, F.S., "Stratospheric Sink for Chlorofluoromethanes: Chlorine AtomCatalyzed Destruction of Ozone," Nature 249, 810 (1974).

Executive Summary, Scientific Assessment of Ozone Depletion: 1994, World Meteorological

Organization, Geneva, 1995^

PART 1

AGENT SELECTION PROCESS

\

\

9

2. SUPPRESSION OF HIGH SPEED FLAMES

AND QUASI-DETONATIONS

Grzegorz W. Gmurczyk

Science Applications International Corporation

and

William L. Grosshandler

Building and Fire Research Laboratory

Contents

Page

2. SUPPRESSION OF HIGH SPEED FLAMES

AND QUASI-DETONATIONS 9

2.1 Background 10

2.2 Technical Approach and Task Summary 11 2.3 Experimental Set-up 12

2.3.1 Design 12

2.3.2 Operation 16

2.4 Experimental Results 16

2.4.1 Conditions 16

2.4.2 Measurement Signals 20

2.4.3 Combustion Characteristics 20

2.4.3.1 Ethene/Air Mixtures 23

2.4.3.2 Propane/Air Mixtures 26

2.4.4 Suppression Characteristics 26

2.4.4. 1 C2HF5 Performance 32

2.4.4.2 C3Fg Performance 43

2.4.4.3 CF3I Performance 52

2.4.4.4 Relative Performance of C2HF5, C3F8, and CF3I 61

2.4.5 Uncertainty Analysis 61

2.5 Summary of Results 66

2.6 Conclusions 73

2.7 Acknowledgments 74

2.8 References 74
10 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

2.1 Background

A dry bay is a normally confined space adjacent to a fuel tank in which a combustible mixture and an

ignition source could co-exist following penetration by an anti-aircraft projectile. They vary considerably in volume, typically being in the range of 0.2 to 3.0 m^ They are located in the wings and

fuselage, and their shape is most often irregular. Aspect ratios up to 10:1 are not uncommon. The

bays may or may not be ventilated, and are usually cluttered with electronic, hydraulic and mechanical

components. Compared to the events leading to engine nacelle fire suppression, the required timing is two orders-of-magnitude faster for dry bay protection.

The previous study using a deflagration/detonation tube (Grosshandler et ai, 1994) was concerned

with establishing a comprehensive experimental program to screen the performance of over a dozen

agents. The experiments were designed to cover the range of conditions that might occur in a dry bay.

Although actual measurements of fuel concentrations in a dry bay during live-fire testing have never been made, one could envision a worst-case situation in which the fuel is vaporized and partially

premixed with the air just prior to ignition, producing a rapidly moving turbulent flame. If the

suppressing agent were not well mixed and the dry bay geometry were conducive, the turbulent flame

could accelerate, generating a shock wave ahead of it and transitioning to a detonation before

encountering the agent. Ethene was chosen as the fuel in the previous study because it was known to detonate easier than many other hydrocarbons. This provided the most severe test for all the agents

under conditions that were not duplicated in any of the other bench-scale studies. Three chemicals were selected by the Technology Transition Team to be included in the complete

full-scale experimental matrix conducted at Wright Patterson AFB (Carbaugh, 1993): HFC-125,

FC-218 and CF3I. Measurements in the previous detonation/deflagration tube study revealed a volume

factor (VF) for FC-218 that was lower than the VF for HFC-125. More significant was the high over- pressure experienced for HFC-125 mass fractions below 25 %. With the deflagration/

detonation tube operating with a lean ethene/air mixture, 21 % HFC-125 in the test section produced a quasi-detonation with a pressure ratio of 37:1, double the pressure build-up when no agent was

present. The FC-218 behaved quite differently, and effectively reduced the pressure ratio at concentrations near 21 %. These results favored the selection of FC-218 over HFC-125 for dry bay protection.

Few flame suppression experiments had been conducted with CF3I that were applicable to dry

bays. The previous detonation/deflagration tube results indicated an unusual behavior that could also

be observed with CF3Br, but to a lesser extent. Both chemicals were equally effective in low

concentrations at reducing the pressure build-up. At mass fractions greater than about 10 % the

chemistry is altered and the pressures began to rise. Increasing the CF3Br concentration benefitted

suppression at mass fractions greater than 20 %, and total suppression of the flame occurred above

30 %. Pressure ratios in the CF3I tests continued to rise up to a mass fraction of 30 %, reaching a

pressure greater than the uninhibited mixture. That is, adding 30 % CF3I to a lean ethene/air flame

exacerbated the situation. It took a mass fraction of almost 45 % to completely suppress the pressure

build-up.

The maximum pressure ratios observed in full-scale live-fire testing of uninhibited propane air mixtures are less than 7:1, and photographic evidence from full-scale dry bay testing suggests that turbulent flame speeds are below 300 m/s (Bennett, 1993). The previous experiments created

uninhibited pressure ratios up to 25:1 and quasi-detonation velocities over 1100 m/s. By changing the

fuel from ethene to propane, and by adjusting the geometry of the detonation/deflagration tube, the pressure ratio and velocity of the combustion wave can be reduced, allowing determination of whether

or not a dangerous over-pressure arises during suppression under conditions that represent more likely threat scenarios.
2. HIGH SPEED FLAMES AND QUASI-DETONATIONS 11

The specific objectives of the current research project are the following:

a. To determine the effectiveness of HFC-125, relative to FC-218, in suppressing high speed

turbulent propane/air flames using the detonation/deflagration tube apparatus.

b. To determine the conditions in the detonation/deflagration tube (equivalence ratio, tube

geometry) which lead to excessive pressure build-up during suppression by HFC-125 of pro- pane/air mixtures initially at room temperature and pressure.

c. To determine the effectiveness of CF3I, relative to FC-218, in suppressing high speed turbulent

propane/air flames using the detonation/deflagration tube apparatus.

d. To recommend a ranking of the three agents for full-scale dry bay applications based upon the current and previous suppression experiments.

2.2 Technical Approach and Task Summary

The detonation/deflagration tube is a unique apparatus for evaluating a fire suppressant in a highly

dynamic situation. A shock wave precedes the flame, with obstructions in the flow, if any, promoting

intense mixing of the fresh reactants with the combustion products and causing the pressure waves to interact with the mixing region. Given enough distance, the initially subsonic flame (deflagration) can

accelerate dramatically, reaching the supersonic regime (detonation), and increasing the temperature of

the reaction zone behind the shock as well as further adding to the heat release rate. Depending upon

the geometric details, the wave can approach its theoretical Chapman-Jouguet velocity and accompanying high pressure ratio. Even a slight variation in composition of the reactants near the limit of

detonation can cause a dramatic change in the wave velocity and cause destructive pressures to be

attained.

Extensive literature exists describing the kinetics and dynamics of flame/shock wave systems

formed within classical detonation tubes (e.g., Lefebvre et al, 1992; Nettleton, 1987; Lee, 1984; Baker

et al, 1983; Westbrook, 1982). Chapman and Wheeler (1926) were the first to note that a methane/air

flame could be accelerated to a terminal velocity in a shorter distance within a circular tube by placing

obstacles into the flow. Lee et al. (1984) built on this observation to study quasi-detonations in hydrogen/air and hydrocarbon/air mixtures.

A quasi-detonation propagates more slowly than a true detonation due to pressure losses in the

flow, but its structure is more complex than a true detonation, and the mechanism of its propagation is not fully understood. Although obstructed flow is more difficult to analyze than the flow in a smoothwalled tube, the complex obstructed arrangement has been investigated here because it more closely

simulates a potentially damaging condition in the dry bay. The present construction of the detonation/

deflagration tube facility is designed to compare both obstructed and unobstructed flow conditions.

Because the fire extinguishant is unlikely to be released prior to the establishment of a turbulent

flame, the traditional experiment in which the flame inhibitor is premixed with the fuel and air prior to ignition does not replicate the chemistry critical to the actual situation. The NIST facility has been

designed to provide a quiescent air/fuel/agent environment into which an uninhibited, fully turbulent

flame propagates. The desire to rapidly suppress a flame and the associated pressure build up in such

a situation is the primary objective behind this study. An ideal agent would quench the exothermic

chemistry at low concentrations and prevent the further build-up of pressure waves; however, the
12 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

incident mechanical shock can only be reduced through viscous damping, and remains even when the

chemistry has been satisfactorily interrupted. Shock/flame wave velocity and pressure ratio were the

two dependent parameters that were measured as a means to characterize the extent of flame

suppression. The velocity was determined by the time it took for the pressure wave to travel the

distance between two pressure transducers. The pressure ratio was evaluated from the average

amplitude of the first pressure pulse recorded by each transducer, normalized by the initial pressure. A number of specific tasks were performed using the detonation/deflagration tube apparatus.

First, experiments were conducted to determine the range of Mach numbers and pressure ratios obtainable in the tube using propane rather than ethene. The objective of this task was to produce in a

predictable manner high speed turbulent flames (with Mach numbers between 1 and 2 and pressure

ratios between 3 and 10) by manipulating the initial conditions in the tube. The variables at our

disposal were the propane/air ratio, the fuel partial pressure, and the length of the tube and internal

spiral. The conditions which led to repeatable subsonic flames were noted. Next, the pressure ratios and Mach numbers were measured in lean, stoichiometric and rich propane/air mixtures over a range

of HFC- 125, FC-218 and CF3I mass fractions in the test section of the tube. The initial conditions were chosen to produce uninhibited Mach numbers below 2.0 and pressure ratios smaller than 10.

2.3 Experimental Set-up

2.3.1 Design. The two-sectional detonation/deflagration tube was designed (Gmurczyk et ai, 1993,

1994) to examine the performance of the alternative agents in a highly dynamic situation, in which the

pressure effects on the chemistry are thought to be important. Using the detonation/deflagration tube,

the effectiveness of a fire fighting agent in suppressing a high speed, premixed flame or quasidetonation can be rated by the extent to which it decelerates the propagating flame and simultaneously

attenuates the hazardous shock which is always ahead of the flame.

A primary feature of the set-up is that the conditions of the ignition event do not affect the

suppression process itself. Also, because an agent of interest is premixed with the fuel and air in a

section of the tube separated from the ignition event, the influence of entrainment of the agent into the

flame is minimized. The tube is closed to allow the increase in pressure to interact with the combustion chemistry.

The heart of the facility is shown schematically in Figure 1. The left hand side of the picture

shows a fragment of the driver section (flame/shock generation region) of the tube separated by a

partition from the test section (flame/shock suppression/attenuation region) of the tube on the right

hand side of the picture. The flame/shock system propagating within a combustible mixture is fully

established before entering the region occupied by a suppressant premixed with the same combustible

mixture.

The driver section is 5 m long (see Figure 2) and is equipped at the closed end with a spark plug.

This section is filled with the combustible mixture of ethene or propane and air of various compositions. The gas handling system (see Figure 3) consists of a vacuum pumping network; pressurized gas

cylinders for the fuel, oxidizer and agent; and a dual circulating pump (Metal Bellows MB 602 XP).^

Certain trade names and company products are mentioned in the text or identified in an illustration

in order to specify adequately the experimental procedure and equipment used. In no case does such

identification imply recommendation or endorsement by the National Institute of Standards and

Technology, nor does it imply that the products are necessarily the best available for the purpose.
2. HIGH SPEED FLAMES AND QUASI-DETONATIONS 13

1 - SHOCK WAVE, 2 - TURBULENT FLAME, 3 - GATE VALVE,

4 - DRIVER SECTION, 5 - SPIRAL INSERT, 6 - TEST SECTION,

F - FUEL, O - OXIDIZER, A - AGENT

Figure 1 . Schematic of the combustion/suppression process in the detonation/deflagration tube.
2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

c

VP VP

--^15^

oooeooooooo

ooooooooooo o o o o

DS - DRIVER SECTION, TS - TEST SECTION, V - GATE VALVE,

IS - IGNITION SYSTEM, CP - CIRCULATION PUMP,

VP - VACUUM PUMP, E - EXHAUST, F - FUEL, O - OXIDIZER,

A - AGENT, PH - FAST PHOTODIODE, PT - PIEZOELECTRIC

PRESSURE TRANSDUCER, > - BLOCK OF AMPLIFIERS

gure 2. Schematic of the detonation/deflagration tube facility designed and installed at NIST.
HIGH SPEED FLAMES AND QUASI-DETONATIONS

7 7 7

\M

2-J

-H -IXJ

11

1

1 - DRIVER SECTION, 2 - TEST SECTION, 3 - GATE VALVE,

4 - FUEL, 5 - OXIDIZER, 6 - AGENT, 7 - METERING VALVE,

8 - ON-OFF VALVE, 9 - STATIC PRESSURE TRANSDUCER,

10 - DUAL CIRCULATION PUMP, 11 - VACUUM PUMP

Figure 3. Schematic of the detonation/deflagration tube gas handling system.
16 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

The ignition energy is delivered in a micro-explosion of a tin droplet short-circuiting the tips of

nichrome electrodes connected to an 80 V power supply. Spiral-shaped obstnactions made of 6.4 mm

stainless steel rods with a pitch equal to the inner diameter of the tube are inserted into the tube, to produce an area blockage ratio of 44 %, close to the value which is known to promote a high-speed or quasi-detonation regime of combustion.

The second section of the detonation/deflagration tube contained the gaseous agent along with the same fuel/air mixture used in the driver section. The diameter is the same and its length is either

2.5 m or 5.0 m. The flame and shock signals serving to determine velocities and pressures were taken

2.2 m downstream behind the gate valve when the 2.5 m long test section of the tube was installed. The 5 m test section was used without the spiral insert. The additional length was used to eliminate

the reflected shock wave that sometimes interfered with a slower moving primary reaction front. The

flame signals serving to determine velocities in the long tube were taken close to the entrance region

of the test section, 0.3 m downstream behind the gate valve, to better ascertain the immediate impact

of the inhibitors on the flame dynamics; the shock signals were measured 2.2 m into the test section,

which is the same location used for the short tube. The driver and test sections are separated from each other by a 50 mm inner diameter, stainless

steel, gate valve (MDC model GV-2000M-SP), which remains closed until just before ignition. Kistler

model 603B1 dynamic piezoelectric pressure transducers (with dual mode charge amplifiers model

501 OA 10) and Siemens model BPX 65 silicon planar PIN high speed photodiodes are located along

the test section to monitor the strength and speed of the combustion wave. Refer to Figures 4, 5, and 6. Their output was recorded either with a 50 MHz IBM Personal Computer with 32 bit EISA

data acquisition board or with a fast, multi-channel, digital storage oscilloscope (Le Croy

model 9314 M). In the latter case the data were also stored in the computer, since full communication was possible.

2.3.2 Operation. The whole system is evacuated to 10"' Pa before filling the two sections separately

with the desired mixtures, which are attained through the method of static partial pressures measured

with Omega PX811 high accuracy absolute pressure gauges. The fuel/air ratio and total pressures are held constant across the gate valve. After filling, the gases are homogenized independently using a double, spark-free circulating pump, recirculating the entire tube volume a total of 20 times. The

mixtures are then left for five minutes to become quiescent. About 10 s prior to ignition, the gate

valve is opened manually. After ignition, the flame propagates into the driver section and accelerates

quickly due to the intense turbulence created by the interactions of the flow with the obstacles. This

generates a shock wave ahead of the flame. After passing through the open gate valve the

flame/shock system encounters the same combustible mixture and a certain amount of agent in the test section. Depending on the concentration of the agent, the flame may be extinguished (or enhanced)

and the pressure wave may be attenuated (or amplified).

2.4 Experimental Results

2.4.1 Conditions. The following independent parameters were changed during the course of the

experiments:

- type of suppressant (C2HF5, CgFg, and CF3I);

- concentration of suppressant;
HIGH SPEED FLAMES AND QUASI-DETONATIONS

' 1

i

—

V/////////////AVA^ ll^//////////////////A

1 - LOW-NOISE CABLE, 2 - CABLE CONNECTOR, 3 - PIEZOELECTRIC PRESSURE TRANSDUCER, 4 - STUB, 5 - TUBE,

6 - RETAINING NUT, 7 - HOLDER, 8 - SEALS, 9 - SEAL WELD

Figure 4. Schematic of the piezoelectric dynamic pressure transducer mounting.
2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

10

11

2^ I 1 3

4

5

1 - PHOTODIODE, 2 - BNC CABLE, 3 - QUARTZ WINDOW,

4 - "O " RING, 5 - SEAL WELD, 6 - COVER, 7 - BUSH, 8 - CASE

9 - WASHER, 10 - STUB, 11 - TUBE

Figure 5. Schematic of the fast photodiode mounting.
HIGH SPEED FLAMES AND QUASI-DETONATIONS

1 - DIN CONNECTOR, 2 - CIRCUIT BOARD, 3 - OPERATIOANL

AMPLIFIER, 4 - STANDOFF, 5 - SCREW, 6 - TO POWER SUPPLY,

7 - SIGNAL OUTPUT, 8 - GROUND, 9 - BNC CONNECTOR,

10, 11 - CAPACITORS, 12 - NUT, 13 - TO PHOTODIODE

Figure 6. Schematic of the fast photodiode ampHfier.

\
20 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

- type of fuel (ethene or propane);

- equivalence ratio of the combustible mixture (lean, stoichiometric, rich)

- geometry of the tube (2.5 m or 5 m long test section, with or without spiral).

The initial temperature of the mixtures was ambient (22 °C ± 3 °C) and the initial pressure was

100± 0.6 kPa. The oxidizer used in all experiments was breathing grade air. Ethene and propane (CP

grade 99.5 % volume purity) were chosen as the fuels; ethene because it is known that detonations can

be obtamed in a tube of this geometry simply by varying the stoichiometry, and propane because it more closely resembles a jet fuel and can reproduce turbulent flames with pressure ratios under 10:1

and wave speeds down to sonic conditions. The extinguishing compounds were used as supplied by

the manufacturers.

The partial pressure measurements were affected by the uncertainty of the static pressure

transducer (± 0.3 kPa maximum uncertainty after combining non-linearity, hysteresis, repeatability, and

temperature effects), the uncertainty of the digital display device (± 0.015 kPa), the uncertainty

associated with the purity of the gases (± 0.5 % of partial pressure reading), and the uncertainty

associated with possible gas losses in the circulation pump (up to 0.3 % of the partial pressure reading,

in the worst case). Assuming that the errors were directly additive, the partial pressure for any

component in the mixture was uncertain to less than 0.32 kPa plus 0.3 % of the reading.

2.4.2 Measurement Signals. Figure 7 shows representative signals coming from two fast photodiodes which indicate the presence of radiation (peak response at 850 nm) associated with the reaction

front traveling in the tube. Figure 8 displays signals from two piezoelectric pressure transducers which

represent a pressure jump associated with the shock wave ahead of the primary reaction zone. The

time difference between the occurrence of the signals allows one to determine flame and shock

velocities. The amplitude of the pressure signals permits determination of the pressure ratio of the

shock.

The uncertainty of the determination of the shock wave amplitude was affected by the combined

uncertainty of the dynamic pressure transducer (± 1 % of the reading), the combined uncertainty of the

transducer amplifiers (± 0.5 % of the reading), the combined uncertainty of the digital data acquisition

system (± 0.5 %), and the combined uncertainty of the digital readout device (± 0.2 %). Assuming

additivity of errors, the resultant accuracy of determining the shock wave amplitude is ± 2.2 %. The

uncertainty of the determination of the shock time differences was affected by the same elements, as well as the transducer rise time ( < 2 [is). The shock speed can thereby be estimated to be uncertain

to less than ± 4.4 % of the reported reading (accounting for the differential nature of this measurement). The uncertainty of the determination of the flame travel time was affected primarily by the rise time of the photodiode, which is 30 ns. The combined uncertainty of the magnitude of the photodiode

signal is estimated to be ± 2 % of the range.

2.4.3 Combustion Characteristics. The combustion generated in the driver section creates a shock wave followed by a chemically reacting region. The dependent parameters that were used to characterize the combustion within the test section of the tube are the pressure rise across the shock,

the speed of the shock, and the speed of the chemically reacting radiation front. A secondary reaction was sometimes observed following the reflection of the incident shock wave from the end wall. The

incident (or forward-travelling) shock wave speed and pressure ratio were determined from the
HIGH SPEED FLAMES AND QUASI-DETONATIONS 21

3200

X

_l

<

<

<

LU

<

00

o

^ 2800

2500

<

I 2400

o

<

2200

2000

800 1200 1600 2000 2400

COMPUTER TIME COUNT, DT = 3us

Figure 7. Typical measurement signals representing flame radiation detected by the

photodiodes.
22 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

2600

1600 2000 2400 2800

COMPUTER TIME COUNT, DT = 3us

Figure 8. Typical measurement signals representing shock pressure detected by the dynamic

pressure transducers.
2. HIGH SPEED FLAMES AND QUASl-DETONATIONS 23

piezoelectric transducer signals, and the time between activation of the photodiodes was used to calculate the forward-travelling radiation (or flame) front. The repeatability of the measurements was affected by the following factors: preparation of the

mixtures; circulation/homogenization of the mixtures; opening of the gate valve; the ignition

parameters; formation/propagation of the flame/shock; vibrations of the spiral insert; and ambient

temperature changes (ambient air pressure and humidity changes did not affect the results as air was

supplied from a gas cylinder). Because each of these factors has an indeterminate randomness

associated with it, a single test condition was repeated eleven times to quantify the precision of the

experiment: a lean mixture of propane and air, with no suppressant, and the 10 m long tube with no

spiral in the test section. More than twenty replicates would have been required to produce a meaningful standard deviation as specified by Taylor and Kuyatt (1994); thus, in the present study the

maximum, rather than standard, deviation is used to indicate the precision of the inferred results. The mean of the eleven tests and maximum absolute deviations are as follows: flame speed, 334 m/s ± 38

m/s; shock speed, 681 m/s ± 25 m/s; and shock pressure ratio, 8.16 ± 0.38.

2.4.3.1 Ethene/Air Mixtures. Figure 9 shows the dependence of the forward shock wave

velocity versus equivalence ratio of the ethene/air mixture for the cases with and without the spiral insert in the 2.5 m test section. The equivalence ratio was changed to cover the full range of various

combustion/flammability modes detectable by the installed apparatus. The shock wave generated by

an accelerating flame is detectable for equivalence ratios between 0.5 and 2.12 for the two geometric

configurations. The maximum shock velocity of nearly 2000 m/s was recorded for a rich mixture with

the spiral absent. Except for the extreme lean and rich cases, with shock velocities less than 500 m/s,

the flame was intimate with the shock.

Figure 10 displays the respective forward shock pressure ratios in the ethene/air mixture versus equivalence ratio for the two geometric configurations. Interestingly, the maximum pressure ratio of

35 was recorded for the situation with the spiral, which indicates clearly that transverse shock

reflections from the wall play an important role in the whole process. However, in general, the shape

of the pressure ratio curves corresponds well with the shape of the velocity curves. There are four modes of combustion observed in the detonation/deflagration facility:

a. Low-speed deflagration : generates a weak pressure wave in which the flame front is uncoupled. A typical pressure wave velocity is 400 m/s and pressure ratio is 1.5 for the two

geometric configurations.

b. High-speed deflagration : generates a strong pressure wave that is coupled with the flame

front, the velocity of the flame is the same as the velocity of the pressure wave, and the

shock-flame distance is on the order of centimeters. A typical velocity is 800 m/s and

pressure ratio is 16 for the lean mixtures, and 1400 m/s and 20 for the rich mixtures, respectively, when the spiral insert is in the tube. When the spiral is not present in the tube the

respective parameters are as follows: 700 m/s and 7 for the lean mixtures, and 600 m/s and 3 for the rich mixtures.

c. Quasi-detonation : associated with the occurrence of high velocities and pressure ratios. The

flame front is coupled with the pressure wave, the velocity of the flame is the same as the

velocity of the pressure wave, and the shock-flame distance is on the order of millimeters.

Typical velocities are 1200 m/s to 1500 m/s and pressure ratios are 20 to 35 over broad lean

and rich ranges for the situation with the spiral insert in place. When the spiral is not present

the transition from the high-speed deflagration mode to a detonation is gradual. The velocities
24 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

2000

^ 1 800

> 1600

o

o

_i

UJ

>

<

o

o

X

(/) Q

<

o

Li_

<

o

1400

1200 -

1000

800 -

600

400

200

0

0.50 0.75 1.00 1.25 1.50 1.75 2.00

C2H4/AIR EQUIVALENCE RATIO

Figure 9. Combustion modes in the ethene/air mixtures - shock/flame velocities (2.5 m test section).
HIGH SPEED FLAMES AND QUASI-DETONATIONS 25

40

35 -

30

25

^ 20

o

Q 1 5

<

cr2 10

ALL SPIRALS IN

1 SPIRAL OUT

0

0.50 0.75 1.00 1.25 1.50 1.75 2.00

C2H4/AIR EQUIVALENCE RATIO

Figure 10. Combustion modes in the ethene/air mixtures - shock pressure ratios (2.5 m test section).
26 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

and pressure ratios are significantly lower. This mode occurs on the rich side of the ethene/air

mixture.

d. Chapman-Jouguet detonation : occurs only in the rich ethene/air mixture without the spiral

insert. The flame front is coupled with the pressure wave, the velocity of the flame is the same as the velocity of the pressure wave, and the shock-flame distance is undetectable. A

typical velocity is 1900 m/s and pressure ratio is 30. The velocity corresponds to the

theoretical equilibrium thermodynamic estimates; however, the pressure ratios are 1/3 higher,

which may indicate the occurrence of an over-driven detonation mode.

2.4.3.2 Propane/Air Mixtures. Uninhibited propane/air mixtures were evaluated in the 2.5 m

test section as well, with and without the spiral insert in the test section. Figure 11 shows the

dependence of the forward shock wave velocity on equivalence ratio. The shock wave generated by

the accelerating flame was detectable for equivalence ratios between 0.65 and 1.45, both with the

spiral present and absent. The maximum shock velocity of about 1300 m/s was recorded for the

stoichiometric case with the spiral in place, much less than was found in the ethene/air mixture. The

flame velocity was the same as the shock velocity for wave speeds above 800 m/s. This means that the propane generated reaction front can be more easily decoupled from the shock wave than in the

ethene/air mixture. Figure 12 displays the respective forward shock pressure ratios in the propane/air

mixture versus equivalence ratio. Here the maximum pressure ratio of 27 corresponds to the

maximum velocity for the situation with the spiral. In general, the shape of the pressure ratio curves corresponds closely to the shape of the velocity curves. Three of the four combustion modes observed with the ethene/air experiments were identified

when propane was the fuel: low-speed deflagration, high-speed deflagration, and quasi-detonation.

The Chapman-Jouguet condition was not observed, and when the spiral was not present in the test section, the quasi-detonation regime of combustion disappeared. There are other differences between

the two fuels, as well. The combustion modes at higher velocities overlap totally in the ethene/air

mixture for the cases with and without the spiral, while the propane/air mixture is characterized by a

clear separation between the combustion modes for the two arrangements. Also, the regime of

equivalence ratios for which combustion is detectable in the tube is much broader for the ethene/air

mixture. Furthermore, the detonation process is unable to develop in the propane/air mixture when

the spiral insert is missing from the tube. However, it is noteworthy that for the first time a quasidetonation in a propane/air mixture has been recorded in the presence of the spiral obstacle. This

finding extends the results of Lee (1984) and Peraldi et al. (1986).

2.4.4 Suppression Characteristics. The performance of the three extinguishing compounds (C2HF5,

CjFg, and CF3I) are analyzed by comparing the velocity and pressure ratio suppression characteristics

in the lean, stoichiometric, and rich ethene/air and propane/air mixtures, with and without the presence

of obstacles in the test section. (Note: the spiral inserts are always present in the driver section.) Pure

nitrogen was placed in the test section as a benchmark to compare its performance to the above

alternatives. The fully nitrogen-suppressed pressure ratios, shock speeds and combustion wave speeds

are compared to the totally uninhibited fuel/air mixtures in Table 1. In the ethene/air mixtures, as one might expect, the flame velocity is zero when the test section is filled with N2, but only under lean and stoichiometric conditions. In the rich mixture the

velocity is slightly higher than zero. This means that the residual flame from the driver section of the

tube enters the suppression section, since the flame velocity was measured just behind the gate valve

separating the two sections of the tube. The respective forward shock velocities decrease because of
2. HIGH SPEED FLAMES AND QUASI-DETONATIONS 27

1 400

C/)

O

o

_J

UJ

>

<

_J

CJ

o

o

<

CO

X

O

1 200

1000

800

^ 600

<

400

200

0

—

1

'

1

' 1

i

1

'

—

1

\

- " "

1

-

j/ •T

\ /

/

^ k

1

ALL SPIRALS

- - 1 SPIRAL C

IN

)UT

1 1 1,1.1. 1 0.6 0.7 0.8 0.9 1.0 1.1 1.2 1.3 1.4 1.5

C3H8/AIR EQUIVALENCE RATIO ' -

Figure IL Combustion modes in the propane/air mixtures - shock/flame velocities (2.5 m test section).
28 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

Figure 12. Combustion modes in the propane/air mixtures - shock pressure ratios (2.5 m test section).
HIGH SPEED FLAMES AND QUASI-DETONATIONS

4000 -

3500 -

^ 3000

2500 -

2000

4000 -

3500 -

^ 3000

2500 -

2000

5250 5500 5750 6000 6250

TIME. MICROSECONDS

6500

Figure 13. Pressure traces taken in CFjI-inhibited CgHg/air mixture (a) 2.04 m into test section,

and (b) 2.37 m into test section.
30 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

2800

I

2600

CO

LlI Q

O

o 2400

2200 -

2800

g 2600

(7) u

o

o 2400

\-

o

2200

(b)

- I I I I I I I I If I I I I I I I I I I I il I r I I ~ 5250 5500 5750 6000 6250 6500

TIME, MICROSECONDS

Figure 14. Photodiode signals from CFjI-inhibited C3Hg/air mixture, taken (a) 2.04 m into test section, and (b) 2.37 m into test section.
2. HIGH SPEED FLAMES AND QUASI-DETONATIONS 31

Table 1. Pressure ratios, shock speeds and combustion wave speeds comparing fully suppressed

(100% N2) to totally unsuppressed (0% N2) test conditions

0% Nitrogen 100% Nitrogen

Fuel Test Mixture Section Pi/Po

(m/s)

Vcomb

(m/s)

1 \J MlULK (m/s)

Vcomb

(m/s)

2.5 m 18 1170 1170 2.5 440 0

C2H4, (spiral) >K A '7 C 0=0.75

5.0 m

(no spiral)

6.2 710 600 4.1 600 0

2.5 m 26 1400 1400 3.5 450 0

C2H4, (spiral)

Q=l .00 5.0 m

(no spiral)

6.5 720 720 4.7 660 0

2.5 m 35 1530 1530 * * * C2H4, (spiral)

0=1.25

5.0 m

(no spiral)

11 940 930 5.0 730 20

2.5 m 11 815 815 * * *

(spiral)

0=0.00

5.0 m

(no spiral)

8.1 690 330 4.5 575 103

2.5 m 13.5 900 900 * * *

(spiral)

0=1.00

5.0 m

(no spiral)

8.8 695 620 4.7 619 105

2.5 m 16 890 890 * * * C3H8, (spiral)

0=1.25

5.0 m

(no spiral)

8.3 690 510 4.5 586 45

* no data available
32 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

the disappearance of the energy supplied by exothermic chemical reactions, and the forward shock

pressure ratios drop most significantly when the spiral is present. In the propane/air mixtures, in every case the flame velocity is higher than zero, which means

that 100 % nitrogen cannot fully extinguish the flame in the propane/air mixtures within a short

distance. When compared to the flame propagating in the ethene/air mixtures, the initial velocity in nitrogen is lower but it is more stable on contact with an inert environment. The respective forward

shock velocities in the propane/air mixtures are approximately proportional to the initial shock

velocities established in the driver section of the tube. The respective forward shock pressure ratios at the three equivalence ratios correspond well to the relative behavior of the shock velocities.

The interaction of the combustion wave with the inhibited fuel/air mixture is a dynamic process

which does not reach a steady state within the test section. Figure 13 shows two pressure traces taken

0.33 m apart in a CF3I inhibited propane/air mixture. The photodiode measurements taken at the same

two locations are shown in Figure 14. The initial and peak pressures can be seen to increase with the

distance that the wave travels, indicating that the CF3I is promoting the development of the shock

wave. The flame radiation, on the other hand, diminishes as the combustion wave travels from the

2.04 m to 2.37 m position. The on-off behavior of the photodiodes also suggests that the combustion

zone is thick and highly nonuniform in intensity.

The suppression data presented in this section have been measured at two different locations,

depending upon the configuration. The pressure transducers and photodiodes were mounted 2.04 m

and 2.37 m beyond the gate valve in the 2.5 m test section (with the spiral); in the 5.0 m test section

(without spiral) the photodiodes were moved to locations 0.21 m and 0.38 m after the gate valve,

while the pressure transducers remained at the 2.04 m and 2.37 m locations.

2.4.4.1 C2HF5 Performance. The primary motivation for conducting more research in the

detonation/deflagration tube is shown in Figure 15. The data for the lean C2H4/air mixture in the

2.5 m test section containing the spiral insert was collected in the earlier NIST study (Grosshandler

et ai, 1994). The shock pressure ratio reaches a maximum of 37:1 for a 6 % mixture of C2HF5. This

is more than double the pressure increase had no suppressant been added, clearly an untenable

situation were it to occur in a dry bay.

The tube geometry and fuel were altered to reduce the severity of the initial conditions to determine the impact on the combustion dynamics. The data points indicated by triangles in Figure 15 were taken with no spiral insert in the 5.0 m long test section. The initial shock pressure ratio is reduced by a factor of 3, and remains below 9:1 out to a partial pressure fraction of 10 %. The

sensitivity of the C2HF5/C2H4/air mixture to small perturbations in the detonation/deflagration tube

became apparent when the 6 % experiment was repeated and resulted in a detonation. Except for that one case, removing the spiral greatly reduced the severity of the combustion wave. Removing the spiral has a similar effect on the shock/combustion wave speed, plotted in Figure 16. The combustion process is completely extinguished when the C2HF5 partial pressure

reaches 10 %, as compared to about 15 % when the spiral is present in the test section.

Ethene is known to be highly reactive (which is one reason it was selected for the previous

study), but it is less representative of a vaporized jet fuel than propane. The original experiments were

repeated with CgHg instead of C2H4, and these are also plotted (filled circles) in Figures 15 and 16. The CjHg mixture produces significantly lower pressure ratios, never exceeding 15:1 when the spiral is present. The initial shock/combustion wave speed is cut in half by the change in fuel, and the flame

radiation is fully extinguished at partial pressure fractions about 4 % lower.

The equivalence ratio of the C2H4/air mixture was varied from lean to rich in the obstructed tube.

The pressure increase across the shock wave is plotted in Figure 1 7 as a function of the partial

pressure fraction of C2HF5 added to the test section. When the equivalence ratio is increased to 0=1.0
Figure 15. Shock pressure ratio in lean, C2HF5-inhibited hydrocarbon/air mixtures, showing

impact of spiral insert and fuel-type on suppression process.
34 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

600

on

>-o

o

_l

LU

>

LiJ >

<

o

I—<

<

1400

200

1000 -

800

600

400

200

2 4 6 8 10 12 14

C2HF5 PARTIAL PRESSURE FRACTION, %

Figure 16. Shock/combustion wave speed in lean, C2HF5-inhibited hydrocarbon/air mixtures

showing the impact of spiral insert and fuel-type on suppression behavior.
2. HIGH SPEED FLAMES AND QUASI-DETONAnONS 35

0 I ' \

I

, \ , \ > \ I I I I > I 0 2 4 6 8 10 12 14 16

C2HF5 PARTIAL PRESSURE FRACTION/ %

Figure 17. Shock pressure ratios in C2HF5/ethene/air mixtures (2.5 m, with spiral).
36 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

35

30

<

cr 25

0

9 V•

PHI = PHI --

= \.2[

= 1.0(D

PHI == 0.7(

0 2 4 6 8 10 12 14

C2HF5 PARTIAL PRESSURE FRACTION

1 6

Figure 18. Shock pressure ratios in C2HF5/ethene/air mixtures (5.0 m, no spiral).
2. fflGH SPEED FLAMES AND QUASI-DETONATIONS 37

Figure 19. Shock pressure ratios in C2HFg/ethene/air mixtures (2.5 m, spiral).
38 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

O

CL_o

<

crLJ

ct: Z)

LD

Ul

m

CL

o

o

X

00

Q

<

cro

<

00

X

CJ

7

6

5

4

3

^ 2

-

0

V PHI = 1.25 • PHI = 1.00

o PHI = 0.86

0 2 4 6 8 10

C2HF5 PARTIAL PRESSURE FRACTION/ %

Figure 20. Shock pressure ratios in C2HF5/propane/air mixtures (5 m, no spiral).
2. HIGH SPEED FLAMES AND QUASI-DETONATIONS 39

00

>-o

o

o

o

X

(/) Q

<

o

<

X

O

1800

1 600 -

1400

LjJ

> 1200

LJ

<

^ 1000

800 -

600

^ 400 -

200 -

0

V PHI = 1 • PHI = 1

.25

.00

o PH I = 0 .75

0 2 4 6 8 10 12 14 16

C2HF5 PARTIAL PRESSURE FRACTION/ %

Figure 21. Shock/combustion wave speed in C2HF5/C2H4/air mixtures (2.5 m, spiral). 1}
40 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

>-O

o

_J

UJ

>

UJ

<

Q

<

oro

u.<

X

CM

U

600

1400 -

1200 -

1000 -

800 -

600

400

200

2 4 6 8 10 12 14 16

C2HF5 PARTIAL PRESSURE FRACTION/ %

Figure 22. Combustion wave speed in C2HFg/C2H4/air mixtures (5 m, no spiral).
HIGH SPEED FLAMES AND QUASI-DETONATIONS 41

if)

o

o

-J

LU

>

o

o

X

Ul

Q

<

<:O

Li_

<

OO

X

o

1000

900

800

700

600

500

400

300

200

100

0

V PHI = 1.25 • PHI =1.00

o PHI = 0.86

0 2 4 6 8 1 0

C2HF5 PARTIAL PRESSURE FRACTION

Figure 23. Shock/combustion wave speed in C2HF5/C3Hg/air mixtures (2.5 m, spiral).
42 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

400

if)

>-o

o

UJ

>

<

cr<

o

<

OO

X

O

200

000

800

600

400 -

200 -

V PHI = 1. 25

• PHI = 1. 00

o PHI = 0. 86

0 2 4 6 8 10

C2HF5 PARTIAL PRESSURE FRACTION/ %

Figure 24. Combustion wave speed in C2HF5/C3Hg/air mixtures (5 m, no spiral).
2. HIGH SPEED FLAMES AND QUASI-DETONATIONS 43

and 0=1.25, the pressure rise in the uninhibited (partial pressure fraction = 0) test section increases

substantially. However, the suppressant does not significantly worsen the situation for the stoichiometric mixture, and it is much more effective for the rich condition.

When the shorter, obstructed test section is replaced with the 5.0 m long unobstructed test section,

the strength of the shock wave is decreased by a factor of about 3, as can be seen in Figure 18. For

the rich and stoichiometric mixtures, adding as little as 2 % C2HF5 reduces the severity of the pressure

build up. Additional agent has little impact on the pressure ratio. The behavior of the agent under

lean conditions is less deleterious in the test section without the spiral in place. However, a small

enhancement in pressure can be seen at a partial pressure fraction of 5 %. A single test at 6 %

exhibited extreme pressures, although that data point could not be repeated.

Replacing ethene with propane cuts in half the pressures generated by the shock passing through

the 2.5 m test section with the spiral insert. Figure 19 is a plot of Pj/Pq as a function of the amount

of HFC- 125 added and the fuel/air equivalence ratio. The stoichiometric condition is the most

dangerous with propane (compare to Figure 17), but for all equivalence ratios, the propane shock wave

is easier to attenuate than the ethene-generated shock. In the long test section without the spiral insert,

neither stoichiometry nor the amount of C2HF5 have much effect on the shock pressure ratio (see

Figure 20).

The ethene/air mixture produces a quasi-detonation when the spiral is present in the 2.5 m test section. Thus, the shock and combustion wave travel together at the same speed. Figure 19 shows the wave speeds measured for this geometry as a function of stoichiometry and the amount of HFC- 125 in the test section. Notice that the wave speed for the lean mixture decreases about monotonically with

C2HF5 partial pressure, which is in contrast to the more complex behavior exhibited by the pressure

ratio shown in Figure 17. When there is no spiral insert in the test section, the flame becomes

decoupled from the shock. The flame velocities for different C2HF5/C2H4/air mixtures are plotted in Figure 22. The flame speed is much slower than the shock speed and goes to zero when the HFC- 125

reaches a partial pressure fraction of 10 % for O = 0.75 and 1.0. A 16 % partial pressure fraction is required when the fuel/air ratio is rich. The outlying point for the lean mixture with 6 % agent has the

correspondingly high pressure ratio shown in Figure 18. The rather chaotic looking behavior in Figure 22 can be partially attributed to the location of the photodiodes (0.3 m beyond the gate valve),

where the gradient in agent concentration is high and the wake created by the gate valve may be

influencing the flow field. The shock and combustion wave travel together when propane is the fuel and the spiral insert is

in the 2.5 m test section. The behavior of the wave speed (see Figure 23) is qualitatively similar to the ethene system, although the maximum speeds are lower and the amount of C2HF5 required to quench the combustion (i.e., speed < 400 m/s) is a couple of percent lower with propane. Figure 24

shows the combustion wave speed without a spiral insert. The stoichiometric mixture produces the

fastest flame, and is quenched when the C2HF5 partial pressure fraction is 10 %. Comparing this plot

to Figure 22, it is interesting to note that the rich propane flame can be fully suppressed at a concentration less than half the value required to suppress rich ethene combustion.

2.4.4.2 C3Fg Performance. The perfluoropropane (FC-218) does not produce the high overpressures found when C2HF5 is added to the lean ethene mixture in the presence of the spiral insert. Figure 25 shows the shock pressure ratio measured in the obstructed configuration with the C3Fg/

C2H4/air mixture. There is a slight enhancement when the partial pressure fraction is 2 %, but

otherwise the drop-off in Pi/Pg with agent concentration is well-behaved. The lean mixture is the

easiest and the rich mixture the most difficult to suppress with C3Fg, which is opposite to the results observed when C2HF5 was the agent.
44 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

0 2 4 6 8 10

C3F8 PARTIAL PRESSURE FRACTION/ %

Figure 25. Shock pressure ratio in C3Fg/C2H4/air mixtures (2.5 m, no spiral).
HIGH SPEED FLAMES AND QUASI-DETONATIONS 45

O_

CL

o

<

UJ

cr

Z)

CO

(/) LU

cr

CL

o

o

X

CO

Q

<

o

Lj_

<

CNO

16 18

C3F8 PARTIAL PRESSURE FRACTION

Figure 26. Shock pressure ratio in C3Fg/C2H4/air mixtures (5 m, no spiral).
46 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

Figure 27. Shock pressure ratio in C3Fg/C3Hg/air mixtures (2.5 m, spiral).
fflGH SPEED FLAMES AND QUASI-DETONATIONS 47

O

Q_

O

I—

<

LxJ

or:

in

LjJ cr

CJ

o

X

tn

Q

<

o

<

00

o

6

4

3

^ 2

0

0

V PHI

• PHI

o PHI =

1.25

1.00

0.86

2 4 6 8 0

C3F8 PARTIAL PRESSURE FRACTION

Figure 28. Shock pressure ratio in C3F8/C3Hg/air mixtures (5 m, no spiral).
48 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

00

>-o

o

_J

Ld

>

LJ

<

o

oX

Q

cr<

O

<

X

Csl O

1800

1 600

1400<^

1200

1000

800

600

^ 400

200 -

0

V PH]

• PH]

[ = 1.25

[ = 1.00

o PH][ - 0.75

0 2 4 6 8 10

C3F8 PARTIAL PRESSURE FRACTION/ %

\ V

•\

Figure 29. Shock/combustion wave speed in C3Fg/C2H4/air mixtures (2.5 m, spiral).
2. HIGH SPEED FLAMES AND QUASI-DETONATIONS 49

C3F8 PARTIAL PRESSURE FRACTION/ %

Figure 30. Combustion wave speed in C3Fg/C2H4/air mixtures (5 m, no spiral).
50 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

00

o

o

_J

LU

>

o

oX

CO

Q

<

o

<

CX) X

o

1000

900

800

700

600

500

400

300

200

100

0

V PHI = 1.25 • PHI = 1.00

O PHI = 0.86

0 12 3 4 5 6 7

C3F8 PARTIAL PRESSURE FRACTION/ %

Figure 31. Shock/combustion wave speed in C3Fg/C3Hg/air mixtures (2.5 m, spiral).
2. fflGH SPEED FLAMES AND QUASI-DETONATIONS 51

C/1

>-o

o

800

700

600

^ 500

400

LJ

<

_l

Q

cr<

^ 300

o

<

00

X

200

1 00

0

V PHI = 1.25 • PHI = 1.00

o PHI = 0.86

0 2 4 6 8 10

C3F8 PARTIAL PRESSURE FRACTION/ %

Figure 32. Combustion wave speed in CgFg/CgHg/air mixtures (5 m, no spiral).
52 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

By removing the spiral, the pressure rise across the shock is greatly reduced, and the suppression

behavior of CjFg is very similar to C2HF5, as can be seen by comparing Figure 26 to Figure 18. The

pressure spike in the lean mixture containing CjFg cannot be easily explained, but may have

something to do with the sensitivity of the ethene/air reaction to small perturbations. Note that a spike

is also present in Figure 18, but that is hypothesized to be related to the hydrogen in the HFC-125 and

its impact on the chain branching reactions.

The FC-218 does a much better job than HFC-125 in reducing the pressure buildup in the

obstructed propane/air experiments. Figure 27 shows that a 4 % partial pressure fraction of C3Fg

attenuates the shock wave traveling through the rich mixture as effectively as 2.5 times as much

C2HF5. Suppression of the lean propane mixture requires 6 % CjFg, compared to 8 % C2HF5.

Figure 28 compares the shock pressure ratios for the lean, stoichiometric and rich propane/air mixtures

in the 5 m long test section without the spiral insert. The FC-218 has little impact on the strength of

the shock wave under these conditions. The same statement applies to HFC-125 (see Figure 20).

The combined shock and combustion wave speeds are plotted in Figure 29 for the C3Fgsuppressed C2H4/air mixture in the obstructed 2.5 m long test section. The speed decreases with the

addition of FC-218 more quickly, dropping below 400 m/s at a partial pressure fraction of 10 %

(compared to 15 % when HFC-125 is the agent). Removing the insert results in the combustion wave

speeds shown in Figure 30. The very high speed for the lean mixture with 4 % C3Fg corresponds to a Chapman-Jouguet detonation, a result that is difficult to have predicted, especially since full suppression is attained when only 8 % C3Fg is added to the mixture. The rich flame travels more

slowly at low agent concentrations, but is not fully extinguished until the partial pressure fraction is 18 %. The speed of the propane combustion waves are plotted in Figures 31 and 32, with and without

the spiral insert, respectively. The FC-218 out-performs the HFC-125 under both arrangements

(compare to Figures 23 and 24). Under the worst conditions, only 8 % C^Pg is necessary to annihilate

the propane flame.

2.4.4.3 CF3I Performance. The suppression behavior of CF3I is more complicated than either

C2HF5 or C3Fg. This complexity is demonstrated in Figure 33, which shows that small concentrations

of CF3I dramatically reduce the shock pressure ratio generated in the obstructed 2.5 m long test section filled with C2H4/air mixtures. However, the pressures rise as the agent partial pressure fraction

is increased to 6 %, equalling or exceeding the totally uninhibited values. For the most reactive

mixture (O = 1.25), it takes over 13 % CF3I to reduce the pressure ratio to below 5.1, as compared to about 12 % for C2HF5 and less than 10 % for C3Fg. The unpredictable behavior extends to the test configuration with no spiral insert (see Figure 34). In these tests, it is the richest condition that is the

best behaved. For an equivalence ratio of 0.75, the pressure increases to a factor greater than 36:1

when the partial pressure fraction of CF3I is 4 %. This peak shifts to 8 % when the C2H4/air mixture

is stoichiometric. Similar peaks were observed when HFC-125 and FC-218 were used (compare to Figures 18 and 26), but only for O = 0.75. The precision of the CF3I suppression measurements were

checked by repeating four times the 8 % CF3I in stoichiometric ethene/air mixture. The maximum

absolute deviations of flame and shock velocities and pressure ratio were 157 m/s, 15 m/s and 0.97

respectively. This level of precision provides confidence that the trends observed in all the experimental sequences are real and meaningful.

The shock pressure ratio generated in obstructed lean, stoichiometric and rich CF3l/C3Hg/air

mixtures is shown in Figure 35. The behavior is similar to that observed with C2H4, except that the

stoichiometric mixture requires much more agent to get under control. With the spiral removed

(Figure 36), the pressure ratio drops down more quickly when propane is the fuel, but the stoichiometric condition persists.
HIGH SPEED FLAMES AND QUASI-DETONATIONS 53

0 I >

I

.

I

^ \ , \ , \ , I > L_

0 2 4 6 8 10 12 14

CF3I PARTIAL PRESSURE ERACTION/ %

Figure 33. Shock pressure ratio in CF3l/C2H4/air mixtures (2.5 m, no spiral).
54 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

V PHI = 1.25

CN

O

0 I \ \

\ I \ I \ \ ^ L

0 2 4 6 8 1012141618 20

CF3I PARTIAL PRESSURE FRACTION/ %

Figure 34. Shock pressure ratio in CF3l/C2H4/air mixtures (5 m, no spiral).
2. HIGH SPEED FLAMES AND QUASI-DETONATIONS 55

20

O

18

Q_

I—

<

UJcr

ZD

if) 00

UJ

Csl CL

^

O

O

X

00

Q

<

o

Li_

<

CX) X

O

1 6

14

12

10

8

^ 4

2

0

1

-

' 1

' 1

7 V

7 1

\ i\ I

V

•

PHI

PHI

= 1.25 = 1.00

o PHI = 0.86

1

1 0 2 4 6 8 10 12 1416

CF3I PARTIAL PRESSURE FRACTION/ %

Figure 35. Shock pressure ratio in CFgl/CjHg/air mixtures (2.5 m, spiral).
56 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS
fflGH SPEED FLAMES AXD QUASI-DETONAnONS 57

1500

1 600

^ 1400

> 1200

Q

<

o

!_!_ <

200

C

V PHI == 1.25 • PHI -- LOO

O ^ PHI == 0.75

0 2 4 6 8 10 12 14

CF31 PARTIAL PRESSURE FRACTION/ %

Figure 37. Shock/combustion wave speed in CF3FC2H4/air mixtures (2.5 m, spiral).
58 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

0 2 4 6 8 1012141618 20

CF3I PARTIAL PRESSURE FRACTION/ %

Figure 38. Combustion wave speed in CF3l/C2H4/air mixtures (5 m, no spiral).
2. HIGH SPEED FLAMES AND QUASI-DETONATIONS 59

CO

1^

O

o

_i

LU

>

o

X

CO

Q

<

or: o

<

00

X

O

1000

900

800

700

600

500

400

300

200

100

0

0 2 4 6 8 10 121416

CF3I PARTIAL PRESSURE ERACTION/ %

Figure 39. Shock/combustion wave speed in CFjI/CjHg/air mixtures (2.5 m, spiral).
60 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

0 2 4 6 8 10 12 14

CF3I PARTIAL PRESSURE FRACTION/ %

Figure 40. Combustion wave speed in CF3l/C3Hg/air mixtures (5 m, no spiral).
2. HIGH SPEED FLAMES AND QUASI-DETONATIONS 61

The wave speeds are congruent with the pressure ratio results. In Figure 37, the shock and

combustion waves are merged. The local minimum in wave speed occurs at a partial pressure fraction

of 4 %, which is slightly higher than the location of the minimum pressure ratio. The behavior of the

combustion wave can be seen most clearly in Figure 38. Acceleration to a Chapman-Jouguet

detonation is clearly observed for both the lean and stoichiometric C2H4/air mixtures. The rich

mixture requires 20 % CF3I to quench the flame radiation, compared to 16 % and 18 % for the

HFC- 125 and FC-218, respectively.

Figures 39 and 40 show the shock and combustion wave speeds when CjHg is the fuel. The

stoichiometric mixture requires the most CF3I to decelerate, but much less agent is required as compared to C2H4 combustion. Similar to C2^^5 ^3^8' ^'^ detonations are produced using

CF3I when propane is the fuel.

2.4.4.4 Relative Performance of C2HF5, C3F8, and CF3I. Relative performance of the three

compounds is compared as a function of the fuel and test section configuration in Figures 41 to 44.

The stoichiometric mixture is chosen as representative. When ethene is the fuel and the spiral insert is in place, the uninhibited shock/combustion wave speed is 1400 m/s. Figure 41 shows that the CF3I is the most efficient at reducing the wave speed for partial pressure fractions of 4 % and below. The

FC-218 (C3F8) reduces the wave speed most effectively at higher concentrations, attaining a fullysuppressed condition of 400 m/s at less than 10 %. Removing the spiral reduces the initial condition

to a high-speed deflagration traveling at 700 m/s (see Figure 42). The C2HF5 quenches the radiation

completely with a partial pressure fraction of 10 %. The C3Fg is less efficient in suppressing a flame

without the spiral insert, and actually enhances the flame when the partial pressure fraction is 8 %.

The CF3I causes a transition to a detonation at the same partial pressure fraction, producing a combustion wave that travels over 1400 m/s. A partial pressure fraction of 18 % is required for the

CF3I to fully suppress the flame.

The CF3I reduces the shock wave velocity the best of the three agents for the propane/air mixture

as long as the levels do not exceed 4 %, as seen in Figure 43. Beyond 7 %, the CF3I is the poorest

agent, requiring more than 16 % by partial pressure to attenuate the shock wave to less than 400 m/s.

By contrast, C3Fg attains full suppression when the partial pressure fraction is only 6 %. HFC-125 is in between. The data plotted in Figure 44 are taken with stoichiometric C3H8/air mixtures in the 5 m

test section without the spiral. FC-218 causes the combustion wave speed to decrease in a monotonic

manner, with suppression occurring when the partial pressure fraction is 8 %. Full suppression is attained with HFC-125 at a concentration of 10 %; however, 2 % and 6 % levels of C2HF5 strongly

enhance the exothermic reaction. The CF3I is relatively well behaved, but requires the largest amount

(on both a molar and mass basis) of the three agents to fully quench the radiation.

2.4.5 Uncertainty Analysis. The measurement results are subject to experimental uncertainties.

Because no generally applicable methods exist which would provide a measure of the reliability of the

experimental data with absolute certainty, only an estimate of the magnitude of these uncertainties is undertaken. The estimate is affected by the uncertainty of the determination of the independent

variables, such as partial pressures of the components constituting the mixtures under investigation,

and the dependent variables, such as shock wave amplitudes, and shock/flame time differences serving

to determine their velocities. The uncertainty of the determination of the partial pressure of an agent is affected by the

uncertainty of the static pressure transducer which is ± 0.15 % (combined non-linearity, hysteresis,

repeatability, and temperature effects); the combined uncertainty of the digital display device which is ± 0.015 %; the uncertainty associated with the purity of the gases which is ± 0.5 %; and the

uncertainty associated with the gas component losses due to leaks occurring in the circulation pump
62 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

600

1 400

o 1 200

o

1 000

800 -

600

400 -

200 -

0

0

o

•

C^HI

C3F8

? 5

V CF3I

2 4 6 8 10 12 14

AGENT PARTIAL PRESSURE ERACTION/

1 6

Figure 41. Shock/combustion wave speed in stoichiometric C2H4/air mixtures (2.5 m, spiral).
2. HIGH SPEED FLAMES AND QUASI-DETONATIONS 5

00

o

o

_l

LJ

>

1_U >

<

o

hcoZD

DO

:^

o

o

1800

1 600

1 400

> 1200

1000

800

600

400

200

0

o

•V 3I

V

CgH/AIR

iSTOICH.

0 2 4 6 8 10 12 14 16 18

AGENT PARTIAL PRESSURE FRACTION, %

Figure 42. Combustion wave speed in stoichiometric C2H4/air mixture (5 m, no spiral).
64 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

00

>-o

o

_l

LJ

>

LJ

<

O

o

X

m

1000

900

800

700

600

500

400

300

200

100

0

CgHg/AIR

STOICHIOMETRIC

o

V CI

0 2 4 6 8 10 12 14 16

AGENT PARTIAL PRESSURE FRACTION, %

Figure 43. Shock/combustion wave speed in stoichiometric CjHg/air mixtures (2.5 m, spiral).
HIGH SPEED FLAMES AND QUASI-DETONATIONS 65

1 400

c/)

>-o

o

_l

LjJ

>

<

_l

Li_

Q

<

o

<

00

X

o

1200

1000 -

800 -

600

^ 400 -

200 -

AGENT PARTIAL PRESSURE ERACTION

Figure 44. Combustion wave speed in stoichiometric C3Hg/air mixture (5 m, no spiral).
66 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

under operation which is ± 0.3 % (the worst case, assuming that only one component leaks).

Combining the uncertainty components by root-sum-of-squares (Taylor and Kuyatt, 1994), the resultant expanded uncertainty of establishing the composition of the mixture is 1.2 % of the reported value.

The determination of the shock wave amplitude is affected by the combined uncertainty of the

dynamic pressure transducer, which is ± 1.0 %; the combined uncertainty of the transducer amplifier,

which is ± 0.5 %; the combined uncertainty of the digital data acquisition system, which is ± 0.5 %;

and the combined uncertainty of the digital readout device, which is ± 0.2 %. The root-sum-ofsquares of these uncertainties results in an expanded uncertainty in determining the shock wave

amplitude of 2.5 % of the reported value. The uncertainty in the determination of the shock time

differences is affected by the some of the same elements and, after considering the differential nature of the measurement and the uncertainty of the readout device (± 0.1 %), can be estimated as ± 3.6 %.

The uncertainty of the determination of the flame time differences is affected by the following

factors: the combined uncertainty of the photodiode which is ± 1.0 %; the combined uncertainty of the

photodiode amplifier which is ± 0.5 %; the combined uncertainty of the digital data acquisition system

which is ± 0.5 %; the combined uncertainty of the digital readout device which is ± 0.1 %. The

resultant expanded relative uncertainty in the differential measurement of flame velocity is ± 3.5 %.

The precision of the experimental measurements can readily be determined by observing the

agreement between the numerical values obtained in a series of measurements performed in exactly the same way and under the same conditions. Depending on the interpretation of experimental results,

either the mean value and the standard deviation is of importance, or the mean value and the

maximum absolute deviation is of interest. In the present study it seems that the latter is more

meaningful because it indicates the worst possible case that could occur in reality. However, here,

both the standard deviation and the maximum absolute deviation are presented as an illustration. The

precision of the measurements in these experiments is affected by the following factors: preparation of

the compositions of the mixtures; circulation/homogenization of the mixtures; opening of the gate

valve; ignition event; formation/propagation of the flame/shock; vibrations of the spiral insert; and

ambient temperature changes (± 2 °C). Ambient air pressure and humidity changes do not affect the

results as air is supplied from a gas cylinder.

As an example experimental point, the 8 % of CF3I in the stoichiometric 6.5 percent ethene/air

mixture without the spiral insert is analyzed in more detail. For this case, four experiments are taken

into account (it is necessary to treat the analysis as very much approximate as from the statistical point

of view it would be required to include more experimental points). The means of flame/shock velocities and pressure ratio are 1484, 1635 m/s and 32.5, respectively; the maximum absolute deviations of

flame/shock velocities and pressure ratio are 157, 15.2 m/s and 0.97 respectively; the maximum

relative deviations of flame/shock velocities and pressure ratio are 10.6, 0.9 and 3.0 % respectively;

the population standard deviations of flame/shock velocities and pressure ratio are 101, 8.7 m/s and

0.61 respectively; and the sample standard deviations (as the number of experiments here is less than

20) of flame/shock velocities and pressure ratio are 117, 10.1 m/s and 0.70 respectively (the expanded

uncertainties are 233, 20.2 m/s and 1.4). This gives an idea on the precision of the experimental

results obtained with the detonation/deflagration tube facility.

2.5 Summary of Results

The fire threat discussed in this section results from a premixed, uninhibited turbulent flame accelerating into a quiescent, premixed, inhibited fuel/air mixture. The response of this highly dynamic

situation to different suppressants is sensitive to the fuel, stoichiometry and geometry in the test
2. HIGH SPEED FLAMES AND QUASI-DETONATIONS 67

section, as well as to the amount of agent used. A successful agent mixture would quench the

combustion wave, as evidenced by the disappearance of visible radiation, and diminish the speed and

pressure build up across the incident shock wave. However, even with 100 % nitrogen in the test section, a residual shock wave will persist. Short of complete suppression, an agent which reduces the

combustion and shock wave speeds and minimizes the pressure increase is desired.

Twelve alternative agents to halon 1301 for protection of aircraft dry bays were ranked previously

according to how well each could suppress a laboratory turbulent spray flame and a quasi-detonation

(Grosshandler et al, 1994). The experiments were designed to cover the range of conditions that might occur following the penetration by an incendiary device of a fuel cell adjacent to a dry bay.

High over-pressures (37:1) were measured when HFC-125 was used to suppress a lean C2H4/air quasidetonation. The range of initial conditions that lead to a worsening of the situation rather than a

lessening of the threat has been investigated in the current study.

The detonation/deflagration tube facility was modified to operate over a less severe range of

conditions. Pressure ratios below 9:1 were generated routinely for lean, stoichiometric and rich

mixtures. These lower pressures were achieved by removing the spiral insert in the test section and by

replacing the more reactive ethene with propane, which is also a better simulant of vaporized jet fuels. The flame speed was monitored close to the entrance of the test section to better assess the immediate

impact of the suppressant on the flame. Previously, incident shock speeds over 1500 m/s were

recorded. The current experiments with propane as fuel yielded uninhibited flame speeds between 300

m/s and 600 m/s, much closer to the hundreds of meters per second estimated to occur in the full-scale dry bay experiments. A further modification to the facility has been the doubling of the test section

length, to 5 m, which has increased the time available for the incident shock to reflect back into the

turbulent flame front. This arrangement has allowed the incident shock speed and pressure ratio, the

turbulent flame speed, and the conditions behind the reflected shock wave all to be monitored. The

reflected shock wave was always found to be stronger than the incident wave, and, with no agent

present in the test section, led to a detonation for a range of initial stoichiometrics. Thus, with a

single shot, we were able to observe the performance of the suppressant under moderate and highly

dynamic conditions.

Table 2 summarizes the results of all the detonation/deflagration experiments done with the three

agents in this and the earlier NIST study. The suppression conditions are defined as the partial

pressure of agent in the test section necessary to either totally quench the radiation from the reactants

or to reduce the pressure ratio to the value had 100 % nitrogen been used. The peak pressure ratios and reaction wave speeds refer to the maximum in the plots of pressure ratios (or velocities) versus agent partial pressure fractions. The agent percent is the partial pressure fraction where the maximum

is reached. In most cases, small amounts of agent increased the pressure and reaction wave velocity.

A value of 0 % implies that the maximum is attained solely at the uninhibited condition. Generally

speaking, the ethene quasi-detonation requires considerably more agent to extinguish than the turbulent

propane flame; the stoichiometric mixtures require more agent than either rich or lean conditions; C3Fg

(FC-218) requires the lowest partial pressure fraction to totally suppress both quasi-detonations and

turbulent flames; C2HF5 (HFC-125) is the least effective suppressant of a quasi-detonation; and CF3I is the least effective compound for total suppression of stoichiometric and rich turbulent propane flames.

The highest pressure ratio observed was for the lean ethene quasi-detonation with 6 % C2HFg added.

HFC-125, when added to the stoichiometric turbulent propane flame at a partial pressure fraction of

2 %, greatly accelerated the speed of the reaction wave, but did little to enhance the pressure build up. The exact conditions that are likely to exist in a dry bay prior to a fire or explosion are impossible to control. Unfortunately, the relative behavior of the three agents under investigation is strongly dependent upon the initial conditions, causing one chemical to be clearly superior under one

arrangement and the same chemical to perform poorly in another. There are some general statements
68 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

Table 2. Summary of Experimental Results in Detonation/Deflagration Tube

Fuel and Equivalence Ratio

Parameter Agent Ethene" 0 = 0.75 QuasiUCLUilaUUn

Ethene" 0 = 1 QuasiucLuriaiiunEthene"

<D = 1.25 QuasiUClUllallUll

Propane*" 0 = 0.86 Turbulent Flame

Propane*"

<1) = 1.0 Turbulent Flame

Propane*" 0 = 1.25 Turbulent Flame

Maximum

Pressure

none

N2

18 (0%)

2.5 (100%)

26 (0%)

3.5 (100%)

35 (0%)

8

8.1 (0%)

4.5 (100%)

8.8 (0%)

4.6 (100%)

8.3 (0%)

4.5 (100%)

Ratio'^ (@ partial pressure %f

C2HF5 J 1 (.0/0; Ly \\j /O

)

J J \\} /c) O.J /O) 0.0 \£./0) O.J \\J/0)

CiFo i 1 (Z, /o) y.D \l/o) 0.5 {2%)

CF3I ^ i \yj /o

)

J J \\) lO ) 0.1 \\J /O

)

0.0 \\J/C) 0,0 \\J lO)

Maximum

Reaction

none

N2

1170 (0%)

0 (100%)

1400 (0%)

0 (100%)

1530 (0%)

s

330 (0%)

100 (100%)

620 (0%)

100 (100%)

510 (0%)

50 (100%)

Wave

Speed*, m/s

(@ partial pressure %)''

C.HFc

C3F8

CF3I

1170 (0%)

1250 (2%)

1170 (0%)

1410 (3%)

1400 (2%)

1400 (0%)

1530 (0%)

1530 (0%)

1530 (0%)

510 (2%)

460 (2%)

450 (1%)

1180 (2%)

620 (0%)

740 (1%)

510 (0%)

510 (0%)

590 (1%)

N2 40% s g g g g

Suppression

Partial C2HF5 13 to 15% 13 to 15% 13 to 15% 7.5 to 8% 9 to 10% 5 to 6%

Pressure Percent*^ C3F8

CF3I

8 to 10%

> 10%

> 10%

> 12%

> 10%

13 to 14%

5 to 6%

5.5 to 6%

7 to 8%

13 to 14%

3 to 4%

7 to 8%

^ 2.5 m test section, with spiral insert, measurement location 2.2 m into test section

*" 5.0 m test section, without spiral insert, measurement location 0.3 m into test section ± 5% of value relative uncertainty ± 1% absolute uncertainty, and note that 0% implies no enhancement over zero inhibitor conditions ± 11% of value relative uncertainty

' ± 1% absolute uncertainty, based upon no flame radiation or pressure ratio equal to value attained by 100% N;

^ no data available
2. HIGH SPEED FLAMES AND QUASI-DETONATIONS 69

that can be made, though. For example, obstacles in the test section lead to higher shock pressure

ratios and initial velocities; ethene/air mixtures lead to higher shock pressure ratios and initial velocities; a residual shock wave remains even when the combustion wave is extinguished; and the

speed of the combustion wave without obstacles in the flow responds to the agents in a more chaotic manner than the shock pressure ratio. The impact of these statements are reflected in Figures 45, 46 and 47 for HFC-125, FC-218 and

CF3I, respectively, in an attempt to identify the best overall chemical for an uncertain application. The

response parameter, 4^, is defined as

= ^

~

where Xg is the value of the parameter of interest (combustion wave speed, shock speed, or shock

pressure ratio) when no agent is present and x* is the corresponding value when extinction has

occurred. A value of unity for m means that the agent has no beneficial impact on the combustion

process; ^ > 1 implies the agent exacerbates the situation; a performance parameter near zero is desirable, indicating close to total suppression. The volume percents plotted on the abscissa in Figures 45 through 47 are identical to the partial pressure fractions for ideal gas mixtures. This is a

reasonable approximation for the close-to-ideal mixtures under investigation in this work. The volume

percent of agent is also a more conventional way to compare the performance of different compounds.

Each point in Figures 45 through 47 represents one measured parameter for an individual

experiment. For most of the experiments, the normalized radiation wave speed (determined from the

photodiodes) and the normalized shock speed and pressure ratio (measured with the piezoelectric

transducers) are plotted. The averages of ^ are shown as the dotted lines in Figures 45 through 47.

The individual values exceed 2.5 (the maximum plotted) for a few of the experiments using C2HF5

and CF3I. These data are not shown in the Figures but were included when the averages of ^ were

computed. By comparing the three curves one gets an indication of general trends, and where one

agent is likely to out-perform another.

Examining Figure 45, one sees that the HFC-125, on average, cuts the magnitude of the

deflagration/detonation threat in half when its volume fraction is greater than about 8 %. FC-218

(Figure 46) achieves the same level of protection for volume fractions greater than 5 %, while a volume fraction of almost 9.5 % is required of CF3I to reduce the combustion activity to half (Figure

47). All three agents increase the threat for lesser levels. The HFC-125 leads to pressures and wave

speeds higher by a factor of 1.8 at volume fractions near 5 %. The FC-218 and CF3I produce close to a 50 % overshoot when the volume fractions are 2 % and 5 %, respectively. The CF3I is much more

chemically reactive than the other agents, undergoing three transitions between suppression and

enhancement as its concentration is increased. A 90 % reduction in the threat requires 14 % CF3I,

13 % FC-218, and 11 % HFC-125 by volume. Total extinction of the exothermic reaction under all conditions examined in this study requires volume percentages greater than 20 %, 18 % and 16 %,

respectively, of CF3I, FC-218 and HFC-125.

As a benchmark, about 6 % (vol.) of CF3Br was required to reduce the pressure build-up by 90 %

in the earlier NIST study (Grosshandler et ai, 1994). This implies that almost three and a half times

as much liquid FC-218 is necessary for equivalent protection, while about two and one half times

liquid CF3I must be stored. With this measure of performance the HFC-125 appears to require the

least liquid storage of the three halon alternatives tested, requiring just over twice the liquid volume of

CF3Br.
70 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

Ld

I—

Ld

<

<

_

LU

O

o

Li_

LJ

D_

Q

LUM

o

2.5

2.0

1 .5

- o

1 .0

0.5

0.0

0

o

o

o \i

o

o

o

o

o

o

o

o

o

o

o

o

o

o

o

o o

o R

o

o

o

o

o

8

O c^. o

o o

o o

o

o

o

o

-e—eo

o

o

o

o

o

o

o

HFC-125

o

—e-

.Q.^ o

o —e- 10 15

C2HF5 VOLUME %

20

Figure 45. Performance parameter vs. volume fraction of HFC-125 (C2HF5).
HIGH SPEED FLAMES AND QUASI-DETONATIONS 71

2.5

2.0

LU

o 1 .5

o

Li_

^ 1.0^

0.5

0.0

I I I I

o

o

o

o

o

o o

; O

8

O

O

o

o

o

0 \

''f o

o •. o

O 0

o

"

o

Q

o

o

o

o

o

o

-e- -2—1—§—I—©—I—e—

1 I I

o

o

o

FC-218

8 8

-e- o V--o.-e—' 0

0 10

C3F8 VOLUME %

15

o

20

Figure 46. Performance parameter vs. volume fraction of FC-218 (CjFg).
2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

2.5

2.0

1 .5

o

O;

1.0

T 1 1 r T 1 r 1 1 1 r

o

o

O '.oo

o

o

o

o

o

o

Q

o

o

0.5

0.0

o

8

o

o

oI I

o

o

o

o

o

o

o

8

8

o

o

-eo

Q

o

1 o

o

o

o

o

o

8

-eCF3I

o

o

8

o 8

o

-e—^—@—'—e

0 10 15

CF3I VOLUME %

20

Figure 47. Performance parameter vs. volume fraction of CF3I.
2. HIGH SPEED FLAMES AND QUASI-DETONATIONS 73

2.6 Conclusions

The following conclusions are made based on the results obtained:

a. Combustion and suppression processes in premixed hydrocarbon/air systems under highly

dynamic conditions can be more effectively studied in the modified two-sectional tube, permitting

clear discrimination of the combustion modes and performance among various gaseous extinguishing compounds.

b. There is a significant difference in combustion behavior between propane/air and ethene/air

mixmres: the combustion modes at higher velocities overlap totally in the ethene/air mixture for

the two geometric configurations while the propane/air mixture is characterized by a clear

separation between the combustion modes for the two arrangements. Also, the regime of

equivalence ratios for which combustion is detectable in the tube is much broader for the

ethene/air mixture. Furthermore, a detonation was unable to develop in the propane/air mixture

when the spiral insert was taken out of the tube.

c. The ethene/air flame in the quasi-detonation wave under suppression very closely follows the

shock wave with the same velocity. The distance between the flame and the shock increases with

the amount of an extinguishing agent. At extinguishment, the flame disappears, while the residual

shock still exists.

d. The presence of a hydrogen-containing suppressant in the ethene/air mixture results in a signifi- cant increase in pressure ratio relative to that for the pure combustible mixture. The phenomenon

occurs also for the compounds not containing hydrogen atoms at relatively lower concentrations.

The impact is generally weaker for stoichiometric and rich mixtures than it is for lean mixtures.

e. C3Fg is the most effective extinguishing compound in suppressing and attenuating flame/shock

systems in the lean, stoichiometric, and rich ethene/air mixtures under highly dynamic conditions

in the detonation/deflagration tube.

f. Depending on their concentrations, the presence of the three extinguishing compounds in the

propane/air mixtures causes the combustion either to be enhanced or suppressed, often with

complex extrema exhibited. The behavior is diminished, however, when the mixture becomes

richer in fuel content.

g. CF3I is the best agent for attenuating shock pressure ratio in the lean, stoichiometric and rich

propane/air mixtures. Such performance may be attributed to the significant endothermicity of the

decomposition process of CF3I during the passage of the shock through the mixtures under

investigation.

The bottom line is, the conclusions drawn from the previous NIST study have been confirmed.

FC-218 provides the most consistent performance over the widest range of fuel/air mixtures and tube

geometries. The CF3I has the greatest positive impact at low partial pressure fractions, but exhibits

non-monotonic behavior of flame speed and shock pressure ratio at increasing concentrations. The
74 2. HIGH SPEED FLAMES AND QUASI-DETONATIONS

dangerously high over-pressures previously exhibited by HFC- 125 were not observed during suppression under more moderate (and likely) combustion conditions. Based upon the results from this section alone, none of the three agents can be eliminated from consideration as candidates for dry-bay

applications.

2.7 Acknowledgments

This research was supported by the U.S. Naval Air Systems Command and the U.S. Air Force, Wright

Laboratory. Wing Tsang, Anthony Hamins and Valeri Babushok of NIST are gratefully acknowledged

for helpful discussions during the course of this work. The contribution of Wendy Chien to the

experimental phase of this work is highly appreciated. Richard Thorpe, Science Applications

International Corporation, is gratefully acknowledged for his editorial contribution.

2.8 References

Baker, W.E., Cox, P.A., Westine, P.S., Kulesz, J.J. and Strehlow, R.A., Explosion Hazards and

Evaluation, Elsevier, 1983.

Bennett, M., Survivability Enhancement Branch, Wright-Patterson AFB, personal communication,

1993.

Carbaugh, S.G., Survivability Enhancement Branch, Wright-Patterson AFB, Minutes of T2

Meeting, October 14, 1993.

Chapman, W.R. and Wheeler, R.N.V., "The Propagations of Flame in Mixtures of Methane and

Air. Part IV. The Effect of Restrictions in the Path of the Flame," Journal of the Chemical Society,

London, p. 2139 (1926).

Gmurczyk, G., Grosshandler, W., Peltz, M. and Lowe, D., "A Facility for Assessing Suppression

Effectiveness in High Speed Turbulent Flames," Eastern States Section Conference, The Combustion

Institute, Princeton, 1993.

Gmurczyk, G.W., Grosshandler, W.L., and Lowe, D.L., "Suppression Effectiveness of Extinguishing Agents under Highly Dynamic Conditions," Fire Safety Science - Proceedings of the Fourth

International Symposium, T. Kashiwagi, ed., International Association of Fire Safety Science, USA,

pp. 925-936, 1994.

Gmurczyk, G., and Grosshandler, W., "Suppression of High-Speed Flames with Cj-Halocarbons,"

Twenty-fifth Symposium (International) on Combustion, The Combustion Institute, in press, 1995.

Grosshandler, W.L., Gann, R.G., and Pitts, W.M., editors. Evaluation of Alternative In-flight Fire

Suppressants for Full- scale Testing in Simulated Aircraft Engine Nacelles and Dry Bays, NIST

Special Publication 861, Gaithersburg, April 1984.

Lee, J.H.S., "Fast Flames and Detonations," Chemistry of Combustion Processes, The American

Chemical Society, 1984.

Lee, J.H., Knystautas, R. and Chan, C.K., "Turbulent Flame Propagation in Obstacle-filled

Tubes," Twentieth Symposium (International) on Combustion, The Combustion Institute, pp. 1663-1672, 1984.

Lefebvre, M.H., Nzeyimana, E., Van Tiggelen, P.J., "Influence of Fluorocarbons on H2/02/Ar

Detonations Experiments and Modeling," Progress in Astronautics and Aeronautics, AIAA, 1992.

Nettleton, M.A., Gaseous Detonations: Their Nature, Effects and Control, Chapman and Hall,

1987.
2. HIGH SPEED FLAMES AND QUASI-DETONATIONS 75

Peraldi, O., Knystautas, R. and Lee, J.H., "Criteria for Transition to Detonation in Tubes,"

Twenty-first Symposium (International) on Combustion, The Combustion Institute, pp. 1629-1637,

1986.

Sheinson. R.S., Penner-Hahn, J.E., and Indritz, D., 1989, "The Physical and Chemical Action of

Fire Suppressants," Fire Safety Journal 75, pp. 437-450.

Taylor, B.N., and Kuyatt, C.E., Guidelines for Evaluating and Expressing the Uncertainty of N1ST

Measurement Results, NIST Technical Note 1297, Gaithersburg, September 1994.

Westbrook, C.K., "Chemical Kinetics of Hydrocarbon Oxidation in Gaseous Detonations,"

Combustion and Flame 46, 191 (1982).

77

3. PHOTODEGRADATION OF CF3I

Marc R. Nyden

Building and Fire Research Laboratory

Contents

Page

3. PHOTODEGRADATION OF CF3I 77

3.1 Introduction 78

3.2 Experimental 78

3.3 Results 83

3.4 Conclusions 94

3.5 Acknowledgments 94

3.6 References 94
78

3.1 Introduction

3. PHOTODEGRADATION OF CF3I

CF3I has been identified as a leading candidate for the replacement of haion fire extinguishing agents

because of its high degree of effectiveness as a flame suppressant and its short tropospheric lifetime

and low ozone depletion potential (Nyden, 1994). Although this compound is known to undergo rapid

photolysis in the presence of sunlight, the byproducts of the photodegradation process have not been

determined. It is also not known whether this agent will photolyze significantly when it is exposed to radiation from fluorescent lamps and other common sources of indoor light. This knowledge is essential to assess the stability of this agent and the risk of human exposure to it and its byproducts.

The ultraviolet absorption spectrum of CF3I, which is displayed in Figure 1 (Fahr, 1994 ), exhibits a peak which is centered near 270 nm and extends beyond 350 nm. This is the so-called A

band which is attributed to the superposition of electronic transitions from the ground state (X) to three low lying excited states which are designated, in order of increasing energy, ^Qj, ^Qq and 'Qj

(Mulliken, 1935, 1940). These states correspond to the spin orbit couplings which arise from the

promotion of an electron from a lone pair orbital (5 pn) on the iodine to an antibonding orbital (a ) which is localized in the region between the carbon and iodine nuclei.

Photoexcitation of the A band results in the dissociation of the C-I bond with unit quantum yield

in accordance with Equation (1) (Felder, 1992; Van Veen, et ai, 1985).

CF3 + / ePj)

CFJ-_^ .^J. (1)

'3 ^ CF^ + I* i^Pi)

2

The excited spin orbit state of the iodine (I ) correlates with the Qq state, whereas the ground state iodine (!) corresponds to the separated atom limit of the 'Q, and ^Qj potential energy curves. The

dissociation energy of the C-I bond is about 222 kJ/mole (Felder, 1992).

The spectrum of a fluorescent lamp, obtained in this laboratory (Figure 2), indicates a measurable

radiative flux out to about 310 nm. The observation that this spectrum overlaps with the absorption

spectrum of CF3I suggests that this agent should undergo appreciable photolysis when it is exposed to radiation from fluorescent lights. The objectives of this study are to determine the extent of this degradation and the nature of the compounds which are likely to be produced when CF3I is released in an indoor environment.

3.2 Experimental

The photodegradation experiments were conducted in a quartz cell which is transparent to ultraviolet

and visible radiation with wavelengths longer than 200 nm. The cell was cylindrical in shape with a

diameter of 5 cm. The long axis of the cell provided an optical path length of 10 cm for the

spectroscopic measurements which were made before and after irradiation of the sample. The

windows of the cell were made out of CaF2 which is transparent from about 10000 nm (1000 cm"^) in the infrared all the way to about 165 nm (60606 cm"') in the ultraviolet. Independent experiments were conducted using either a 1000 watt xenon arc or a 40 watt fluorescent lamp as the photolytic

source. A schematic of the experimental setup is depicted in Figure 3.
3. PHOTODEGRADATION OF CF3I
3. PHOTODEGRADATION OF CF3I

400 300

wavelength

Figure 2. The near UV-VIS spectrum of a common fluorescent lamp.
3. PHOTODEGRADATION OF CF3I

Photolytic

Source

Spectroscopic

Source

Quartz Cell

Figure 3. Schematic showing the experimental setup for the photodegradation experiments.
82 3. PHOTODEGRADATION OF CF3I

The CF3I used in these experiments was obtained from PGR, Inc. The stated purity was 99 %

which was confirmed by GC/MS analyses performed on a Hewlett Packard 571 OA gas chromatograph

with an open split interface to a HP 5970 mass selective detector. Samples containing approximately

6.7 kPa of the agent, either by itself or mixed with a variety of other gases, were irradiated for 30 min

with the high intensity Xe arc lamp. The experiments designed to measure the extent of degradation

in the radiation field generated by the fluorescent lamp, however, were conducted over a much longer

period of time (several days) because the rate of photolysis was much slower than it was with the Xe

arc. The effects of adding various gases including O2, Ar and air were also investigated.

The products of the photolysis of GF3I were characterized by ultraviolet (UV), visible (VIS) and

infrared (IR) spectroscopy. The UVA^IS spectra were measured over the range extending from about

250 nm to 850 nm using a fiber optic spectrometer manufactured by Ocean Optics, Inc. in conjunction

with either a xenon flash (UV) or a tungsten halogen (VIS) source. The IR spectra, which were

measured at 1 cm"^ resolution and signal averaged over 64 scans, were obtained using a Midac open

path Fourier transform infrared (FTIR) spectrometer.

Quantitative estimates of the fraction of unreacted agent (f) were obtained from least squares fits of the spectra of the original to the spectra of the photolyzed mixtures in accordance with

Equation (2).

fA^iv)A^iv)dv

f - ^^7 . (2)

where Aq(v) and Ap(v) are the measured values of the IR absorbance of the original and photolyzed

mixtures at the frequency, v, and the integration is over the band in GF3I which extends from 2272 - 2230 cm"'. The extent of degradation is the fraction of reacted agent which is given by 1-f. The

uncertainties in the spectroscopic measurements were computed as the standard error of the estimated

concentration using Equation (3) (Haaland, 1980).

^ , (3)

in - DfA^i^fdv

where n is the number of data points in the specified frequency range.

Gertain trade names and company products are mentioned in the text or identified in an illustration

in order to specify adequately the experimental procedure and equipment used. In no case does such

identification imply recommendation or endorsement by the National Institute of Standards and

Technology, nor does it imply that the products are necessarily the best available for the purpose.
3. PHOTODEGRADATION OF CF3I 83

In addition, there were a number of sources of systematic error which might have affected the

reproducibility of the experimental results. These include variations in the composition of the sample

mixtures resulting from leaks and contamination of the photolytic cell, as well as variations in exposure times and in the intensity of the photolytic source which could have affected the amount of

energy absorbed by the sample. The magnitudes of the resulting errors are difficult to quantify,

however, past experience suggests that they can be as much as an order of magnitude larger than the

uncertainties in the spectroscopic measurements calculated from Equation (3).

3.3 Results

The first set of experiments was designed to determine the extent of degradation of the pure agent in the presence of high intensity radiation from the Xe arc lamp. The expected reactions are summarized

in Equation (4).

In fact, a comparison of the IR spectra measured before an after the irradiation (Figure 4) indicate that the net effect of the photolysis was minimal, presumably because the reverse reaction is favored over

the reactions leading to the formation of new products. This is consistent with the entries in the NIST

kinetic database which indicate that the rate constant for the formation of CF3I from CF3 + I is about

twice as fast as the corresponding value for the formation of C2F^ + I2 (Mallard, 1993). The observed

extent of degradation, as estimated by least squares analysis over the region between 2230 cm"' - 2272 cm'\ was only 5.5 % (Table 1). The formation of a small amount of C2Fg, however, is apparent

from the presence of a sharp peak centered at about 1250 cm'* in the difference spectrum displayed at the top of Figure 5. This assignment may be confirmed by comparing the spectrum in Figure 5 with

the reference spectrum of C^^^^ in Figure 6 (Sadtler, 1993).

Additional measurements were performed to determine whether the presence of an inert gas would

have an effect on the photolysis of CF3I. In these experiments, a mixture consisting of 8.6 kPa of Ar

and 6.7 kPa of CF3I was exposed to the radiation from the Xe arc lamp for 30 min. The extent of

degradation was about 4.1 % (Table 1); which is comparable to the value obtained in the photolysis of

the pure agent. The small discrepancy in the extent of degradation (1.4 %) is indicative of the range

of values obtained in repetitive experiments and is most likely due to the systematic errors delineated

in the previous section. On this basis, it was concluded that the effect of the inert gas was minor.

The presence of even a small amount of O2, however, has a dramatic effect on the

photodegradation chemistry of CF3I. This is apparent in the comparison of the UV spectra (Figure 7)

of a mixture containing 6.7 kPa of CF3I and 0.94 kPa of O2 taken before (top) and after (bottom)

irradiation for 30 min with the Xe arc lamp. A thick, white smoke formed almost immediately upon

applying voltage to the arc. Shortly thereafter, a brown residue was observed to form in the crevices

where the windows were joined to the cell. The IR spectra displayed in Figure 8 exhibit a significant

loss of spectral intensity in the region between 2230 cm'' and 2272 cm'\ which correlates with

depletion of the agent, and a concomitant increase in absorbance in the region between 2300 cm"' and

2400 cm"', which is associated with the formation of CO2. At the same time, the presence of I2 was

indicated by the appearance of a new feature centered at about 530 nm in the VIS spectrum which is

2CF3/ - 2CFj» + 2/. - (4)
84

3. PHOTODEGRADATION OF CF3I

Figure 4. The IR spectra of CF3I taken before (top) and after (bottom) irradiation for 30 min

with the Xe arc lamp. . . , .

.
3. PHOTODEGRADATION OF CF3I

.2^

2000 1500

Wavenumbers (cm-1)

Figure 5. The differences in the IR spectra taken before and after irradiation with the Xe

lamp of pure CF3I (top) and a mixture of CF3I and O2 (bottom).
86 3. PHOTODEGRADATION OF CF3I

Figure 6. Library spectrum of C2Fg.
3. PHOTODEGRADATION OF CF3I

245 315

VAu« length

4^ 501

lOO.Idte l.rA lO.ns 1. io l.or 4^ -dn -tdo

Figure 7. The UV spectra of a mixture of CF3I and O2 taken before (top) and after (bottom

irradiation for 30 min with the Xe arc lamp.
88 3. PHOTODEGRADATION OF CF3I

displayed in Figure 9. The extent of degradation, based on the change in optical density in the region

between 2230 cm'^ and 2272 cm"^ in the spectra of the original and photolyzed mixtures in Figure 8, was approximately 32 %. This result implies that the stoichiometry of the photo-oxidation process is 2.3 mol of CF3I per mole of O2 (i.e., 6.7(.32)/0.94). A ratio of approximately 2:1 is obtained if the

5 % degradation of the agent, which would be expected to occur upon depletion of all of the available

O-,, is taken into account. The peak centered at 1950 cm in the difference spectrum displayed at the bottom of Figure 5 is attributable to COF2. A spectrum of this compound, which was measured in this laboratory, is displayed in Figure 10. Carbonyl fluoride is an extremely toxic compound (Sax and Lewis, 1987)

which can react further with to produce HF as summarized in Equation (5).

The smaller peak at about 1850 cm was not definitively assigned but it is consistent with the

presence of compounds possessing a carbonyl group (C=0). There was an indication in the IR spectra

that ethanol, which was used to clean the cell, was present as a contaminant during some of the

photolytic experiments. It is possible, therefore, that this peak is associated with a product of a photochemical reaction involving the ethanol and that is not related to the photodegradation of the

agent.

A mechanism for the photo-oxidation of CF3I was put forward in a recent paper by Clemitshaw

and Sodeau who used FTIR spectroscopy to measure the quantum yield of the formation of COF2 in the photolysis of CF3I at 253.7 nm (Clemitshaw and Sodeau, 1995). Their mechanism, which is consistent with the reaction stoichiometry of 2 mol of CF3I to 1 mol of O2 which was observed in this study, is summarized in Equation (6).

Another set of experiments were performed to determine the extent of degradation of CF3I in the

presence of radiation from a common fluorescent lamp. The photolysis of a mixture of 6.7 kPa of

CF3I and 82 kPa of O2 was monitored for about 5 days. During this period of time, about 28 % of

the CF3I was transformed into products. The brown residue, which had been observed in the previous

photo-oxidation measurements, formed after the first few hours of exposure to the lamp.

COF2 + Hp - CO^ + 2HF (5)

CF^I"-^ CF^* + /•

CF^* + O2 + M - CFp^* + M

CFp^* + !• + M - CFpOl + M

CFpOI - COF^ + FOI

FOI + M - I* + OF + M

!• + !• + M - + M

OF + OF + M - F^ + + M

(6)
3. PHOTODEGRADATION OF CF3I 89

Table 1. Extent of Photodegradation Based on Eq.(2)

Conditions iAo(v)dv jAp(v)dv % Degradation

Vacuum

Xe arc lamp

4.55 4.32 5.5 ± 0.1

8.6 kPa Ar

Xe arc lamp

5.80 5.53 4.1 ± 0.2

0.94 kPa O2

Xe arc lamp

3.59 2.37 31.7 ± 0.3

82 kPa O2

fluorescent lamp

3.20 2.30 28.1 ± 0.2

95 kPa air fluorescent lamp4.33 3.34 23.2 ± 0.1

In the final experiment, the valve on the cell containing 6.7 kPa of CF3I was opened to the

atmosphere and allowed to equilibrate. The valve was closed after a few minutes and the resulting

mixture of CF3I and air was placed under the fluorescent lamp for approximately 24 hours. The idea was to simulate as closely as possible the environment inside of a building during a release of the

agent. The IR spectra before (top) and after 24 hours of exposure to the radiation from the fluorescent

lamp (bottom) are displayed in Figure 11. A comparison of these spectra indicate a 23 % reduction in optical density in the region between 2230 cm'^ and 2272 cm'' which is attributable to the loss of

CF3I, and a concomitant increase in optical density in the region between 2300 cm"' and 2400 cm"'

which is due to the formation of CO2.

The first order rate constant obtained from the observation of a reduction of 23 % in the

concentration of agent in 24 hours is 3 x 10"^ s"'. Not surprisingly, this is of the same order of

magnitude as the value which was calculated for the rate constant of the photolysis of CF3I in direct

sunhght (Nyden, 1994).

It is then useful to estimate the extent of degradation of CF3I and to assess the potential hazard to occupants should an accidental discharge take place. The magnitude of the threat to public safety will

vary significantly depending on the ventilation and lighting in the vicinity of the release. Thus, for

example, the rupture of a cylinder of CF3I in a well-lit storage area will pose a more serious threat than a controlled release of the same agent into a dark dry bay.

As an example of the former, consider an accidental release into 100 m^ room which is uniformly

lit with a series of 40 watt fluorescent fixtures. This is a conservative scenario as the mean photon

flux incident on the agent in this room (at ground level) would be expected to be far lower than it was

in the experiment where the photolytic cell was positioned only a few centimeters from the lamp. If the contents of a 5 kg storage cylinder were to be distributed uniformly throughout the space, the

concentration of agent would be about 0.6 % by volume (assuming a ventilation rate of zero). The

degree of degradation of the agent 5 min after the release, based on the value of the rate constant for
90 3. PHOTODEGRADATION OF CF3I

.5-

u

o

0-

2400 2200

Wavenumbers (cm-1)

Figure 8. The IR spectra of a mixture of CF3I and taken before (top) and after (bottom)

irradiation for 30 min with the Xe arc lamp.
3. PHOTODEGRADATION OF CF3I

Figure 9. The VIS spectrum of a mixture of CF3I and O2 taken after irradiation for 30 min

with the Xe arc lamp.
92 3. PHOTODEGRADATION OF CF3I
3. PHOTODEGRADATION OF CF3I

Wavenumbers (cm-l)

Figure 11. IR spectrum of a mixture of CF3I and air after irradiation for 24 hours with

fluorescent lamp.
94 3. PHOTODEGRADATION OF CF3I

photodegradation given above, would then be only about 0.1 %. Even if all of the fluorine atoms

were converted to HF (and none deposited on the walls), the concentration of HF would only be about

20 |iL/L by volume. Alternatively, a COF2 concentration of about 5 |iL/L would be possible if all of

the agent were to be converted to this compound (and none were to react further to form HF). Given

the conservative assumptions used in this estimation, it seems unlikely that anyone who evacuated

promptly would be seriously harmed by exposure to the degradation products. It would, however, be

beneficial to assess the effect of exposure to low concentrations of HF on the integrity of sensitive

electronic equipment in the vicinity of a release.

3.4 Conclusions

CF3I photolyzes in the presence of radiation from common fluorescent lights with a rate constant which is comparable to that in photolysis in direct sunlight. The potential byproducts of this process

include COFj and HF. The high level of toxicity associated with these compounds merits their

inclusion when considering the impacts of accidental releases of CF3I in well-lit, occupied spaces. However, the experiments performed as part of this study indicate that the indoor rate of photolytic

decomposition is quite low, and a calculation based on conservative estimates further indicates that the

resulting concentrations of the most hazardous degradation products (i.e., HF and COF2) would also be

expected to be quite low.

3.5 Acknowledgments

The GC/MS analyses were performed by Mr. Richard Harris of the Fire Science Division of NIST.

The author also wishes to acknowledge the helpful advice and guidance offered Dr. Bama Laszlo and

Dr. Robert Huie of the Chemical Kinetics and Thermodynamics Division of NIST.

3.6 References

Ciemitshaw, K.C. and Sodeau, J.R., "The quantum-yield of formation of CF2O in the gas-phase

photo-oxidation of CF3I at 253.7 nm," J. PhotoChem. and PhotoBio A:Chem. 86, 9-14 (1995).

Fahr, A., Nayak, A.K., Huie, R.E., "Ultraviolet absorption Spectrum of CF3I," Halon Options

Technical Conference: Albuquerque, NM, 465 (1994).

Felder, P., "The influence of the molecular beam temperature on the photodissociation of CF3I at 308 nm," Chem. Phys. Letters 197, 425-432 (1992).

Haaland, D.M. and Easterling, R.G., "Improved Sensitivity of Infrared Spectroscopy by the

Application of Least Squares Methods," Appl. Spectrosc. 34, 539-548 (1980).

Mallard, W.G., Westley, F, Herron, J.T., Hampson, R.F and Frizzell, D.H., "NIST Chemical

Kinetics Database - Ver. 5.0," NIST Standard Reference Database 17, (1993).

Mulliken, R.S., "Electronic Structures of Polyatomic Molecules. VII. Amonia and water type

molecules and their derivatives," J. Chem. Phys. 3, 506-514 (1935).

Mulliken, R.S., "Intensities in Molecular Electronic Spectra. X. Calculations on mixed-halogen,

hydrogen halide, alkyl halide, and hydroxyl spectra," J. Chem. Phys. 8, 382-395 (1940).
3. PHOTODEGRADATION OF CF3I 95

Nyden, M.R. "Flame Inhibition Chemistry and the Search for Additional Fire Fighting Chemicals," in Evaluation of Alternative In-Flight Fire Suppressants for full-scale Testing in Simulated

Aircraft Engine Nacelles and Dry Bays, NIST SP 861, National Institute of Standards and Technology,

Gaithersburg, MD, 467, 1994.

Sadtler Search Software, "IR Vapor Phase Library," Bio-Rad Laboratories, Inc. (1993).

Sax, N.I., Lewis, R.J., "Dangerous Properties of Industrial Materials," Van Norstrand Reinhold,

New York (1987).

Van Veen, G.N.A., Bailer, T., De Vries, A.E. and Shapiro, M., "Photofragmentation of CF3I in the A band," Chem. Phys. 93, 277-291 (1985).

97

4. EFFECT OF SUPPRESSANTS ON METAL FIRES

Thomas J. Ohlemiller and John R. Shields

Building and Fire Research Laboratory

Contents

Page

4. EFFECT OF SUPPRESSANTS ON METAL FIRES 97

4.1 Introduction 98

4.2 Description of Apparatus 99

4.2.1 Bum Configuration 99

4.2.2 Combustion Chamber 99

4.2.3 Flow System 102

4.2.4 Sample Ignition 104

4.2.5 Test Materials 104

4.2.6 Data Acquisition 105

4.3 Results and Discussion 105

4.3.1 Magnesium 105

4.3.2 Titanium 108

4.4 Conclusions 117

4.5 References 118
98 4. EFFECT OF SUPPRESSANTS ON METAL HRES

4.1 Introduction

This study is focussed on two specific metals, magnesium and titanium. These metals, usually in the

form of various alloys, are utilized in several locations in modem military aircraft. Magnesium is used

extensively in structural areas of aircraft; titanium also has structural uses but is most commonly found

in jet engines, for example, as compressor and turbine blades.

These metals, and alloys dominated by them, are flammable. Ignition of these metals requires

that they be heated to near their melting points: ca. 650 °C for magnesium and ca. 1800 °C for

titanium (somewhat lower for alloys). Their ease of ignition is much less than that of familiar solid

organic materials such as polymers which typically ignite when heated to 300-400 °C. Bringing the

metal surface to these temperatures is also more difficult than it is for organic materials {i.e., it requires a higher heat flux) because the metals have much higher thermal inertia values, thus requiring more heat to achieve a given surface temperature (Elrod et al., 1980) . Once ignited both metals are capable of intense, high temperature burning' as a result of the very

high heats of formation of their oxides. Titanium requires more restrictive conditions for continued

combustion than does magnesium, at least in part because its oxide accumulates in the molten metal

surface region, slowing the rate of oxidation (Clark, Moulder and Runyan, 1975). As a result, continued combustion requires either a significant pre-heating of the metal, elevated oxygen partial pressure

or shearing away of the molten oxide/metal layer (Clark, Moulder and Runyan, 1975; Rhein and

Baldwin, 1980). All of these conditions can be found in jet engines; the shearing is provided by the

high velocity of the gases and centrifugal forces (Fox, 1976; Anderson and Manty, 1978; Elrod and

Lyon; 1979; Strobridge et al., 1979; Rhein and Baldwin, 1980; Elrod et al., 1982; Baldwin et al., 1984).

The current study is directed toward the interaction of suppressants with these burning metals.

Although it is possible for the metal alone to be the burning material in need of suppression, the concern here is broader. Military aircraft fires in general are most frequently suppressed with halons.

A fire may start with a fuel or hydraulic fluid leak and spread, involving parts composed of magnesium or titanium. This fire will be manually or automatically flooded with halon compounds (or, in the future, their replacements). The use of halogen-containing suppressant agents on metal fires is considered hazardous and is not recommended (Andrews, 1981) but in the mixed materials environment of a military aircraft, where a fuel fire is the more likely threat, they are used, regardless of the

possible involvement of metal aircraft components. Experience to date, based on halon suppressants,

shows that this is an effective approach. The work here specifically addresses the concern of whether

candidate halon replacements will pose any greater hazard in their interaction with a metal fire than do

the currently used halons.

Halogens, particularly fluorine, can combine with magnesium or titanium with an evolution of more heat than is obtained from combining with oxygen. The heat of formation of magnesium

fluoride is 83 % higher than that of magnesium oxide. For titanium there are several possible products

but the most relevant comparison is probably titanium tetrafluoride which has a heat of formation that

is 70 % higher than that of titanium dioxide. Simple calculations (below) suggest that under some

conditions either of these metals might strip the halogens from a halon such as halon 1301 and evolve more heat than could be obtained from reaction with oxygen. Thus the suppressant could become an

In pure oxygen at one atmosphere pressure, the adiabatic combustion temperatures of magnesium and

titanium are both approximately 3000 °C (Strobridge et al., 1979).
4. EFFECT OF SUPPRESSANTS ON METAL FIRES 99

accelerant. There are warnings in the literature regarding this (McCutchan, 1954; Maykuth, 1964;

Kuchta et ai, 1965 & 1966; NASA, 1970; Rhein & Baldwin, 1980; Tapscott, et al., 1986). Unfortunately, most of these references describe experiments in which the effect of a halogen-containing agent on a metal fire cannot be clearly quantified. The fires are qualitatively observed to "accelerate" or

"flare-up" when sprayed with the halon agent from an extinguisher. The possible role of purely

physical factors, such as intensification due to scattering of the burning metal by flash vaporization of

liquid droplets of the halon, cannot be assessed.

Kuchta et al. (1966) did perform some more quantifiable experiments with heated titanium wires

in hot gas streams containing the decomposition vapors of various halocarbons. Various halon vapors were heated to 450 °C and then passed over titanium wires at 600 °C; the increase in wire temperature was measured. This increase should be reflective of the heat release rate due to reaction of the halon

vapors with the wire surface. Halon 1301 gave the least increase in wire temperature of any halon

examined though it was still significant. This result does confirm that halons will chemically react with titanium in an exothermic manner (under these conditions). The variation of effect with halon

type implies kinetic effects, perhaps both in extent of vapor decomposition and in rate of attack on the

metal. The data are too limited to resolve such issues. The present study is limited to examining the effect of the vapors of various halogen-containing

suppressants on the burning of pure magnesium and titanium rods in a slow flow of oxygen-containing

gas. The test configuration was instrumented to allow quantitation of the effect of the vapors on the

rate of burning. As will be seen, the results were not as the literature leads one to expect in that the

metal burning process was always either slowed or extinguished under the range of conditions

examined.

The agents examined in the presence of burning magnesium included halon 1301 (the reference

case), HFC- 125, HFC-227ea, FC-218 and CF3I. For titanium, time limitations narrowed the study to halon 1301, HFC-227ea^ and HFC- 125.

4.2 Description of Apparatus

4.2.1 Burn Configuration. For both metals, the burning configuration was upward propagation along

a thin rod with the oxidant gas flowing in the same (upward) direction. This configuration was chosen

so that gravity, by causing dripping of the molten metal in the case of titanium, would provide the

necessary fresh metal surface to yield sustained combustion. It was not feasible, in the time allotted for this study, to set up the high speed, elevated temperature flows used previously to study titanium

combustion propagation under jet engine conditions. Upward propagation was chosen because

Sato et al. (1983) showed that downward propagation can produce much more erratic results {i.e.,

highly variable propagation rates) due to the dripping molten metal haphazardly making contact with

the rod as it flows/falls downward, causing spot ignition. Oxidant flow velocity levels were limited to low values by practical constraints on mass flows borne of the use of bottled gases.

4.2.2 Combustion Chamber. Figure 1 is a not-to-scale, cross-sectional sketch of the overall

experiment; Figure 2 is a scaled, cross-sectional assembly drawing of the combustion chamber. The

chamber is made of brass and consists of three sections, assembled in such a way as to allow relatively

convenient cleaning and sample replacement. The bottom section accepts the incoming gas flow and

HFC-227ea is referred to throughout the remainder of this discussion simply as HFC-227.
100 4. EFFECT OF SUPPRESSANTS ON METAL HRES

Figure 1. Schematic of metal combustion experiment.
EFFECT OF SUPPRESSANTS ON METAL HRES

Figure 2. Scaled assembly drawing of combustion chamber.
102 4. EFFECT OF SUPPRESSANTS ON METAL RRES

directs it upward. This section also includes fittings for a pair of electrodes used in the metal ignition process. This section includes, as well, a small cup to catch and quench the drips of molten metal.

The middle section contains the sample rod (16 cm long) suspended in the gas flow, between two

windows for clear observation. The front window, through which a video camera records the burning

process, is Pyrex. The rear window is Vycor which provides light transmission out to about 2.5 \im.

A calibrated thermopile-type radiometer (Dexter Research Center, Model 2M-HS)^ views the burning

rod through this latter window (a non-imaged view), providing a signal proportional to the radiative

emissions of the bum zone. Early tests indicated that small burning particles emitted from the burning

titanium rods could pit the windows (which are only about 2.5 cm from the burning rod). Thus the

windows were lined with a thin (0.25 mm), reusable sheet of Kel-F polymer film which is essentially

transparent to all of the radiation passing through the Vycor or Pyrex windows. The upper section of

the chamber contains a bare wire thermopile consisting of four 0.125 mm diameter type S thermocouples wired in parallel; the four thermocouple junctions are arrayed in the flow. This thermopile samples

the average temperature of the gases passing through the chamber downstream of the combustion

process; it thus provides a measure of the convective heat output of the burning process. This measure

could be made absolute by calibration with an electrical energy source in the sample burning region;

this was not done here because only relative performance (to halon 1301 as a baseline) was of interest. The chamber can be operated at pressures up to 1.48 MPa (200 psig). A burst disk is incorporated in the system because information on the possible flammability of mixtures of the

candidate halon replacement agents in elevated oxygen atmospheres is not available. (It is known that some halogenated hydrocarbons are flammable in some oxygenated mixtures (Perlee, Martindill and

Zabetakis, 1966).) A practical limit of 5 grams of incoming oxidant gas per second was imposed to accommodate the use of bottled gases and the limited supplies of some of the halon replacement

agents. This fixed mass flow defines the outer envelope of achievable pressure and velocity conditions. The maximum oxidant flow velocity in the sample section thus varies from 3.7 m/s at 0.1 MPa

(atmospheric pressure) to 0.27 m/s at 1.48 MPa.

4.2.3 Flow System. Figure 3 is a schematic of the flow system in its simplest configuration, used

when the agent vapor pressure at room temperature was sufficient to supply the desired add-on level to the oxidant flow. All pressures were measured with gages rated by the manufacturer at ± 0.5 %

accuracy. All flows were measured with rotameter-type flowmeters which had been calibrated with a

dry test meter in the pressure and flow range of usage; flow accuracy is estimated to be ± 2-3 %.

Figure 3 shows that the agent vapor could be added upstream or downstream of the combustion

chamber. As will be described below, this bypass feature was necessary because the agent vapor

tended to attack the sample ignition system. When the agent was introduced upstream of the chamber,

a static mixer assured concentration uniformity in the oxidant flow prior to the chamber.

When flow of an oxidant stronger than air was needed, it was obtained by mixing air and pure

oxygen. This was done using a two tube, mixing rotameter for which both tubes were calibrated as above. It should be noted that oxygen level data reported here are those prior to the agent vapor addon.

In order to operate the chamber at pressures above the room temperature vapor pressure of the

agent of interest, it is necessary to meter and introduce the agent as a liquid, mix it with the oxidant

Certain trade names and company products are mentioned in the text or identified in an illustration

in order to specify adequately the experimental procedure and equipment used. In no case does such

identification imply recommendation or endorsement by the National Institute of Standards and

Technology, nor does it imply that the products are necessarily the best available for the purpose.
4. EFFECT OF SUPPRESSANTS ON METAL RRES 103

Burst Disk

Chamber

Pressure

Air From

Bottle

Exhaust

Emergency

Pressure

Dump

Static

Mixer

Agent

Flow

Rate

Agent

Figure 3. Flow system associated with combustion chamber.
104 4. EFFECT OF SUPPRESSANTS ON METAL FIRES

gas in a heater section then cool the total flow back to room temperature. A system to achieve this was in fact built. However, when it was found that the agent vapors attacked the ignition system, this approach had to be discontinued. In the flow system shown in Figure 3, the agent vapor needs to be

added a short distance upstream of the chamber so that the flow time to the chamber is short

(compared to the sample bum time). The system required to vaporize the agent with heating and then

cool the entire gas stream was several meters in length. In some conditions, the sample could have

burned excessively before the agent vapor reached it. Thus all tests were confined to pressure conditions in which the desired agent level could be achieved using the vapors in the headspace above a

liquid agent reservoir.

4.2.4 Sample Ignition. Ignition of the small diameter metal rods proved to be more problematical

than expected. For magnesium, copper electrodes were used in the base of the combustion chamber to supply electrical current to a nichrome heating wire from a low voltage DC power supply. One end of

a length (about 4 cm) of magnesium ribbon (0.25 mm thick by 3 mm wide) was wound around the

bottom few mm of the sample rod; the other end was folded over the nichrome wire. Heating of the

nichrome wire was to ignite the magnesium ribbon and the ribbon would ignite the rod. This was

satisfactory as long as there was no agent vapor in the oxidant flow. When agent vapor was present, it attacked the nichrome wire, frequently causing it to break before the magnesium ribbon could ignite.

This behavior necessitated the agent vapor bypass path shown in Figure 3. With this in place, the

agent vapor was introduced into the chamber exhaust line during the sample ignition process; this allowed the agent flow system to achieve a steady-state condition. When the sample rod had clearly

ignited and propagated combustion 1-2 cm, the agent vapor was switched into the chamber inlet line,

reaching the sample bum zone a few seconds later. Titanium, with its much higher melting point, could not be ignited in the above manner, even

though the sample rod was only about 1 mm in diameter. An attempt was made to use an electric arc welder boosted by a high frequency, high voltage arc stabilizer in the manner of Baldwin, Beach and

Rhein (1984). The arc welder itself was rated at 230 amperes AC; the arc stabilizer superimposed

3500 VAC of radio frequency oscillations to assist in arc formation and retention. The arc stabilizer

proved to be totally ineffective for reasons which could not be established; it would not assist in arc formation even for atmospheric air gaps of a small fraction of a millimeter. Furthermore, it was very

disruptive to the data acquisition system despite careful attempts at electromagnetic shielding.

Ultimately, it was found that the titanium rods could be ignited by the contact resistance heating between tungsten electrodes and the surface of the sample rod. This required about 200 amperes and

took as much as 20-30 seconds to yield ignition.

The combustion of the sample was recorded with a Hi-8 video camera which viewed the sample

through a #12 welder's light filter. Progress of the propagation was also sensed by three blackened

thermocouples (0.13 mm diameter; chromel/alumel) recessed into the sidewall of the middle section of

the combustion chamber at intervals of 4.0 cm. Radiative heating of these was sufficient to yield a peak in voltage output as the combustion zone passed the thermocouple location.

4.2.5 Test Materials. The 1.6 mm diameter magnesium rod was 99.9 % pure. It appeared to have

been extruded. The titanium rod was 1.14 mm in diameter; it was commercially pure grade, used for

welding purposes. It appeared to be wire-drawn. The initial focus of the titanium work had been on

alloy 6-2-4-2 (6 % aluminum, 2 % molybdenum, 4 % zirconium and 2 % tin) which is one of several

forms in which titanium is frequently used in aircraft. However, it was found that ignition of the

available 3 mm square rod required an atmosphere of more than 60 % oxygen. This alloy could not be obtained in the small diameter which the above ignition system could more readily handle. Thus

the focus was shifted to the pure metal, obtainable in the smaller diameter form. The qualitative
4. EFFECT OF SUPPRESSANTS ON METAL HRES 105

conclusions of this work are believed to be applicable to high titanium alloys as well as the pure metal

used here.

4.2.6 Data Acquisition. All data channels were hard-wired to a Strawberry Tree Workmate data

acquisition board through an isothermal terminal board. Strawberry Tree Quicklog software was used

as the data acquisition program. Data were acquired at 2 Hz with 14 bit resolution.

4.3 Results and Discussion

4.3.1 Magnesium. An initial short series of tests was conducted using pure air with no added agent

to explore the possible sensitivity of the burning rate to other test parameters. Table 1 shows the

results of varying both the chamber pressure and gas flow velocity in the test section. Chamber

pressure was varied from 0.24 MPa (20 psig) to 0.79 MPa (100 psig); air flow velocity in the test section was varied from 13 to 42 cm/s. Note that the mass flow rate of air was held constant at one

of two values which differed by 42 %.

Chamber pressure variation at a constant air mass flux will hold the mass transfer rate of oxygen

to the buiTi zone constant while varying the gas phase chemical reaction rates between magnesium

vapor and oxygen. This is done in the first four tests of Table 1. In the next three tests, the pressure

(and thus the chemical reaction rate) is held constant but the mass flux is increased to enhance the

oxygen mass transfer rate. Here the 42 % increase in mass flux will yield roughly a 20 % increase in mass transfer rate since the mass transfer coefficient is approximately proportional to the square root of Reynolds Number for this configuration (Bird, Stewart and Lightfoot, 1960).

No effect on burning rate from either of these changes clearly emerges from the data of Table 1. The average burning rate at the lower air mass flow rate is 3.9 ± 0.3 mm/s; at the higher mass flow

rate it is 3.6 ± 0.4 mm/s. Here the uncertainties given are one standard deviation. The overlapping

uncertainties preclude any inference of a trend. It is noteworthy, however, that the much larger

percentage change in pressure {ca. 300 % for absolute pressure) did not clearly affect burning rate. This implies that the gas phase chemical reaction rate is not the controlling step here. The data are ambiguous on the possible control by oxygen mass transfer, but they are not inconsistent with this idea.

It is also noteworthy that the burning rate is not particularly sensitive to gas mass flow rate or to pressure so that the effects of any slight variations in these will not be confused with agent vapor

effects in the results below. Unfortunately, one reason that no effects of these variables emerge in Table 1 is that the data are somewhat noisy. The source of the noise appears to have been the rather

complex nature of the combustion zone. The magnesium oxide did not simply condense to an aerosol

and float away in the gas flow. Rather, a large fraction of the oxide condensed locally around the

immediate area where the magnesium vapors were emerging from the burning rod, ultimately leaving a rough replica of the rod behind, attached to the unbumed portion. (This tendency for the oxide to form complex solid structures in the region of the burning zone has been noted in previous studies

(Law and Williams, 1974).) Here, the condensation process seemed to interfere with the vaporization/burning process so that the bum zone was irregular and extended of the order of ten rod

diameters. Varying heat transfer to the metal through the irregular oxide mass probably caused the

burning rate to vary. This irregular bum zone also had the potential to fool the three sensor thermocouples in the side of the chamber, originally intended to provide buming rate information. Thus all of

the buming rate information was obtained instead from the video tapes of the tests where the leading

edge of the bum zone could be easily followed.
106 4. EFFECT OF SUPPRESSANTS ON METAL FIRES

Table 1 . Magnesium Burning in Pressurized Air

Test

I.D.

Ambient

pas

iviass now raie of gas

(g/s)

vjas now

velocity

(cm/s)

Chamber

pressure

(MPa/psig)

Bum rate

^111111/ o J

Mg-24 AIR 1.33 42 0.24/20 3.66

Mg-21 22 0.44/50 4.23

Mg-4 ti 13 0.79/100 4.14

Mg-5 3.75

Mg-50 1.89 36 0.27/25 4.14

Me-52 „ 3.45

Mg-53 „ ti M 3.33

Mg-16 AIR + NITROGEN

(10 % ADD-ON)

1.46 13.9 0.79/100 3.75

Mg-19 It n fi 3.53

Mg-20 It M M n 3.57

Table 2 shows the result of adding on 5 or 10 weight percent of various agent vapors to the

pressurized air flowing past the burning magnesium. Note that at least two replicates were run at each

test condition. The surprising result in all cases was extinction of what was initially a vigorously

propagating burning zone. (Recall that the sample rod was burning for about 2 cm before the agent

vapor reached it.) Extinction typically occurred in 5-10 seconds, which is probably not much longer

than the time to fully displace the agent-free atmosphere around the burning zone. Note that these

add-on levels are low compared to those previously found necessary to extinguish various hydrocarbon

flames (Hamins et ai, 1994) in a fairly similar situation (cup burner). For hydrocarbons, an add-on

level of 25-30 % agent vapor to air was typically necessary. (Halon 1301 and CF3I were exceptions,

with the former requiring only 12-28 % add-on by weight and the latter, equally effective on a molar

basis, about 25 % less than this on a mass basis.)

The reason for the present results is not clear. There is no obvious chemical basis for expecting

the agents to act as suppressants here. Halon 1301 is believed to achieve its suppressant effect in hydrocarbon flames by trapping hydrogen radicals (Nyden and Linteris, 1994). Such radicals are not involved in magnesium burning (Tapscott et al, 1986). Most, if not all, of the extinguishing

"^wo tests, Mg-44 and Mg 45, with FC-218 were appreciably slower, with roughly 2/3 of the rod

consumed before extinction occurred.
4. EFFECT OF SUPPRESSANTS ON METAL HRES

Table 2. Magnesium Burning in Pressurized Air with Agent Vapor^

107

Test vT D

Agent

(Add-on,

wgt %)

Gas flow

velocity

(cm/s)

Chamber

Pressure (MPa/psig)

Bum Rate

(_mm/s)

Mg-11, 12,13, 14 HALON-1301

(10 %)

13 0.79/100 EXTING.

Mg-25,26,27 37 0.27/25 EXTING

Mg-47,48 HALON-1301

(5 %)

13 0.79/100 EXTING.

Mg-39,40 37 0.27/25 EXTING.

Mg-41,42 HFC- 125 (5 %) 13 0.79/100 EXTING.

Mg-37,38 37 0.27/25 EXTING.

Mg-39,40 HFC-227 (5 %) 37 0.27/25 EXTING.

Mg-30,31 FC-218 (10 %) 37 0.27/25 EXTING.

Mg-43,44,45 FC-218 (5 %) 14 0.72/90 EXTING.

Mg-32,33 II 37 0.27/25 EXTING.

Mg-34,35,36 CF3I (5 %) 37 0.27/25 EXTING.

^ Air mass flow rate held constant at 1 .33 g/s for all tests.

effectiveness in hydrocarbon fires of the halon alternative agents examined here (except CF3I) was

attributed to the simple physical effect of thermal dilution (Hamins et ai, 1994). That is, these agents,

by their heat capacity effects, simply reduce the flame temperamre to the point of extinguishment. A

similar effect may be responsible here but the substantially lower agent concentrations (<5 % by

weight) make this less likely.

A few tests were run with nitrogen added-on at 10 weight percent to the air flow. Table 1 shows

that this had no significant effect on the measured burning rate at 0.79 MPa (100 psig). This result is ambiguous, however, because nitrogen, which, of course, is the dominant gas in the air, may not be a

passive diluent in this system. Magnesium nitride has a heat of formation that is 75 % of that of

magnesium oxide. Thus magnesium can burn in pure nitrogen. Whether nitride formation is significant (and thus whether the add-on was a diluent or a reactant) in the present experiments
108 4. EFFECT OF SUPPRESSANTS ON METAL HRES

depends on its ability to compete with the oxygen/magnesium kinetics. Experiments in argon/oxygen

mixtures could probably resolve the issue of whether an equivalent thermal capacitance^ of a truly

inert diluent is capable of extinguishing the burning magnesium. There was insufficient time in the

present study to pursue this issue further. Some simple calculations, reported below with the titanium results, do, in fact, suggest a very

weak slowing effect on the burning rate (not extinction) is possible for the conditions used here. As

will be seen, the basis is oxygen dilution and retarded mass transfer of the alternative oxidant.

In spite of the incomplete picture as to the causes underlying the results in Table 2, those results do indicate that, at least down to the 5 % add-on level, there is no difference in the impact of the old

and new agents. Halon 1301 and the four candidate agents tested here are equally prone to extinguishing the burning process in this configuration. Differences could exist in the minimum add-on level required for extinguishment. This would require substantially more testing to establish.

It should be noted that the burning configuration used here incorporates no radiative self-feedback

of energy among surface elements of the fuel. Such feedback is stabilizing to the combustion of any

solid fuel, making it more difficult to extinguish. Thus the present result indicating that 5 or 10

weight percent of these agent vapors can extinguish a burning magnesium rod does not necessarily

carry over to more complex configurations, such as would be encountered in a burning aircraft structural fire.

4.3.2 Titanium. The test conditions and burning rate results for titanium are summarized in Table 3. The tests here are less extensive than with magnesium due to time constraints. Note that, in the tests done with no agent present, cutting the absolute pressure by greater than a factor of two (at a fixed mass flow rate) and increasing the ambient oxygen level by only 20 % yielded a roughly 30 %

increase in the linear burning rate. This suggests (but does not prove) that mass transfer of oxygen is an important controlling factor. Since the oxidation reaction is evidently mainly heterogeneous here,

within the molten oxide/metal mix (Clark et al, 1975), it is likely to be largely uninfluenced by the

noted pressure change.

Note that the no-agent burn velocities for titanium are generally significantly less than those for

magnesium, despite that fact that the ambient oxygen level is considerably higher in the titanium cases. Burning rate models for metals such as that of Sato and Hirano (1988) should aid in explaining such

differences but the requisite kinetic data on the oxidation reactions are not available. Thus the reasons

for these differences are not clear at present but a wide variety of property differences between the two

metals could be contributing.

In all of these tests (including those with agents present in the ambient atmosphere) the behavior

of the burning metal was non-steady. A glow zone would progress up the metal rod several (up to about ten, depending on agent presence) diameters, sometimes seeming to proceed in steps of about one diameter. Surface tension in the elongated glowing zone behind the leading front would

periodically cause this part or all of this zone to contract into a sphere. The extent of contraction to a sphere depended on the presence of agent vapors, as discussed below. This sphere would frequently

begin to eject numerous fine, glowing particles. When the mass of the sphere exceeded the tensile

strength of the glowing zone to which it was attached, it dropped off, taking much of the glowing

material away below the propagation front. The front nevertheless continued and the cycle repeated

itself. This occurred 3 to 6 times during the course of a typical full length bum.

Thermal capacitance here means mass of the diluent times its heat capacity (J/g °C). Argon has less than half the heat capacity of nitrogen so its weight percent add-on would have to be increased

accordingly.
4. EFFECT OF SUPPRESSANTS ON METAL HRES

Table 3. Titanium Burning in Elevated Oxygen Atmospheres

109

Test

I.D.

weignt

percent

oxygen^

Chamber

pressure

(MPa/psig)

Gas flow

velocity

(cm/s)

Agent

(add-on

wgt %)

Bum rate (mm/s)

Ti 1 A 11-14

Ti-15

/in u. /y/ luu 1 1 11.94

Ti-20

Ti-21

14 NITROGEN

(10 %)

1.63

1.67

Ti-16

Ti 1 8 1 1- 1 o Ti-19

II 1 1I

J

HALON-1301 1.11 CA 1 UNO.

EXTING.

1 1-ZZ Ti-23

Ti-24

ri-25

H 0.79-0.91

0.79-0.95

12 HFC- 125

(10 %)

1.3/

EXTCsTG.

EXTING.

1.92

Ti-39

Ti-40

50 0.31/30 33 NONE 2.53

2.59

Ti-37

Ti-38

34 HALON 1301

(15 %)

1.07

1.05

Ti-34

Ti-36

I. 34 HFC- 125

(15 %)

1.07

0.96

Ti-32

Ti-33

34 HFC-227

(15 %)

1.88

1.90

Ti-26

Ti-27

50 0.38/40 27 NONE 2.48

2.41

Ti-28

Ti-29

28 HALON 1301

(15 %)

1.24

1.20

This is the oxygen percentage by weight in the oxidant gas flow exclusive of the presence

of the agent.
no 4. EFFECT OF SUPPRESSANTS ON METAL HRES

The enhanced heat release rate and/or heat transfer rate to the gas as the spheres developed and

dropped was sometimes clearly visible in the output from the gas stream temperature sensor. Figure 4

shows such a case;^ this is for 50 % oxygen and 15 % halon 1301 vapor. Note the ambiguity in the

preceding sentence as to the cause of the spikes; the shape changes and varying area available for

convective heat transfer may be mimicking variations in heat release rate. A further source of

ambiguity in the signal from this sensor, caused by the dripping phenomenon, is a possible erratic contribution from drips which hit and stuck to the ignition electrodes,^ continuing to bum for several

seconds, rather than falling down into the quenching cup at the base of the chamber. Thus the output

from this sensor has to be viewed with caution for the titanium tests; it cannot clearly differentiate heat

release rate contributions from the primary burn zone from other sources of fluctuation. As a result,

other data from this sensor are not reported here. Figure 5 shows the simultaneous output (same test as Figure 4) from the radiation thermopile

viewing the sample through the back window of the chamber. The timing of the molten metal drips is indicated. Note that the halon 1301 vapor control valve was switched 8 seconds after the initial ignition spike and would be arriving in the neighborhood of the sample about one second later (at

about 14 seconds on the time scale of Figure 5). Except for the first drip, every drip is accompanied

by a sharp drop in radiation output as the burning area shrinks sharply. As it builds up again, here in the form of a growing drop attached to the propagating front, the radiation again increases. However,

each successive peak is smaller. It appears that the main reason for the monotonic peak shrinkage is erratic depositing of particles ejected from the burning zone on the rear chamber window, partially

blocking some of the radiation. (The last peak is reduced by vignetting from the window frame.)

Figure 6 shows the radiation sensor output for a case at the same conditions as in Figure 5 except

that there was no agent vapor present. The differing shape of the curve apparently corresponds to the

different shape of the overall burning zone. When no agent was present, the burning zone consisted of

several diameters of glowing rod, at the bottom of which the sphere of molten material accumulated

before dropping off. When the sphere did drop off there was still an extensive glowing area that maintained the radiation signal at a fairly high level. When the agent vapors were present, as in Figure 5, the glowing rod section above the sphere was minimal in length leaving the sphere as the

principal radiation source. When it dropped off the radiation signal dropped by a greater percentage.

The mechanism whereby the agent vapors had this effect on the burning zone is not clear. One can

speculate that it involves the relative rates of propagation (faster in the no-agent case) and of sphere

formation.

In spite of these complications, there was a general tendency for the average radiation signal for

the no-agent cases to be higher than for the cases with an agent vapor present. It is probable that this was due in part to the average temperature of the burning zone being hotter in the no-agent cases, but

the confounding effects of shape changes and non-quantifiable window deposits make this unprovable

with the available data. A lower burning zone temperature with the agent vapors present would, of

course, be consistent with generally slower burning which Table 3 shows to occur in such cases, as compared to the no-agent cases. Before proceeding to examine the titanium burning rate data further, it is of interest to make a

simplified estimate of the potential effect on heat release rate from the burning metal due to participation of the agent vapors, if the burning process is taken to be mass-transfer limited. As was indicated

Time zero is arbitrary in this Figure; ignition occurs when the sensor signal becomes non-zero.

^The chamber was tilted backwards slightly to minimize the occurrence of this phenomenon but it could not be avoided entirely.
4. EFFECT OF SUPPRESSANTS ON METAL HRES 111

TITANIUM ROD (TEST Ti-29)

AGENT: HALON 1301 (15%)

CHAMBER PRESSURE = 40 (psig)

OXIDANT GAS = AIR + O (50% O TOTAL)

20 T 1 1 1 1 1 1 1

1 1 1 1

1 1 r -| 1 r

0)

COca

0)

oc

CO

CD I

>

O

CD>

C

o

O

o

CO

90 - O

3

Q.

O

o

c

CD

60

30 -

Arc welder

ignition transient

0

_l I I.

ARROWS INDICATE

TIME OF MELT DRIPS

1^ ^ t f f

0 20 •40 60 80

Relative Time (s)

00 120

Figure 4. Output from convective heat release rate sensor.
112 4. EFFECT OF SUPPRESSANTS ON METAL FIRES

0.25

TITANIUM ROD (TEST Ti-29)

AGENT: HALON 1301 (15%)

CHAMBER PRESSURE = 40 (psig)

OXIDANT GAS = AIR + (50% TOTAL)

~i I I 1 I I I

I I I II

\ 1 I

I

1 I r I

CME

o

0.2

ARROWS INDICATE

TIME OF MELT DRIPS

o 0.15

w

c

c

o

(0

cr

0.

0.05

0

20

J I L 40 60 80

Relative Time (s)

100 20

Figure 5. Output of radiation sensor for case with halon 1301 vapor present.
4. EFFECT OF SUPPRESSANTS ON METAL FIRES 113

TITANIUM ROD (TEST Ti-26)

NO AGENT

CHAMBER PRESSURE = 40 (psig)

OXIDANT GAS = AIR + O (50% O TOTAL)

0.25

_ 0.2

o 0.15

0.

0.05

1—I—I—I—1—

r

n I 1 I I I 1 -|—I—

I—I—I—I—I—

|-

0

ARROWS INDICATE

TIME OF MELT DRIPS

_l I I LJJ I I I I I i L J I I I \ L 40 50 60 70 80 90

Relative Time (s)

100 110 120

Figure 6. Output from radiation sensor for case with no agent vapor present.
114 4. EFFECT OF SUPPRESSANTS ON METAL FIRES

previously, the data are not inconsistent with control by oxidant mass transfer through the boundary

layer around the burn zone. The literature leads one to expect the impact of the agent to be burning

rate enhancement. The source of such a potential burning rate or heat release rate enhancement is, of

course, the higher heat of formation of the metal fluorides compared to the oxides.

Here we denote rate of heat release as RHR, where:

RHR = m,H^= kA, (F,,^^ -

y„,^„^) / (1)

Here is the mass burning rate of the metal (g/s); is the heat of combustion of the metal by

oxygen per unit weight of metal burned (J/g); is the stoichiometric coefficient for oxidation (g 02/g

metal burned); k is the overall mass transfer coefficient between the flowing ambient atmosphere and

the burning metal surface (g/cm^ s); is the burning area of the metal; Y^^^^^ is the mass fraction of

oxygen in the ambient gas and Y^^ is the mass fraction of oxygen at the surface of the burning

metal. Note that RHR and mass burning rate, m^, are proportional; RHR is generally a more relevant measure of fire intensity. When the agent vapors are also present in the ambient gas they potentially

comprise a second oxidizer which can attack the metal surface in parallel with the oxygen. However,

note that they also displace some oxygen. Then one can write the total heat release rate as:

RHR' = A,i(HX(Y:^^^ - Y:^^^^) I ^ HMY^^^ - Y^^^^) I nj) (2)

Here subscript "a" refers to the agent vapors. Note also that the prime on the oxygen mass fraction

symbols indicates that they have been changed from their values in Equation 1 by the addition of the

agent vapor. That is,

y' = Y / (I + Y ) (3)

'^ox^as ox^as ' ^ a^as'

If the chemistry involving both of the oxidizers is fast compared to their mass transfer rates through the boundary layer, the surface values of both become very small and can be dropped. We

can estimate the effect of the agent addition by taking the ratio of Equation 1 to Equation 2:

RHRJRHR' = (HXY^^^JnJ / (HXYL^Jn^ - "AYa^M ^"^^

Note that the bum area parameter has been cancelled out on the assumption that it is unaffected by the

agent vapor. The discussion above regarding the radiation output implies that the average bum area probably does decrease somewhat in the presence of the agent vapor but there is no way to quantify

this at present. After dividing through, one obtains the following expression.
4. EFFECT OF SUPPRESSANTS ON METAL HRES 115

RHRIRHR' = Y^^^^ / (7^^^ . (HJHJW(nJnWJ (5)

Now it is possible to see that the effect of an add-on mass fraction of amount is modulated not

only by the relative heat of reaction but also by the relative rate at which these new molecules can

diffuse to the reaction zone and by the relative mass which must diffuse there (dictated by the reaction

stoichiometry). The diffusion coefficients differ because the agent is typically a much bigger molecule

than oxygen. Using diffusivity data for naphthalene and anthracene as molecules of appropriate

weight, and a mid-temperature estimate in the boundary layer of 1500 °C, one finds that {kjk^ is about 1/3. For halon 1301 versus oxygen, {njn^ is 0.216; for HFC- 125 it is 0.335 (both based on the

type of overall reaction indicated below).

An estimate of requires that some estimate of the products be made for the interaction with the

agent vapor. Lacking the requisite thermodynamic data to make detailed equilibrium calculations for

this system, we assume the following plausible overall reaction scheme for halon 1301.

4 77 + 4 O2 + 4 CF^Br - 3 TiF^ + TiBr^ + 4 COj

The evolved heat from this reaction is 1.06 MJ/g-mole of titanium, 16 % higher than that for titanium

dioxide formation. A comparable calculation for HFC- 125 yields an estimate that is about 50 %

greater than for Ti02 formation. (Necessary data on the heat of formation for HFC-227 could not be

found so an estimate for this agent has not been made.) Note that the above overall reaction implies

that there is some competition of the agent vapor for the oxygen that is diffusing to the metal surface,

since the carbon in the agent is oxidized in the gas phase. Thus there is not a clean separation into

two purely parallel, non-interacting paths of attack by the two oxidizing agents, as assumed in Equation 2. It seems likely that some of the carbon dioxide formed in the boundary layer would lose

its oxygen to the titanium when it diffuses to the surface; this would somewhat lower the overall

exotherm of the above reaction scheme and counter the effects of the neglected interaction. Any

dissociation of the products would also lower the available heat. With these caveats, one can use the preceding reaction heats in Equation 5 to estimate that the

heat release rate effect for a 10 % add-on of halon 1301 is a net decrease of about 6 For a

10 % add-on of HFC- 125, the estimated decrease in RHR is about 2 %. The estimated changes in mass burning rate are comparable. Given the various assumptions that went into these estimates they

are probably accurate to no better than ± 50 %. (Furthermore, the model predictions cannot be any more accurate than the underlying assumptions allow them to be.) In any event, for the add-on levels of the order of 10 % they do predict a small slow-down of the burning process. The reason that this prediction runs counter to the heat release enhancement one might expect is implicit in the main assumption behind the simplified model. Oxidant mass transfer is taken to be the

rate controlling step. The large agent molecules cannot diffuse to the reaction zone at more than a

fraction of the rate at which oxygen can. The displacement of some oxygen by the agent vapor

reduces the titanium dioxide production rate and heat release; the halogen reactions, retarded by the

A comparable calculation for magnesium and halon 1301 yields a 4 % decrease in RHR but one

should note that the model assumptions are less appropriate here due to the gas phase chemistry, rather

than heterogeneous chemistry.
116 4. EFFECT OF SUPPRESSANTS ON METAL FIRES

relatively slow agent molecules, cannot proceed fast enough to make up the difference, in spite of their

energetic advantage.

This simple model can make a prediction of what add-on level of agent would be needed to yield a net increase in heat release rate. This just implies that the denominator exceeds the numerator in Equation 5. Interestingly, the model says that for halon 1301 there is no add-on level of this agent

which can achieve a greater heat release rate than one gets with oxygen as the sole oxidant gas; this holds for oxygen levels from that in air up to at least 50 For HFC- 125 the model says that this agent will yield a RHR enhancement at add-ons above about 40 % if the starting gas is air; there is no

add-on level capable of yielding enhancement when the starting gas has 40 % or 50 % oxygen.

The above estimates of the impact of agent addition stand in contrast to the actual observed

behavior as shown in Table 3. As with magnesium, the actual effect of the agent vapors in all cases was strongly suppressive, not just slightly so. The reasons behind this are difficult to discern, given

the limited data. The strong effect suggests a greater role for chemistry than the above model allows.

Perhaps the oxidation reactions are more marginal (and thus more controlling) than has been assumed

or there are suppressive chemical effects of the agent on the metal burning which are not known to the

present authors.

In titanium burning, nitrogen is somewhat more akin to an inert diluent than it was with

magnesium because the heat of formation of titanium nitride is one third that of titanium dioxide.

Table 3 shows that a 10 % nitrogen add-on yields a 17 % reduction in burning rate (in 40 % oxygen

at 0.79 MPa). A 10 % add-on of halon 1301 (with only about half of the thermal capacitance of the

nitrogen add-on) yields either a much greater burning rate reduction or even extinction.^^ A 10 %

add-on of HFC- 125 gave very erratic results under these same conditions, from no effect to extinction.

(The pressure variation in these tests was the result of partial clogging of the downstream orifice by

particles emitted from the burning titanium.) In terms of thermal capacitance or ability to lower

temperature by pure dilution, a 10 % add-on of nitrogen is comparable to HFC- 125 and about twice as

effective as halon 1301. These results thus suggest that the apparent suppressive effects of the agents

are beyond simple thermal dilution and well beyond the calculated effects if mass transfer was the

controlling mechanism.

The ambient oxygen level was increased to 50 % so that the chamber pressure could be lowered

to the point where HFC-227, with its comparatively low vapor pressure, could be tested alongside

HFC- 125 and halon 1301. Table 3 shows that the more vigorous burning that resulted from the

oxygen enhancement resisted extinguishment by any of these agents. However, they again all slowed

the burning process strongly. Here halon 1301 and HFC- 125 appear about equally effective in slowing

the burning. HFC-227 is distinctly less effective, though this inference is based on very limited data. One might cautiously extrapolate this result to infer that HFC- 125 would produce a lesser hazard in aircraft applications than would HFC-227 but more data on this issue would certainly be desirable.

Again it is necessary to add the caveat noted previously with the magnesium results. The ability

of these agent vapors to partially suppress the burning of titanium in the present experimental

conditions does not mean that they will do so in a more complex burning configuration or at all

For magnesium and halon 1301, this calculation says that agent add-on levels to air of nearly 80% are needed to yield an increase in heat release rate above that found in pure air.

'^The extinction process here occurred after 50-80 % of the rod had burned and thus was much more

marginal than in the case of magnesium. It occurred when the bum zone was in the contracted molten

ball stage. A build-up of titanium dioxide near the ball surface (Clark et ai, 1975) could make this a slower reaction stage and thus a more vulnerable part of the burning cycle.
4. EFFECT OF SUPPRESSANTS ON METAL FIRES 117

admixture levels. Also, any agent sprayed at a metal fire in liquid form has the potential to cause a

flare-up due solely to droplet "explosions" which could enlarge the metal area exposed to oxidation.

4.4 Conclusions

Halon 1301 (the reference case), as well as the candidate replacement agents HFC-125, HFC-227,

FC-218 and CF3I, extinguished the strongly- established burning of thin rods of magnesium when

added-on at the 5 % and 10 % level to pressurized air flowing past the bum zone. In nearly all cases

the extinction occurred in a matter of seconds. FC-218 was appreciably slower in causing extinction

in two replicate tests at one condition; this was the only differentiation among the various agents seen

with burning magnesium. In this sense the candidate replacement agents, other than FC-218, have

shown no greater tendency to interact with burning magnesium than does halon 1301.

The general suppressive effect on magnesium burning runs counter to the burning enhancement

effects reported in the literature for halogen-containing compounds sprayed onto more complex

magnesium burning configurations. A simplified calculation, based on the assumption of mass transfer

control of the burning rate and an assumed reaction stoichiometry, suggests that a very weak suppressive effect could be expected here, largely as a result of oxygen dilution. The cause of the much

stronger suppressive effects seen here is not clear at present. This same calculation suggests that enhancement of magnesium burning could occur at much higher agent add-on levels. Halon 1301, HFC-125 and HFC-227 also showed substantial suppressive capabilities toward the

burning of a thin titanium rod in elevated oxygen atmospheres. Here extinction was seen only in 40 % oxygen, not in 50 % oxygen. Nevertheless, even at the higher oxygen level, substantial

reductions in bum propagation rate were seen. Here HFC-125 at a 15 % add-on level slowed the bum

rate equally as much as did halon 1301 and much more than did HFC-227. This was the only

differentiation among agents seen with titanium buming. It provides a very limited basis on which to infer that HFC-125 might be less reactive with titanium fires than HFC-227 (and comparable to halon 1301).

The simplified calculation, mentioned above but here done only for halon 1301 and HFC-125,

again implies only a very weak suppressive effect of 10 to 15 % add-on levels of these agents, via

oxygen dilution. The much stronger effect seen experimentally cannot be explained on the basis of the

available information. This calculation implies that, at elevated oxygen levels, no add-on level of

agent can yield buming enhancement. In air, according to this model, HFC-125 could yield en- hancement of the buming rate at add-on levels above 40 %. This raises a note of caution about the

relatively lower reactivity of HFC-125 compared to HFC-227 due to the limited number of hard facts available.

Since the results here for isolated, single metal rods mn counter to the reported effects of

halogen-containing agents on more massive magnesium and titanium fires, it appears that further

experiments would be desirable, involving larger masses of buming metal, preferably in a linear

bummg mode as here, but incorporating radiative interchange among fuel elements. This would allow

quantitation of the buming progress, allowing definitive determination of the impact of agent vapors

and agent liquid/vapor sprays. Pre-heating and pressurization of the gas stream would also make it possible to adhere to oxygen levels closer to air with titanium so that one could determine more

definitively the impact of halogen-containing agents under conditions most pertinent to jet aircraft

fires.
118 4. EFFECT OF SUPPRESSANTS ON METAL FIRES

4.5 References

Anderson, V. and Manty, B., "Titanium Alloy Ignition and Combustion," Naval Air Development

Center Report No. NADC 76083-30, Warminster, PA, January, 1978.

Andrews, R., "Metals," Section 4, National Fire Protection Handbook, Fifteenth Edition, Chap.

10, Section 4, National Fire Protection Association, Quincy, MA, p. 4-98, (1981).

Baldwin, J., Beach, C. and Rhein, R., "Extinction of Titanium Fires in Pressurized Flowing Air,"

Naval Weapons Center Report No. NWC TP 6439, China Lake, CA, Sept., 1984.

Bird, B. Stewart, W. and Lightfoot, E., Transport Phenomena, Wiley, New York, p. 647 (1960).

Clark, A., Moulder, J. and Runyan, C, "Combustion of Bulk Titanium in Oxygen," Fifteenth

Symposium (International) on Combustion, The Combustion Institute, Pittsburgh, PA, p. 489 (1975)

Elrod, C, Funkhauser, M. and Glickstein, M., "Compendium on Titanium Combustion," Air

Force Wright Aeronautical Laboratories Report No. AFWAL-TR-82-2014, Wright-Patterson Air Force

Base, OH, Nov., 1980.

Elrod, C. and Lyon, S., "Self-Sustained Combustion Evaluation of Titanium 6A1-4V in an Aerodynamic Environment," Air Force Wright Aeronautical Laboratories Report No. AFAPL-TR-79-

2047, Wright-Patterson Air Force Base, OH, August, 1976.

Fox, D., "Investigation of Titanium Combustion Characteristics and Suppression Techniques," Air

Force Wright Aeronautical Laboratories Report No. AFAPL-TR-75-73, Wright-Patterson Air Force

Base, OH, July, 1975.

Hamins, A., Gmurczyk, G., Grosshandler, W., Rehwoldt, G., Vazquez, I., Cleary, T., Presser,

C. and Seshadri, K., "Flame Suppression Effectiveness," Chapter 4 in "Evaluation of Alternative Fire

Suppressants for Full-Scale Testing in Simulated Aircraft Engine Nacelles and Dry Bays," National

Institute of Standards and Technology Special Publication 861, April, 1994.

Kuchta, J. Martindill, G. and Spolan, I., "Fire and Explosion Hazards of Flight Vehicle Combustibles," Quarterly Progress, U. S. Bureau od Mines, Pittsburgh, PA, June to August, 1965.

ibid, November to February, 1966.

Law, C. and Williams, F, "Combustion of Magnesium Particles in Oxygen-Inert Atmospheres,"

Combustion and Flame, 22, p. 383 (1974).

Maykuth, D., "Methods of Controlling and Extinguishing Titanium Fires," Defense Metals

Information Institute Technical Note, Batelle Memorial Institute, Columbus, OH, July, 1964.

McCutchan, R., "Investigation of Magnesium Fire-Extinguishing Agents," Wright Air Development Center Technical Report 54-5, Wright-Patterson Air Force Base, OH, Jan., 1954.

National Aeronautics and Space Administration, "Detonation Hazards with 'Safe' Industrial

Solvents," NASA Langley Research Center LAR-10299, Hampton, VA, 1970.

Nyden, M., Linteris, G,. Burgess, D., Westmoreland, P., Tsang, W. and Zachariah. M., "Flame

Inhibition Chemistry and the Search for Additional Fire Fighting Chemicals," Chapter 5 in "Evaluation

of Alternative Fire Suppressants for Full-Scale Testing in Simulated Aircraft Engine Nacelles and Dry

Bays," National Institute of Standards and Technology Special Publication 861, April, 1994.

Perlee, H., Martindill, G. and Zabetakis, M., "Flammability Characteristics of Selected Halogenated Hydrocarbons," U. S. Bureau of Mines TN23.U7, Pittsburgh, PA, 1966.

Rhein, R. and Baldwin, J., "Literature Review on Titanium Combustion and Extinction," Naval

Weapons Center Report No. NWC TP 6167, China Lake, CA, Jan., 1980.

Sato, J. and Hirano, T, "Fire Spread Limits Along Metal Pieces in Oxygen," Flammability and

Sensitivity of Materials in Oxygen-Enriched Environments: Third Volume, ASTM STP 986, D. Scroll

(ed.), American Society for Testing and Materials, Philadelphia, PA, p. 158 (1988).
4. EFFECT OF SUPPRESSANTS ON METAL FIRES 119

Sato, J., Sato, K. and Hirano, T., "Fire Spread Mechanisms along Steel Cylinders in High

Pressure Oxygen," Combustion and Flame, 51, p. 279 (1983).

Strobridge, T., Moulder, J. and Clark, A., "Titanium Combustion in Turbine Engines," Federal

Aviation Administration Report No.FAA-RD-79-51, Washington, DC, July, 1979.

Tapscott, R., Beeson, H., Lee, M., Plugge, M., Zallen, D., Walker, J, and Campbell, P., "Extinguishing Agent for Magnesium Fire: Phases I-IV," New Mexico Engineering Research Institute Report

WA3-15, Albuquerque, NM, July, 1986.

121

5. COMPATIBILITY WITH METALS

Mark R. Stoudt, James L. Fink, James F. Dante and Richard E. Ricker

Metallurgy Division

Materials Science and Engineering Laboratory

Contents

Page

5. COMPATIBILITY WITH METALS 121

5.1 Introduction 122

5.2 Mass Change Exposure Tests 123

5.2.1 Materials 123

5.2.2 Experimental Procedure at 150 °C 123

5.2.3 Experimental Procedure at 20 °C 127

5.2.4 Results at 150 °C 127

5.2.5 Results at 20 °C 139

5.3 Environmentally Induced Fracture Experiments (Slow Strain Rate Tensile Tests) 152

5.3.1 Materials 152

5.3.2 Experimental Procedure at 150 °C 153

5.3.3 Experimental Procedure at 20 °C 153

5.3.4 Results at 150 °C 153

5.3.5 Results at 20 °C 161

5.4 Post Deployment Corrosion 168

5.4.1 Materials 172

5.4.2 Exposure Environment 172

5.4.3 Experimental Procedure 172

5.4.4 Results 172

5.5 Electrochemical Measurements 176

5.5.1 Materials 177

5.5.2 Experimental Procedure 177

5.5.3 Agent Resistivity 179

5.5.4 Defining a Reference Potential 180

5.5.5 Cp2^e^ Electron Kinetics 183

5.5.6 Corrosion Rates of Storage Vessel Alloys in HCFC-124, HFC- 125, and CF3I 185

5.5.7 Corrosion Rates of Metals in Other Halon Replacement Agents 189

5.6 Conclusions 190

5.7 Acknowledgments 191

5.8 References 191

Appendix A. Mass Change Measurements at 150 °C 193

Appendix B. Mass Change Measurements at 20 °C 197
122 5. COMPATIBILITY WITH METALS

5.1 Introduction

Under current military practice, fire suppressants are stored for up to five years. Since the corrosivities of the chemicals currently under consideration as replacements for halon 1301 are not well known,

there is a need for compatibility information between these chemicals and the array of materials used on high performance aircraft. As a result, experiments were conducted to assess the relative compatibility of the potential fire suppressants with the different metallic materials used for storage, distribution, and handling of aircraft fire suppressants.

On board an aircraft, the metallic materials will be exposed to fire suppressants during storage

(storage vessels, rupture disks, etc.), deployment (distribution tubes, nozzles, etc.) and after deployment

(structural and engine components exposed to combustion by-products). Of these possible exposures,

the greatest concern for the safe operation of the aircraft lies in the possibility that the fire suppressant

might cause attack of the storage vessels resulting in a loss of agent or prevention of proper operation

of the deployment system.

In flight, the storage vessels, which are designed to contain the agent for up to five years, will

experience pressure excursions up to 5.9 MPa at operational temperatures which will range from below

ambient to those as high as 150 °C. At the maximum in-flight temperature, it is likely that the agent

will surpass the critical temperature and will, therefore, be a single phase. At the lower in-flight and

ground temperatures, the agent will be an equilibrium mixture of the liquid and vapor phases. Since

corrosion behavior is a function of many variables and is also strongly dependent on temperature, it follows that the corrosion behavior at the ground conditions are likely to be very different from those

in flight (Fontana, 1987).

Corrosion is defined as a degradation of the properties of a material resulting from a chemical

reaction with the environment and can generally be classified into eight different categories based on

the morphology of the corrosion damage. These eight forms are: general corrosion, pitting corrosion,

crevice corrosion, intergranular corrosion, environmentally induced fracture, dealloying, galvanic

corrosion, and erosion corrosion (Fontana, 1987). The definitions and descriptions of these forms of

corrosion are detailed in an earlier report (Ricker, 1994). Of these eight corrosion failure modes, six are of potential concern for the storage, distribution, and post deployment corrosivity of fire suppressant agents on aircraft: general corrosion, pitting corrosion, crevice corrosion, intergranular corrosion,

environmentally induced fracture, and dealloying (Ricker, 1994). Erosion corrosion may be a concern

for nozzles, but the duration of the flow through the nozzle is very short and damage by this mechanism should be noted in the fire suppressant efficiency tests if it occurs. Galvanic coupling may

influence corrosion damage accumulation, but due to the relatively low conductivity of these agents,

galvanic effects should be minimal. As a result, the experiments were designed in a manner that would allow for the evaluation of the occurrence of the remaining six modes of damage (Ricker,

1994).

Very little is known about the corrosion behavior of the metals in these environments, and there is virtually no information available in the literature regarding the compatibility of these agents with the

alloys used in the storage systems. As a result, corrosion experiments, which emulate these storage

conditions, are a vital component in the selection of a replacement to halon 1301.

In the first phase of this investigation, a series of twelve possible replacements for halon 1301 were evaluated on the basis of a general ranking of the corrosivity with respect to storage vessel

materials (Ricker, 1994). For the second phase, three candidates from the original list of twelve were

selected for further study. These were: HFC- 125, HFC-227ea, FC-218. An additional candidate, CF3I, was also included in this corrosion study. The primary objective of the Phase 11 analysis was to assess the compatibility of this select group of replacement candidates over a broader range of environmental
5. COMPATIBILITY WITH METALS 123

conditions and longer exposure times. Since limited corrosion data is available for the selected alloys

in 1301 in the literature, and the service performance record of this agent has been acceptable, this agent was included in Phase II as a reference environment.

The potential for corrosion damage to the structural materials of the aircraft by residual suppressant or by combustion by-products is also of great concern. These chemicals have been shown to form aggressive halide acids or salts during fire suppression (Nyden, 1994). Residual suppressant or

the combustion by-products left on the surface of the structural components after deployment could

result in extensive corrosion damage before they are cleaned.

This analysis was designed and conducted to further evaluate the compatibility of the replacement

candidates selected for additional study under probable and potential service conditions. The

experiments consisted of exposure tests which emulated the in-flight and on-the-ground conditions

under which corrosion might occur followed by careful examination and evaluation of the resulting

corrosion damage. Standard experimental techniques were employed to assess the potential for failure of the storage container alloys by the aforementioned six possible corrosion mechanisms in normal

service. The techniques used for this evaluation consisted of immersion tests and slow strain rate

tensile tests and were similar to those utilized in Phase I (Ricker, 1994). A new electrochemical

technique was also developed to predict the corrosion behavior over a broader range of potential

service conditions and incorporated into the analysis.

5.2 Mass Change Exposure Tests

Experiments were conducted to assess the changes in mass of container alloys as a result of a long

term exposure to the replacements candidates. These measurements provided a basis to determine the

rate of formation of surface scales or the rate of removal of metallic species due to the exposure to the

replacement candidate. Also visual and optical microscopic examinations following the exposures

allowed for further evaluation of the occurrence of pitting, intergranular corrosion, and dealloying.

5.2.1 Materials. The materials selected by the sponsors for evaluation in Phase n of this study

represented the range of alloys currently in service as or under consideration for use as storage vessels. These were: 304 stainless steel, stainless steel alloy 21-6-9 (Nitronic 40)\ titanium alloy 15-3-3-3 and

6061-T6 aluminum alloy. Shortly after the immersion tests began, a decision was made to include

AISI 4130 alloy steel in the test matrix and the exposure times for this alloy were shortened appropriately. After HFC- 125 was selected as the most likely candidate to replace halon 1301, several

additional alloys were included for further testing in HFC-125 and halon 1301 in order to provide

compatibility information over a broader range of m.aterials. Those materials were: Inconel alloy 625,

CDA-172 copper/beryUium alloy, CDA-110 copper alloy, 321 stainless steel, and AM-355 stainless

steel. The compositions of these alloys are given in Table 1

.

5.2.2 Experimental Procedure at 150 °C. All of the samples used for the weight loss immersion

tests were flat coupons of the geometry shown in Figure 1 (ASTM G 01-93). The sample preparation

Certain trade names and company products are mentioned in the text or identified in an illustration in order to specify adequately the experimental procedure and equipment used. In no case does such

identification imply recommendation or endorsement by the National Institute of Standards and

Technology, nor does it imply that the products are necessarily the best available for the purpose.
124 5. COMPATIBILITY WITH METALS

W = 25.4 mm (1.0 in) L = 50.8 mm (2.0 in)

t = 1 .6 mm (0.0625 in) d = 6.7 mm (0.265 In)

General corrosion coupon design (ASTM G-01).

Figure 1. General corrosion coupon sample design (ASTM G 01-93).
5. COMPATIBILITY WITH METALS 125

Table 1. Composition of alloys tested in Phase II (mass fraction in percent)

Element Nitronic 304 Al Titanium AISI Inconel AM-355 Cu-Be 321 Cu Alloy 40 SS 6061 15-3-3-3 4130 625 SS CDA 172 SS CDA 110

Al -- -- bal 3.50 0.04 0.23 - 0.04 - - C 0.02 0.06 -- 0.01 0.30 0.02 0.12 - 0.04 -

Cr 20.20 18.39 0.34 3.30 0.91 21.71 15.28 0.01 17.22 -

Cu 0.03 0.27 0.40 - 0.01 - - bal 0.10 99.95

Fe bal bal 0.70 0.11 bal 3.97 bal 0.06 bal - Mn 9.07 1.41 0.15 - 0.48 0.08 0.80 " 1.61 - Mo 0.07 0.21 - - 0.17 8.82 2.60 - 0.14 -

Ni 7.01 8.15 -- - 0.01 61.39 4.23 0.06 9.85 -

Ti -- -- -- bal - - ~ - 0.43 -- V -- -- -- 14.70 - - - - - ~ Mg -- -- 1.20 - - - - - - -

Zn - - 0.25 - - - - - -

B - - -- - - - - - - N 0.31 0.03 0.01 0.12 0.02

Sn 3.00 Co 0.14 0.21 Nb 3.41 W

Be 1.94

technique used for these experiments was identical to that used for Phase I (Ricker, 1994). Three

separate weighings were taken at approximately 30 s intervals and averaged. This average value was

then referred to as the initial weight of the coupon. The balance used for these weight measurements was self-calibrating to maintain an accuracy to within ± 10 x 10'^ g with a reproducibility of no less than ± 15.1 X

10'^

g. Representative photographs of the surface of each alloy were also taken prior to the start of testing.

Prior to the start of each exposure test, three coupons of each alloy were mounted on a

polytetrafluoroethylene (PTFE) rod with PTFE spacers between the samples. This was done (1) to separate and (2) to electrically isolate the samples so that the effects from galvanic couplings would be

minimized. PTFE shields were also placed between the individual sets of alloys to protect them from

contacting any corrosion products that may have formed on other alloy sets. The samples were next

placed in a custom designed PTFE liner and placed into a 2 L pressure vessel, as shown in Figure 2, for exposure. The test vessel was sealed, connected to a mechanical vacuum pump and evacuated for a minimum of 45 min.

The mass of agent required to produce a nominal pressure of 5.9 MPa at 150 °C was determined

with the same mass balance used for the elevated temperature tests in Phase I which is shown in the

following equation: (Ricker, 1994)
5. COMPATIBILITY WITH METALS

General Corrosion General Corrosion

at150°C at20°C

Liquid Phase

1 = AI Alloy AA 6061 -T6

2 = Tltanuim 15-3-3-3

3 = 304 Stainless Steel

4 = Nitronic 40

5 = AISI 4130 Steel

;ure 2. Schematic diagram of the immersion testing chamber used for general corrosion

experiments.
5. COMPATIBILITY WITH METALS 127

where Mj is the total mass of the system, pj is the density of the liquid agent at 20 °C, V, is the

volume of the liquid agent at 20 °C, is the density of the vapor phase at 20 °C, and is the

volume of the vapor at 20 °C. To facilitate charging of the agents, the test vessels were chilled in a bath of either ice and water, or dry ice and alcohol, depending on the temperature necessary to maintain the liquid phase. The vessels were then placed on a balance so that the mass of the agent

could be monitored during filling. Upon completion of the charging step, the vessels were placed in proportionally controlled, calrod-type heaters for the duration of the exposure period.

Mass-change measurements were performed on the coupons after one, two, four, eight, and twelve

month elapsed times during the one year exposure. To minimize agent loss during the extraction

process, the vessels were immersed in a bath of dry ice and alcohol for approximately three hours in order to condense as much of the agent as possible. The vessels were immediately sealed and returned

to the cooling bath upon the extraction of the PTFE rod assemblies. After extraction, the samples were allowed to warm naturally to ambient temperature, promptly weighed using the standard

weighing procedure, and then immediately returned to the test vessels. At the end of the exposure tests, the agents were released, the coupons were extracted and

immediately re-weighed, again using the standard weighing procedure. The average of the three

weight measurements was then referred to as the final weight of the coupon. Representative

photographs of the surfaces were again taken and compared to those of the initial condition.

A series of statistical analyses were performed on the weight loss coupon data. The results of

these analyses were then used to evaluate the potential for general corrosion induced failure for each

alloy in the replacement candidates.

5.2.3 Experimental Procedure at 20 °C. The sample charging procedure used for the ambient

temperature experiments differed slightly from the one utilized for the tests at 150 °C. For these tests, the samples were not separated according to alloy type, but were stacked so that each set contained one coupon from each alloy. This sample arrangement was selected because it exposed two coupons

from each alloy to the liquid phase, and one to the vapor phase as also shown in Figure 2. As with

the tests conducted at 150 °C, each coupon was separated by a PTFE spacer and each alloy set was

separated by a PTFE shield. The agent filling procedure used for the ambient temperature tests was also slightly different from

the high temperature test. For these tests. Equation (1) was used to determine an agent volume which

located the interface region between the two phases near the midpoint of the vessel. The remainder of

the procedure for these experiments was identical to the procedure used for the tests at 150 °C.

5.2.4 Results at 150 °C. The rate of mass change was calculated for each alloy/environment

combination using the mass change measurements and the same relationship from Phase I (Ricker,

1994)

(2)
128 5. COMPATIBILITY WITH METALS

where aM is the mass change, A is the total exposed area of the sample and t is the exposure time in days. These values are presented in Table 2. If the mass change rate is assumed to be essentially constant throughout the entire exposure period, then this relationship will provide a reasonable estimate

of the average mass change rate for the duration of the exposure. While the magnitudes of these

measurements are small they are larger than the minimum required to be statistically significant.

Linear regression analyses were also performed on the mass change rate values. Table 3 contains

the results of a regression performed on mass change per unit area versus time basis {i.e., aM/A vs t). In this table, the slope of the linear regression, m, corresponds to the estimated corrosion rate for each

alloy/environment combination, R is the correlation coefficient for the linear regression, and the

standard error is an estimate of the standard deviation of the estimated slope. In this table, it can be

seen that the correlation coefficients reflect the high degree of scatter in the measurements. At this temperature, 4130 alloy steel in halon 1301 had the best correlation (R value > 0.9) and the 4130 alloy

steel in HFC- 125 was the worst (R < 0.1).

If the corrosion rates were indeed constant, then the weight changes would be linear and the

regression slope would then equal the mass change rate for each exposure. However, in most cases

the corrosion rate slows with time due to mass transport through the surface film thereby limiting the

reaction rate. In this case, the corrosion rate should decrease with time following a square root of time

dependence. As a result, a second regression was performed on the mass change data on a square root of time basis. If one assumes that transport through a surface film is rate limiting, then the rate of mass change would be inversely proportional to the thickness of the film according to the following

relation:

where the film thickness, C, is proportional to the mass change per unit area. Solving this differential

equation yields the equation used for the second regression:

The results of this regression are presented in Table 4. The correlation coefficients for this analysis

also exhibited a wide scatter ranging from > 0.9 for Al-6061 in HFC-125 to -0.1 for 4130 in the same

agent. While the correlations of this regression were better in some cases, the overall fit was not much better than the simple linear regression. As a result, a third regression was also performed on

this data. This analysis assumes that part of the samples surface is bare and exhibits linear reaction

kinetics while part is filmed and exhibits square root of time behavior so that:

This regression produced very good correlations for most of the environment/alloy combinations as shown in Table 5 and are also shown along with the raw mass change data as derived from the

measurements made on the three samples of each of the alloys in each of the agents at 150 °C,

(Figures 3-7). (The raw data used for these figures is presented in tabular form in Appendix A.)

d(AM\ 1 a — (3)

dt\ A H

(4)

(5)
5. COMPATIBILITY WITH METALS 129

Table 3. Mass change rate per unit area at 150 °C (g/meter^ • day) estimated by the slope for a

linear regression of mass change per unit area on exposure time (assumes constant, bare

surface, reaction kinetics)

HFC- 125 HFC-227ea Halon 1301

Alloy m R Std Err m R Std Err m R Std Err

Nit 40 n 00076 0 S09 0 00036 n 777

Al 6061-T6 0.00233 0.800 0.00044 0.00474 0.861 0.00078 0.00237 0.681 0.00064

304 SS 0.00039 0.692 0.00010 0.00072 0.608 . 0.00026 0.00066 0.693 0.00017

Ti 15-3-3-3 0.00029 0.651 0.00009 0.00063 0.333 0.00049 0.00096 0.768 0.00020

AISI 4130 -0.00037 0.035 0.00407 -0.08934 0.413 0.07450 0.01737 0.937 0.00245

321 SS + t t * * *

X X CDA 110 i ± t

* *

X X X CDA 172 +

X

+ +

* * X X X AM 355 J.+ t t

* * X X X

IN 625 ± t

* * X X X

Table 3. (continued)

Alloy

FC-218 CF3I m R Std Err m R Std Err

Nit 40 0.02284 0.855 0.00384 0.07057 0.435 0.04622

Al 6061-T6 0.01996 0.454 0.01085 0.05121 0.864 0.00944

304 SS 0.00140 0.848 0.00024 0.19160 0.887 0.03156

Ti 15-3-3-3 0.00578 0,792 0.00124 0.00973 0.410 0.00819

AISI 4130 0.03982 0.804 0.01112 -0.45896 0.792 0.13382

321 SS * * * * * * CDA 110 * * * * * CDA 172 * * * * AM 355 * * * *

IN 625 * * * * * *

* No measurements were made in this alloy/environment combination

X Only one exposure time used for these measurements, see the average mass change rate table m = estimated mass change rate for linear regression of mass change per unit area on exposure time (g/meter^ • day)

R = correlation coefficient for regression Std Error = estimated standard deviation for the mass change rate (regression slope)
130 5. COMPATIBILITY WITH METALS

Table 4. Mass change rate per unit area at 150 °C (g/meter^ • day) estimated by the slope for a

linear regression of mass change per unit area on square root of the exposure time

(assumes a uniformly filmed surface)

Alloy

HFC- 125 HFC-227ea Halon 1301

k R Std Err k R Std Err R Std Err

Nit 40 0.00552 0.831 0.00092 0.01415 0.661 0.00446 0.01849 0.922 0.00194

Al 6061-T6 0.05442 0.923 0.00566 0.07080 0.922 0.00827 0.05913 0.839 0.00958

304 SS 0.00220 0.855 0.00147 0.01347 0.811 0.00269 0.00243 0.875 0.00232

Ti- 15-3-3-3 0.00802 0.710 0.00221 0.01466 0.555 0.00609 0.02338 0.922 0.00246

AISI 4130 0.01836 0.140 0.49160 -1.10539 0.419 0.90484 0.22373 0.991 0.01171

321 SS t t t

* * * t t t CDA 110 t t t

* * * t t t CDA 172 t i t

* * * i i t AM 355 t * t

* * * t t t IN 625 t t t

* * t t t

Table 4. (continued)

Alloy

FC-2!8 CF3I

k R Std Err k R Std Err

Nit 40 0.30768 0.818 0.06006 1.16426 0.657 0.42281

Al 6061-T6 0.26926 0.440 0.15261 0.53902 0.819 0.11951

304 SS 0.00450 0.783 0.00396 2.20300 0.918 0.30045

Ti- 15-3-3-3 0.08392 0.824 0.01598 0.13561 0.343 0.14058

AISI 4130 0.48701 0.807 0.13449 -6.09099 0.862 1.35119

321 SS * * * * * * CDA 110 * * * * * * CDA 172 * * * * * * AM 355 * * * * *

IN 625 * * * * *

* No measurements were made in this alloy/environment combination

t Only one exposure time used for these measurements, see the average mass change rate table k = estimated mass change rate for linear regression of mass change per unit or square root of the exposure time

(g/meter^ • day"^)

R = correlation coefficient for linear regression Std Error = estimated standard deviation for the mass change rate (regression slope)
5. COMPATIBILITY WITH METALS 13

Table 5. Mass change rates at 150 °C estimated from a regression model that assumes mixed

bare and filmed surface kinetics

HFC- 125

Alloy iVlQ M, Mj \A ( f*XTC\T\ IVl^V^tl 1 \J\ )

IN It 4U -U.UUZUl u.uuyso -U.UUUjO U.UUUl 1 u.uizsz U.UUZl V u.y 1

J

Al OUO 1-10 -U.UOZjj U.UOj'+l -u.uuzi J u.uwu / z

304 SS -0.00825 0.01503 -0.00065 0.00017 0.0220. 0.00334 0.931

Ti-15-3-3-3 -0.00179 0.01893 -0.00065 0.00021 0.02397 0.00420 0.920

AISI 4130 0.000 0.35385 -0.07341 0.03582 0.57720 0.26956 0.658

321 SS i t t t X X CDA 110 ± t t t X X CDA 172 i t + i X X AM 355 ± t t t t X X

IN 625 + t t t i X X

Table 5. (continued)

Halon 1301

Alloy Mo Mq(error) M, M, (error) M2 M2(error) R

Nit 40 -0.01110 0.01654 -0.00095 0.00018 0.03665 0.00367 0.974

Al 6061-T6 -0.06996 0.10502 -0.00390 0.00115 0.13374 0.02331 0.912

304 SS 0.00554 0.01651 -0.00126 0.00018 0.04079 0.00366 0.972

Ti-15-3-3-3 -0.01031 0.01885 -0.00129 0.00021 0.04813 0.00418 0.979

AISI 4130 1.15e-'° 0.09515 0.00234 0.00964 0.33181 0.07258 0.993

321 SS X X X X X X CDA 110 X X X X X X X CDA 172 X X X X X X AM 355 X

J. X X X t t

IN 625 X

J.+ X X X X X
132 5. COMPATIBILITY WITH METALS

Table 5. (continued)

FC-218

Alloy Mo0 Mq(error) M,1 M| (error) M,2 M2(error) R

Nit 40 -0.17990 0.56982 0.02352 0.01314 -0.00991 0.18321 0.855

Al 6061-T6 -0.14014 1 .61 140 0.01692 0.03715 0.04445 0 51810 0.455

304 SS 0.00294 0.03554 0.00182 0.00082 -0.00613 0.01143 0.852

Ti-15-3-3-3 -0.04398 0.17003 0.00050 0.00392 0.07724 0.05467 0.825

AlSl 4130 0.00001 1.25580 0.08739 0.12758 0.10738 0.96200 0.813

321 SS * * * * * * * CDA 110 * * * * * * * CDA 172 * * * * * * * AM 355 * * * *

IN 625 * * * * * * *

Table 5. (continued)

HFC-227ea

Alloy Mo Mq(error) M, M, (error) M2 M2(error) R

Nit 40 -0.00571 0.04056 -0.00197 0.00094 0.03995 0.01304 0.762

Al 6061-T6 0.00940 0.08720 -0.00099 0.00201 0.08400 0.02804 0.923

304 SS -0.00153 0.01084 -0.00212 0.00025 0.04168 0.00348 0.975

Ti-15-3-3-3 -0.00582 0.04084 -0.00402 0.00094 0.06807 0.01313 0.852

AISI 4130 0.000 8.55920 -0.15585 0.86606 -0.56139 6.52660 0.420

321 SS * * * * * CDA no * * * * * * * CDA 172 * * * * * * * AM 355 * * * * * * *

IN 625 * * * * * * *

Dl
5. COMPATIBILITY WITH METALS • 133

Table 5. (continued)

CF3I

iviQ^^crrui ^ M, MJ (error) M2(error) K

Nit 40 u.uu 1 1 y

1 1

1

-U.J5 IsJ A ACO A£ 4.9333 f\ C DA 1 0 0.58918 0.953

Al oUol - 1 0 A r\A "70 7 -U.Uo /a / u.o /yoo U.Ujz / 1 A A'3 0 t A U. (Jji 14 A A 1 TC7 -U.Ul /5 / 0.35687 0.864

f\ IOC/:.'? -U. 1 Zjo / A AT AAA A AA 1 -1 "7 u.uyi4/ 1 .8759 1.0157 0.919 A f\A 1 A 1 -U.U41Ui 0.22519 0.915

AISI 4130 -2.12el0 12.199 0.50302 1.2265 -13.27 9.2174 0.876

321 SS * * * * * * * CDA 110 * * * * * * + CDA 172 * * * * * * * AM 355 * * * * * *

IN 625 * * * * *

* No measurements were made in this alloy/environment combination

X Only one exposure time used for these measurements, see the average mass change rate table m^ = g/meter^ m, = g/meter^ • day

mj = g/meter^ • day"^

Figure 3 shows the mass change per unit exposed area for each sample in FC-218. The performances of the alloys in HFC-227ea, HFC- 125 and in halon 1301 are shown in Figures 4, 5, and 6, respectively. Examination of these figures revealed that the majority of the observed changes were mass

increases. Mass increases are not uncommon during immersion testing and are generally an indication

of the formation and growth of a scale on the surface of the samples. Visual examinations confirmed

the presence of such surface films; the samples were discolored in many cases. During immersion

tests in aqueous solutions, the scales that form are usually oxide and/or hydroxide films. In the case

of exposures to the replacement candidates, the films may be the result of reaction with the agent

residual gasses, decomposition products, or just deposits of decomposition products. The large

variations in mass exhibited in these figures could be due to cracking and/or spalling of these surface

scales during the exposure period or during sample handling necessary for the measurements.

Generally, the 4130 alloy steel exhibited the largest mass changes in every agent and the poorest

correlation coefficients. Again, these were usually mass increases; however, the 4130 steel did exhibit

significant mass decreases in the CF3I (Figure 7). These decreases in mass are believed to be the

result of active corrosion processes. At this temperature, the CF3I appeared to be substantially more aggressive than the other agents

evaluated. The magnitudes of the mass changes observed in this agent were larger as a result of the

formation of substantially thicker surface films. This hypothesis was supported by the results of the

visual inspections performed on the coupons. While visual examination found that films were also

present on the 4130 steel samples, they were layers of loose corrosion products. In general, the

magnitudes of the mass changes for the 4130 steel were consistently higher than the other alloys tested

in the same environments.
134 5. COMPATIBILITY WITH METALS

Figure 3. Mass change versus exposure time in FC-218 at 150 °C.
COMPATIBILITY WITH METALS 135

as

04

as

02

oi

•

" V'""i""l" ,'

1 » NIT~40O150C 1

1 •

•

• • 1

1.4 I"

I

ft 0.8 %

3 0.6 I

C 0.4

I

0.2

0

•

1 AA6061O150C \

• •

•

50 100 ISO 200 250 300 350 400

ExpoMjr* Tina, Day*

0.2S 1

f

02

0.16

I

OI

006

•

•

1 « SS304ei50C 1

f

•

50 100 150 200 250 300 350 400

Expo«ura Tlma, Day* iiiiiiiiiiiiiviiigi

SO 100 ISO 200 250 300 350 400

EjqMMra Tim*, Day*

50 100 150 200 250 300 350 400

Exposura Tim*, Day*

10

I

-10

-20

-30

...................

1 — AISI 4130ei50C 1

•

•

•

0 so 100 ISO 200 250 300 350 400

Expoaw* TInM, Day*

Figure 4. Mass change versus exposure time in HFC-227ea at 150 °C.
Figure 5. Mass change versus exposure time in HFC- 125 at 150 °C.
COMPATIBILITY WITH METALS 137

Figure 6. Mass change versus exposure time in halon 1301 at 150 °C.
Figure 7. Mass change versus exposure time in CF3I at 150 °C.
5. COMPATIBILITY WITH METALS 139

As with the Phase I results, the evaluation of the relative performances of the alloy/agent

combinations was hampered by the comparison of mass loss and mass gain measurements in the same

environment. Typically, mass loss measurements are evaluated by assuming that all of the lost mass

was the result of corrosion and that no corrosion products were left on the surface to generate errors in

this determination. Then, the quantity of metal reacting and rate of the reactions can be calculated

directly from the mass loss. Similarly, mass gain measurements can be evaluated if the reaction and

the reaction product stoichiometry are known and ii is assumed that none of the scale spalls off on the

samples weighed before and after scale removal. For this study, descaling techniques were examined,

but it was found that these techniques removed significant quantities of metal with respect to the

magnitude of these measurements or did not remove the surface scales. For analysis of the relative

corrosion behavior of the candidates, it is assumed that the relative magnitudes of the mass changes

are proportional to the corrosion rates of the different alloy/agent combinations. It should be noted

that the electrochemical measurements are the only means available for checking this assumption.

Visual inspections were also conducted on immersion test coupons. Those results are presented in Table 6. The attack observed on these coupons was relatively minor in nature and mainly consisted of

staining and/or discolorations, but some pitting was observed. The observed staining may be an

indication of general corrosion but this form of corrosion could not be distinguished from these

measurements. CF3I, again, appeared to be the most aggressive of the agents at 150 °C since pitting was observed in every alloy except for Nitronic 40 stainless steel. The 4130 alloy steel exhibited some slight pitting in every environment except for in the CF3I where the most severe pitting was

observed.

5.2.5 Results at 20 °C. The mass change rate, estimated for each alloy/environment combination

using the same relationships from the 150 °C analyses, are presented in Table 7. As expected, the

magnitudes of these values are much smaller at ambient temperature than at 150 °C. Similar linear

regressions were also performed on these data. Table 8 shows the results of the mass versus simple

time regression where m, the slope, is the estimated mass change rate, and R is the correlation

coefficient. The standard error, again is the estimated standard deviation of the mass change rate values. The correlations at this temperature were somewhat better than those at 150 °C, but once

again, the scatter resulting from the loss of surface films had a large influence. The worst value

f< 0.1) occurred in the Al-6061 in CF3I. The second regression, (aM/A = k(t)''^) produced slightly better results (Table 9), but once again, the correlations were not encouraging. As in the case with the

150 °C data, the correlation coefficients of this regression demonstrate that the mass changes at 20 °C

1 /9 also followed a combined linear and t dependence. The results of this regression are shown in Table 10. The mass change data from the experiments performed at ambient temperatures are presented in similar fashion to those performed at 150 °C and the corresponding curves from the multiple

regression are once again included with the data. The raw data used for these analyses are presented

in Appendix B. Figures 8-12 are the mass change per unit exposed area for each alloy in FC-218,

HFC-227ea, HFC-125, halon 1301 and CF3I respectively. The magnitudes of the mass changes

observed at ambient temperature were significantly smaller than those observed at 150 °C. These

large variations are particularly evident in the aluminum data. The CF3I did not appear to be as aggressive an environment at ambient temperature as it was at 150 °C. At this temperature, the mass

changes appear to be due to the formation, growth and spalling of surface scales. Visual inspections were also conducted on the coupons tested at 20 °C (Table 11). As in the case of the evaluation performed on the coupons tested at 150 °C, the attack observed on these

coupons was relatively minor in nature and mainly consisted of staining. Some pitting was observed

but, it was generally less severe at 20 °C. While the 6061 -T6 aluminum alloy did not exhibit any
140 5. COMPATIBILITY WITH METALS

Figure 8. Mass change versus exposure time in FC-218 at 20 °C.
COMPATIBILITY WITH METALS 141

Figure 9. Mass change versus exposure time in HFC-227ea at 20 °C.
142 5. COMPATIBILITY WITH METALS

Figure 10. Mass change versus exposure time in HFC- 125 at 20 °C.
COMPATIBILITY WITH METALS 143

Figure 11. Mass change versus exposure time in halon 1301 at 20 °C.
144 5. COMPATIBILITY WITH METALS

Figure 12. Mass change versus exposure time in CF3I at 20 °C.
5. COMPATIBILITY WITH METALS 145

Table 6. Results of visual inspection of coupons tested at 150 °C

Alloy HFC- 125 HFC-227ea Halon 1301 FC-218 CF,I

Nit 40 0 0 0 0 0

Al 6061-T6 0 0 0 0 1

304 SS 0 0 0 0 2

Ti 15-3-3-3 0 0 0 0 1

AISI 4130 1 2 1 1 3

321 SS 0 * 0

* *

CDA-110 0 * 0

* *

CDA-172 0 * 0

* *

AM 355 0 * 0

* *

Inconei 615 0 * 0

* *

(0) No visual evidence attack

(1) Average observed pit depth < 0.06 mm

(2) Average observed pit depth between 0.06 mm and 0.1 mm

(3) Average observed pit depth > 0.1 mm

(*) Alloy not tested in this environment
146 5. COMPATIBILITY WITH METALS

Table 7. Mass change rate per unit area (g/meter • day) at 20 °C estimated from the mass

change measurements divided by the exposure time and sample area. This analysis assumes the reaction rate does not vary with time (bare surface kinetics)

HFC125 HFC-227ea Halon 1301 FC-218 CF3I

Alloy

Std Std Std Std Std R Dev K Dev R Dev K Dev K Dev

Nit 40 0.00084 0.00078 0.00039 0.00055 0.00090 0.00106 0.00057 0.00046 0.00031 0.00110

Al 6061-T6 0.00354 0.00277 0.00322 0.00183 0.00548 0.02867 0.00474 0.00253 0.22589 0.50962

u.uuizo nu.uui nni onvu 0.00049 0.00045 yj.yJV I U'f 0.00056 0.00055 0.00248 0.00238

Ti 15-3-3-3 0.00058 0.00036 0.00062 0.00046 0.00080 0.00052 0.00084 0.00036 0.00149 0.00182

AISI 4130 0.00974 0.00884 -0.00352 0.01229 0.02156 0.02336 0.00731 0.00271 0.71189 0.23330

321 SS 0.00030 0.00027 * * 0.00069 0.00010 * * * * CDA 110 0.00039 0.00073 * * 0.00140 0.00078 * * * * CDA 172 0.00006 0.00058 * * -0.00020 0.00031 * * * * AM 355 -0.00001 0.00077 * * 0.00022 0.00047 * * * *

IN 625 0.00004 0.00012 * * 0.00016 0.00015 * * *

(+) Values mean mass increase (scaling)

(-) Values mean mass decrease (corrosion)

* No measurement made in the agent for this alloy R = Average mass change rate (g/meter^ day)
5. COMPATIBILITY WITH METALS 147

Table 8. Mass change rate per unit area (g/meter^ • day) at 20 °C estimated by the regression

slope for a linear regression of mass change per unit area on exposure time(assumes

constant, bare surface, reaction kinetics)

Alloy

HFC- 125 HFC-227ea Halon 1301 m R Std Err m R Std Err m R Std Err

Nit 4U 0.00078 0.637 0.00023 0.00049 0.825 0.00009 0.00042 0.660 0.00012

A 1 <^ru^ 1 T£. A\ oUol - 1

0

0.791 0 00348 0.934 n 00748

304 SS 0.00061 0.495 0.00027 0.00046 0.824 0.00009 0.00060 0.691 0.00016

Ti 15-3-3-3 0.00053 0.900 0.00006 0.00071 0.934 0.00008 0.00051 0.886 0.00007

AISl 4130 0.00894 0.664 0.00380 -0.00352 0.241 0.00709 0.02791 0.603 0.01395

321 SS + T t

* * % t t CDA 110 + ±

* * * t t t CDA 172 +

+

J.+

* * * t t AM 355 +

i

tT

* * X t %

IN 625 * * * t %

Table 8. (continued)

Alloy

FC-218 CF3I m R Std Err tn R Std Err

Nit 40 0.00048 0.753 0.00012 0.00149 0.798 0.00036

Al 6061-T6 0.00459 0.900 0.00062 0.03963 0.066 0.18881

304 SS 0.00058 0.800 0.00012 0.00409 0.919 0.00055

Ti-15-3-3-3 0.00082 0.944 0.00008 0.00257 0.860 0.00048

AISI 4130 0.00731 0.920 0.00156 0.55297 0.949 0.06940

321 SS « * * * * CDA 110 * * * * * * CDA 172 * * * * * * AM 355 * * * * *

IN 625 * + * * *

* No measurements were made in this alloy/environment combination

+ Only one exposure time used for these measurements, see the average mass change rate table m = estimated mass change rate for linear regression of mass change per unit area on exposure times (g/meter^ • day)

R = correlation coefficient for linear regression Std Error = estimated standard deviation for the mass change rate (regression slope)
148 5. COMPATIBILITY WITH METALS

Table 9. Mass change rate per unit area (g/meter^ • day) at 20 °C estimated by the slope for a

linear regression of mass change per unit area on square root of the exposure time

(assumes a uniformly filmed surface)

Alloy

HFC- 125 HFC-227ea Halon 1301

k R Std En- k R Std Err k R Std Err

Nit 40 0.01491 0.602 0.00494 0.00696 0.823 0.00133 0 00922 0.714 0 00226

Al 6061-T6 0.04856 0.829 0.008 1

8

0.04711 0.907 0.00606 -0.18978 0.289 0.01570

304 SS 0.0133! 0.530 0.00532 0.00622 0.809 0.00126 0.01327 0.761 0.02820

Ti 15-3-3-3 0.01084 0.913 0.00121 0.00965 0.904 0.00126 0.01078 0.925 0.00110

AISI 4130 0.10716 0.654 0.04691 -0.02725 0.241 0.05492 0.31117 0.552 0.17774

321 SS t +

* * * t t t CDA 110 t

* * * t t t CDA 172 t t

* * * t t t AM 355 t t t

+ * * t t t IN 625 t t t

+ * * t t t

Table 9. (continued)

Alloy

FC-218 CF3I

k R Std Err k R Std Err

Nit 40 0.00688 0.766 0.00160 0.01350 0.663 0.00483

A! 6061-T6 1.63021 0.894 1.63021 0.46755 0.070 2.09612

304 SS 0.00785 0.776 0.00177 0.04108 0.831 0.00869

Ti 15-3-3-3 0.01113 0.917 0.00135 0.02562 0.772 0.00668

AlSl 4130 0.05658 0.920 0.01209 6.92014 0.975 0.59885

321 SS * * * * * CDA 110 * * * * * * CDA 172 * * * * * * AM 355 * * * * *

IN 625 * * * * * *

* No measurements were made in this alloy/environment combination

t Only one exposure time used for these measurements, see the average mass change rate table k = estimated mass change rate for linear regression of the mass change per unit area on square root of exposure time

(g/meter^ • day"^) R = correlation coefficient for linear regression Std Error = estimated standard deviation for the mass change rate (regression slope)
5. COMPATIBILITY WITH METALS 149

Table 10. Mass change rates at 150 °C for a regression model that assumes mixed bare and filmed

surface kinetics

HFC- 125

Alloy M„ MQ(en"or) M, M, (error) M2 M2(error) R

Nit 40 0.00035 0.06907 0.00075 0.00076 0.00047 0.01533 0.637

Al 6061-T6 -0.05633 0.11893 0.00014 0.00131 0.04588 0.02640 0.830

304 SS -0.00965 0.07734 -9.53e-5 0.00085 0.01513 0.01716 0.531

Ti- 15-3-3-3 -0.01411 0.01693 0.00020 0.00019 0.00700 0.00376 0.920

AISI 4130 1.05e-10 0.4376 0.02643 0.04432 -0.03015 0.33373 0.666

321 SS 1: t t t

T*T CDA 110 J. f f t t CDA 172 J. t t

+ AM 355 t t

+T ?

IN 625 t + t +

Table 10. (continued)

Halon 1301

Alloy M„ M()(error) M, M, (error) M2(error) R

Nit 40 -0.00814 0.03247 -0.00012 0.00036 0.01154 0.00721 0.720

Al 6061-T6 -0.39446 2.1183 -0.03640 0.02326 0.50713 0.47012 0.461

304 SS -0.02406 0.04067 -0.00026 0.00045 0.01815 0.00903 0.767

Ti-15-3-3-3 -0.00970 0.01602 5.18e-5 0.00018 0.00978 0.00355 0.926

AISI 4130 7.29e-15 1.593 0.15669 0.157 -0.68495 1.1102 0.614

321 SS t t t t t t t CDA 110 ± T t

J.+ t CDA 172 t + t t t AM 355 t t ± i

IN 625 t t t t i t i
150 5. COMPATIBILITY WITH METALS

Table 10. (continued)

FC-218

Alloy Mo Mo(error) Ml M| (error) M2(error) R

Nit 40 -0.00226 0.01695 0.00022 0.00039 0.00378 0.00545 0.764

Al 6061-T6 -0.03239 0.08801 0.00268 0.00203 0.02799 0.02829 0.908

304 SS -0.00210 0.01788 0.00047 0.00041 0.00154 0.00575 0.801

11-15-3-3-3 -U.U02U5 U.UI 165 U.UWJ6/ n AAAT7 O.UOUz/ 0.00225 0.00374 0.946

AISI 4130 n/a n/a n/a n/a n/a n/a n/a

321 SS * * * * * * CDA 110 * + * * * * * CDA 172 * * * * * * * AM 355 * * * * * * *

IN 625 * * * * *

Table 10. (continued)

HFC-227ea

Alloy Mo Mq(error) M, M| (error) M, M2(error) R

Nit 40 -0.00502 0.01389 0.00037 0.00032 0.00177 0.00447 0.827

Al 6061-T6 -0.01883 0.05417 0.00280 0.00125 0.00987 0.01742 0.936

304 SS -0.00142 0.01276 0.00032 0.00030 0.00194 0.00410 0.828

Ti- 15-3-3-3 -0.00467 0.01119 0.00060 0.00026 0.00172 0.00360 0.935

AISI 4130 n/a n/a n/a n/a n/a n/a n/a

321 SS * * * * * * * CDA 110 * * * * * * * CDA 172 * * * * * * AM 355 * * * * * *

IN 625 * * * * *
5. COMPATIBILITY WITH METALS 151

Table 10. (continued)

CF3I

Alloy Mo MgleiTor) M, M, (error) M2 MjCerror) R

Nit 40 -0.00019 0.0267 0.00353 0.00098 -0.02389 0.01083 0.874

Al 6061-T6 -3.2761 15.868 -0.83038 0.57971 9.2353 6.4376 0.436

304 SS -0.00955 0.04828 0.00597 0.00176 -0.0220 0.01959 0.929

Ti- 15-3-3-3 -0.00551 0.04308 0.00392 0.00157 -0.0159 0.01718 0.872

AISl 4130 I.2!e-I0 5.605 0.71758 0,56594 5.443 4.232 0.9706

321 SS * * * * * * * CDA 110 * * * * * * CDA 172 * * * * * * * AM 355 * * * * * *

IN 625 * * * * * * *

* No measurements were made in this alloy/environment combination

t Only one exposure time used for these measurements, see the average mass change rate table. niQ = g/meter^ m, = g/meter^ • day

mj = g/meter^ • day"^
152 5. COMPATIBILITY WITH METALS

Table 1 1 . Results of visual inspection of coupons tested at 20 °C

Alloy HFC- 125 HFC-227ea Halon 1301 FC-218 CF3I

Nit 40 0 1 0 0 0

Al 6061-T6 1 1 3 0 0

304 SS 0 0 0 0 0 .

Ti 15-3-3-3 0 0 0 0 0

AISI 4130 3 1 1 1 3

321 SS 0 * 0

* *

CDA-110 0 * 0

* *

CDA-172 0 * 0

*

AM 355 0 * 0

* *

Inconel 615 0 * 0

*

(0) No visual evidence attack observed

(1) Average observed pit depth < 0.06 mm

(2) Average observed pit depth between 0.06 mm and 0.1 mm

(3) Average observed pit depth > 0.1 mm

(*) Alloy not tested in this environment

pitting in the halon 1301 at 150 °C, it did however, exhibit some pitting in that agent at 20 °C. This

result was unexpected and may be an indication that this alloy has some compatibility problems with

the halon 1301. The 4130 alloy steel, once again, exhibited some pitting in every environment with

the most severe pitting observed in the CF3I.

5.3 Environmentally Induced Fracture Experiments (Slow Strain Rate

Tensile Tests)

The environmentally induced fracture resistance of the alloys was evaluated by conducting exposure

tests in each of the replacement candidates. This was accomplished by loading cylindrical tensile

specimens of the alloys in tension, by a slow increase in the applied strain, until failure occurred by

either a normal mechanical or an environmentally assisted means. Comparisons between the stresses and strains required to cause failure in an inert environment and in the agent at the same temperature

provided a measure of the propensity for that agent to cause an environmentally induced failure.

5.3.1 Materials. The materials selected by the sponsors for Phase II were: 304 stainless steel,

stainless steel alloy 21-6-9 (Nitronic 40), and beta-titanium alloy 15-3-3-3. These alloys were selected
5. COMPATIBILITY WITH METALS 153

because they are presently in service, or under consideration for future use, in the agent storage and

distribution systems on board jet aircraft. The compositions of these alloys are also given in Table 1 (Section 5.2.1).

5.3.2 Experimental Procedure at 150 °C. The alloys were evaluated for stress corrosion cracking

susceptibility in the replacement candidates by the slow strain rate (SSR) tensile test technique. This

technique was selected because it generates intrinsic mechanical properties data for a given alloy and it also reveals any interactions that may have occurred between that alloy and the testing environment

within a relatively short time frame (Ricker, 1994, Stoudt and Fink 1994, and Stoudt and Ricker,

1994).

All of the samples used for these experiments were machined with the tensile axis parallel to the

rolling direction of the plate stock (Figure 13) and tested in the "as received" condition. The sample

preparation consisted of a measurement of the appropriate quantities (e.g., gauge length) followed by

degreasing in acetone and alcohol. The vessels used for these experiments were 250 ml volume

autoclaves with a design similar to those used for the exposure testing, except that these vessels were

modified so that load could be applied to the tensile specimen in-situ under constantly maintained

environmental conditions as shown in Figure 14 (Ricker, 1994). The testing conditions used for this portion of the analysis were nominally 5.86 MPa at 150° ± 1 °C.

The test vessels were evacuated and charged with agent in the same manner as the exposure test vessels, but due to the reduced capacity of the test vessels, the appropriate mass was obtained by a

calculation based on the ideal gas law. This approach was selected over the mass balance shown in Equation (1) (Section 5.2.2) for two reasons. First, the number of moles of gas could be held constant allowing for a consistent agent concentration for each test. Second, this model generated a slightly

lower pressure in the cells reducing the risk of a rupture.

The mechanical tests were conducted using a computer controlled SSR testing system which

operated at a constant crosshead speed of 2.54 x

10'^ m/s (1.0 x 10"^ in/s). The computer was

configured to sample and record the applied load, the crosshead displacement and the elapsed time at 90 s intervals. After failure, the agent was released, the vessels were then allowed to cool to ambient

temperature, and the samples were removed from the vessel and stored in a desiccator until analyzed.

The fracture surfaces were sectioned from the broken samples and prepared for analysis. The

influence of the agents on the ductility of the alloys was determined from reduction in area (RA)

measurements performed on the fracture surfaces with an optical measuring microscope with a ± 0.5 X 10"^ m resolution. All samples were examined visually and scanning electron microscopy was

performed on selected samples to verify the presence of stress corrosion cracking. The results of these

experiments were used to assess the potential for failure by stress corrosion cracking for each of the

alloys in a replacement candidate.

5.3.3 Experimental Procedure at 20 °C. A second series of slow strain rate tests was conducted at ambient conditions in order to determine the possible influences of the liquid phase and/or the

meniscus regions on the stress corrosion cracking resistance of these materials. To do this,

Equation (1) was utilized to determine an appropriate mass of agent so as to locate the liquid/vapor

meniscus region at the approximate midpoint of the gauge section of the tensile specimen. The

resulting cell pressures varied according to the density of each agent, but were on the order of

1.38 MPa at 20 ± 1 °C. The remainder of the procedure was identical to the one outlined in the

previous section.

5.3.4 Results at 150 °C. Environmentally induced fracture is the initiation and/or propagation of a

crack at stresses well below those normally required to cause fracture resulting from simultaneous
154

5. COMPATIBILITY WITH METALS

^llllllllllll ODUUIUU

M6x20 Threads • (1/4 X 20)

D = 19mm (0.25 in) L = 178 mm (7 in) g = 25.4 mm (1 in) h = 57 mm (2.25 in) s = 19 mm (0.75 in)

Figure 13. Slow strain rate tensile sample design (ASTM E-8, G-49).
155

Figure 14. Schematic diagram of the slow strain rate test chamber.
156 5. COMPATIBILITY WITH METALS

exposure to a chemically reactive environment and a mechanical stress (Stoudt, 1995 and Stoudt,

1994). This particular form of corrosion occurs in specific material/environment combinations and

while attack may not be readily apparent, failures can be sudden and catastrophic.

The same three parameters used to indicate the susceptibility of an alloy to environmentally

induced failure used for Phase I were applied to the Phase n data. These are: the ultimate tensile

strength (UTS), the strain to failure (e^), and the reduction in cross sectional area (RA) (Ricker, 1994).

The UTS is an indication of the fracture strength of the sample and it is determined from the

maximum load observed during the tensile test according to the relationship:

where P^^^^ is the maximum load supported by the sample during the test and is the initial crosssectional area of the gauge section of the sample (Deiter, 1986). If cracking, or some other environmental interaction that promotes crack initiation or propagation at lower stresses, occurs, it will

produce a noticeable change in this parameter (Stoudt, 1994). Environmental interactions can also

alter deformation processes on the surfaces and at tips of cracks formed in the sample. One of the two

parameters used to measure the ductility is the engineering strain to failure (e^) which can be

determined from the relationship:

where Lf is the total change in sample length during the experiment and is the initial gauge length

(Deiter, 1986). This measurement includes both the elastic and plastic deformation components

required to induce failure, which are measured by displacement gages located outside the environmental chamber increasing measurement errors. As a result, a ductility determination based solely on this quantity may contain significant experimental error. Another means for determining the ductility of a

material is the reduction in area (RA). This quantity is obtained from physical measurements

performed on the fracture surface after completion of the tensile test from the relationship:

where Af is the final cross-sectional area of the fracture surface and A^ is the initial gauge section

cross-sectional area (Deiter, 1986). Unlike the strain to failure (STF), the RA measurement only

includes the plastic deformation component required to induce failure and does not include measure- ment errors which may have occurred during the time the cell was ascending to the test temperature.

Therefore, this is a better measure of the environmental interactions. Typically, ail three of these

parameters are analyzed by forming a ratio of the value observed in the environment to the value

observed in an inert reference environment, argon (Ar), (Ricker, 1994 and Stoudt et ai, 1994).

Another approach for analysis of this data is to estimate the statistical significance of the

difference between the mean determined for the parameter in the inert reference environment and the mean determined for the agent (Ricker, 1994). This analysis consisted of the calculation of Student's

t-statistic for the significance of the difference between the mean determined in each environment and
5. COMPATIBILITY WITH METALS 157

the mean for the same parameter and alloy tested in argon at the same temperature. This statistic is calculated from the relationship:

where: ygnvir mean of the tests in the agent, y^^ is the mean of the tests in argon, S^^^-^ is the

variance of the tests conducted in the agent, S^^ is the variance of the tests conducted in argon, n^^yy.

is the number of samples tested in the agent, and n^^ is the number of samples tested in argon all at the same temperature (Mendenhall, 1992).

Table 12 gives the average UTS determined for each alloy in argon at 150 °C, and in each agent

at 150 °C. The standard deviations for these calculated values are presented in Table 13. The ratio of

the average UTS in each agent to the average UTS in argon at the same temperature is given in Table 14. In this table, it can be seen that 18 of the 20 agent/alloy combinations have a higher

average UTS values in the agent than in argon. Table 15 contains the results of a t-test estimate of the statistical significance of the observed

differences using the alloy standard deviations. For a confidence level of 99.5 %, t determined in this manner (4 degrees of freedom) should be less than -4.60 to conclude that the environment significantly

reduced the strength or ductility of the alloy. Similarly, a value greater than +4.60 would indicate a

significant increase in strength or ductility caused by the environment. Values between -4.60 and

+4.60 are considered statistically insignificant. A significant decrease in the average UTS may be an

indication of cracking, but it could also be the resuk of corrosion reactions which reduced effective

cross-section or otherwise assisting deformation. An increase in the average UTS is unusual and this may suggest that sample/environment interactions are inhibiting deformation and fracture, but it could

also be an indication of an interference between the corrosion products being generated on the sample

and the seal of the autoclave through which the sample must slide. The most important point is that no agent caused statistically significant changes in all of the alloys which demonstrates that suitable

materials can be identified for the containment of any of these candidates (Ricker, 1994).

The ductility results for the 150 °C SSR tests are given in Tables 16 and 17. Table 16 gives the

average strain to failure for each alloy/environment combination, and in argon at 150 °C. Table 17

shows the standard deviations for the STF calculations. The ratio of the average strain to failure for

each alloy in each environment to that observed in the same alloy in argon at the same temperature is presented m Table 18. The significance of the difference is again presented as a function of the alloy

standard deviations in Table 19. Environmental induced cracking, or stress corrosion cracking (SCC),

is usually evidenced by a reduction in ductility. In Table 18, it can be seen that strain to failure ratio generally increased for Nitronic 40 and for 304 stainless for most of the alloy/environment combinations. The titanium 15-3-3-3 alloy demonstrated significant decreases in ductility for all of the

environments. This is best reflected in the CF3I data. The sharp decrease in the ductility was the

result of severe cracking which was observed on two of the three specimens tested in this agent.

However, these strain to failure measurements are based on load frame displacement measurements

taken by a transducer positioned outside the autoclave during the experiment. As a result, the

reduction in area measurements, which are based on measurements performed on the fracture surface

with an optical microscope after the experiment, are a more reliable indication of changes in ductility.

Table 20 presents the RA data for each alloy in argon gas at 150 °C and each agent at 150 °C.

Table 21 gives the standard deviations for these measurements. The ratio of the average reduction in

(9)

'envir ^Ar +
158 5. COMPATIBILITY WITH METALS

Table 12. Average ultimate tensile strength at 150 °C (MPa)

Environment Nitronic 304 Titanium 40 SS 15-3-3-3

Argon jyy o4y 497

Halon 1301 603 671 497

FC-218 602 646 504 HFC-227ea 611 655 534 HFC- 125 611 651 552

CF3I 607 657 505

Table 13. Standard deviation of the ultimate tensile strength measurements at 150 °C (MPa)

Environment Nitronic 304 Titanium 40 SS 15-3-3-3

Argon 1.5 0.6 6.0

Halon 1301 2.3 1.5 3.4

FC-218 3.7 0.8 4.7

HFC-227ea 4.8 2.0 5.7 HFC- 125 3.6 2.4 2.0

CF3I 0.9 2.0 2.2

Table 14. Ultimate tensile strength ratios at 150 °C [(UTS in Agent)/(UTS in Ar)]

Environment Nitronic 304 Titanium 40 SS 15-3-3-3

Argon * *

Halon 1301 1.007 1.035 1.000

FC-218 l.OOo 0.996 1.015

HFC-227ea 1.021 1.010 1.076 HFC- 125 1.020 1.003 1.112

CF3I 1.013 1.012 1.017
5. COMPATIBILITY WITH METALS 159

Table 15. t-Test for significance in UTS change at 150 °C

Environment Nitronic 304 Titanium 40 SS 15-3-3-3

Argon

Halon 1301 0.281 2.109 0.003

FC-218 0.226 -0.224 0.276

HFC-227ea 0.809 0.582 1.377 HFC- 125 0.776 0.198 2.035

CF3I 0.514 0.745 0.304

Table 16. Average strain to failure at 150 °C (%)

Environment Nitronic 304 Titanium 40 SS 15-3-3-3

Argon 43.0 17.2 6.6

Halon 1301 44.9 18.1 5.7

FC-218 43.8 16.9 6.1

HFC-227ea 43.7 17.0 5.9 HFC- 125 43.1 17.2 6.0

CF3I 41.7 18.1 3.8

Table 17. Standard deviation of the strain to failure measurements at 150 °C (%)

Environment Nitronic 304 Titanium 40 SS 15-3-3-3

Argon 0.5 2.3 1.9

Halon 1301 3.8 1.7 0.8

FC-218 2.6 1.5 0.8

HFC-227ea 2.1 0.6 0.6 . HFC- 125 1.0 0.4 0.6

CF3I 1.0 4.0 2.5
160 5. COMPATIBILITY WITH METALS

Table 18. Strain to failure ratios at 150 °C [(STF in Agent)/(STF in Ar)]

Environment Nitronic 304 Titanium 40 SS 15-3-3-3

Ar^on * *

Halon 1301 1.045 1.056 0.871

FC-218 1.019 0.986 0.927 HFC-227ea 1.018 0.991 0.899 HFC- 125 1.004 1.000 0.912

CF3I 0.972 1.056 0.571

Table 19. t-Test for significance in STF change at 150 °C

Environment Nitronic 304 Titanium 40 SS 15-3-3-3

Argon * * *

Halon 1301 1.119 0.619 -0.695

FC-218 0.464 -0.159 -0.394

HFC-227ea 0.457 -0.098 -0.543 HFC- 125 0.098 -0.005 -0.472

CF3I -0.712 0.623 -2.312

Table 20. Average reduction in area at 150 °C (%)

Environment Nitronic 304 Titanium 40 SS 15-3-3-3

Argon 77.6 75.0 54.8

Halon 1301. 80.2 68.4 56.2

FC-218 80.5 68.4 58.5

HFC-227ea 79.2 68.2 55.7 HFC- 125 68.2 80.9 54.0

CF3I 79.4 66.6 44.9
5. COMPATIBILITY WITH METALS 161

Table 21. Standard deviation of the reduction in area measurements at 150 °C (%)

Environment Nitronic 304 Titanium 40 SS 15-3-3-3

Ar^on nV/. 7/ 7 7

Halon 1301 1-2 0.7 3.2

FC-218 0.7 1.5 8.4

HFC-227ea 1.1 0.9 2.7 HFC- 125 0.9 0.5 3.6

CF3I 0.4 2.4 29.9

area measured for each alloy/agent combination to argon at 150 °C is presented in Table 22. The

statistical significance of the difference is again presented as a function of the alloy standard deviations

in Table 23. In Table 22, it can be seen that the reduction in area values for Nitronic 40 increase in every agent except for HFC- 125, while 304 stainless has the opposite behavior. The titanium alloy

performed well in halon 1301, FC-218 and HFC-227ea, but exhibited significantly lower values in HFC- 125 and in CF3I.

In general, each alloy performed well in one or more of the agents evaluated at 150 °C and the

titanium alloy exhibited brittle cracking in two of the three samples tested in the CF3I. Figure 15 is a low magnification scanning electron micrograph exhibiting the crack morphology observed in the

titanium 15-3-3-3 samples tested in CF3I at 150 °C. In this figure it can be seen that the entire

fracture surface consists of brittle, transgranular cleavage cracks, which indicates this alloy has a

strong sensitivity to this agent (Mills, 1987). Figure 16 is a higher magnification view of the observed

cracking. This cracking could be the result of interactions between the titanium alloy and one or more

of the decomposition products on the CF3I. However, it is also possible that this cracking may have

been due to an interaction with an impurity present in the CF3I. Figures 17 and 18 are scanning

electron micrographs of the alloy tested in argon at the same temperature included for comparison. In these figures, it can be seen that the fracture surfaces are entirely composed of microvoid coalescence

which indicates the failures occurred by completely ductile processes. Figure 18 is a higher magnification view of the microvoid coalescence (Korb, 1987, Mills, 1987, and Boyer, 1987). These figures

indicate that the titanium alloy may not be a suitable alloy for the containment of CF3I.

5.3.5 Results at 20 °C. The slow strain rate data from the tests conducted at 20 °C are presented in a fashion similar to those used for the 150 °C tests. Table 24 gives the average UTS determined for

each alloy in argon at 20 °C, and in each agent at 20 °C. The standard deviations for these calculated

values are presented in Table 25. The ratio of the average UTS in each agent to the average UTS in argon at the same temperature is given in Table 26. In this table, it can be seen that all of the 20

agent/alloy combinations have a higher average UTS in the agent than in argon. Tables 27 contains the results of an estimate of the significance of the differences. As before, a

significant increase in the average UTS is unusual and this suggests that sample/environment

interactions are inhibiting deformation and fracture, but it could also be an indication of an interfer- ence between the corrosion products being generated on the sample and the seal of the autoclave

through which the sample must slide. The most important point is, once again, suitable materials can

be identified for the containment and reliable distribution of any of these candidates (Ricker, 1994).
162 5. COMPATIBILITY WITH METALS

Figure 15. Scanning electron micrograph of the transgranuiar cleavage-like fracture observed in Ti 15-3-3-3 when tested in CF3I at 150 °C.
COMPATIBILITY WITH METALS

Figure 16. Higher magnification view of the transgranular cleavage-like crack morphology

observed in titanium tensile specimen tested in CF3I at 150 °C.
164 5. COMPATIBILITY WITH METALS

Figure 17. Scanning electron micrograph of the ductile microvoid coalescence type of fracture

observed in Ti 15-3-3-3 when tested in argon at 150 °C.
5. COMPATIBILITY WITH METALS 165

Figure 18. Higher magnification view exhibiting the microvoid coalescence observed on the

titaniunn tensile specimens tested in argon at 150 °C.
166 5. COMPATIBILITY WITH METALS

Table 22. Reduction in area ratios at 150 °C [(RA in Agent)/(RA in Ar)]

Environment Nitronic 304 Titanium 40 SS 15-3-3-3

Argon + * *

Halon 1301 1.035 0.912 1.026

FC-218 1.038 0.912 1.068

HFC-227ea 1.021 0.909 1.015 HFC- 125 0.879 1.078 0.985

CF,i 1.024 0.888 0.820

Table 23. t-Test for significance in RA change at 150 °C

Environment Nitronic 304 Titanium 40 SS 15-3-3-3

Argon * * *

Halon 1301 0.699 -1.548 0.142

FC-218 0.768 -1.562 0.370

HFC-227ea 0.431 -1.615 0.084 HFC- 125 -2.451 1.377 -0.084

CF3I 0.476 -1.977 -0.983

Table 24. Average ultimate tensile strength at 20 °C (MPa)

Environment Nitronic 304 Titanium 40 SS 15-3-3-3

Argon 718 793 579

Halon 1301 734 806 597

FC-218 735 805 627

HFC-227ea 726 799 633 HFC- 125 746 803 606

CF3I 724 803 603
5. COMPATIBILITY WITH METALS 167

Table 25. Standard deviation of the ultimate tensile strength measurements at 20 °C (MPa)

Environment Nitronic 304 Titanium 40 SS 15-3-3-3

Argon 3.0 2.1 2.1

Halon 1301 3.0 2.5 2.4

FC-218 3.0 2.3 0.6

HFC-227ea 3.4 1.9 3.8 HFC- 125 0.3 2.9 3.8

CF3I 3.0 2.8 4.6

Table 26. Ultimate tensile strength ratios at 20 °C [(UTS in Agent)/(UTS in Ar)]

Environment Nitronic 304 Titanium 40 SS 15-3-3-3

Argon * *

Halon 1301 1.023 1.016 1.031

FC-218 1.024 1.016 1.084

HFC-227ea 1.012 1.008 1.093 HFC- 125 1.039 1.013 1.048

CF3I 1.009 1.012 1.042

Table 27. t-Test for significance in UTS change at 20 °C

Environment Nitronic 40

304 SSTitanium

15-3-3-3

Argon * * *

Halon 1301 1.072 1.076 0.840

FC-218 1.111 1.051 2.232

HFC-227ea 0.540 0.520 2.486 HFC- 125 1.830 0.856 1.270

CF3I 0.414 0.816 1.109
168 5. COMPATIBILITY WITH METALS

The ductility results for the 20 °C SSR tests are given in Tables 28-30. Table 28 gives the

average strain to failure for each alloy/environment combination, and in argon at 20 °C. Table 29

shows the standard deviations for the STF calculations. The ratio of the average strain to failure for

each alloy in each environment to that observed in the same alloy in argon at the same temperature is presented in Table 30. The significance of the difference is presented as a function of the alloy

standard deviations, in Table 31. In Table 30, it can be seen that the STF ratios decreased for

Nitronic 40 in all of the environments. However, both the 304 stainless steel and the titanium alloy

exhibited increases in the STF ratio values in most of the environment/alloy combinations. This

suggests that deformation may have been easier in these environments than in argon at this temperature. Table 32 presents the RA data for each alloy in argon gas at 20 °C and each agent at 20 °C.

Table 33 gives the standard deviations for these measurements and the ratio of the average reduction

in area measured for each alloy/agent combination to argon at 20 °C is presented in Table 34. The

statistical significance of the difference is again presented as a function of the alloy standard deviations

in Table 35. In Table 34, it can be seen that the reduction in area values for Nitronic 40 increase in every agent. The 304 stainless steel alloy exhibited slight decreases in the RA values for all agents

and the titanium alloy exhibited increases in every agent except for HFC-227ea, where it shows a

significant decrease.

Based on these results, none of the alloys appear to be susceptible to environmentally induced

fracture in these agents at 20 °C. Unlike the tests conducted at 150 °C, no evidence of environmentally induced cracking of any kind was observed on the fracture surface of any specimen tested at 20 °C. The values presented in these tables do not readily indicate any susceptibility to environmentally induced fracture at 20 °C.

5.4 Post Deployment Corrosion

When a fire suppressant is applied to a fire, the metals in the aircraft engine nacelle or the dry bays

may become covered with deposits of the fire suppressant or fire suppressant combustion products.

An additional evaluation of the relative corrosivity of the expected combustion products was included

in the investigation. As a result, an evaluation of the relative corrosion behavior of representative

structural materials with the combustion by-products was included in this investigation.

All of the potential replacements investigated in this study contain fluorine in their molecules and

will produce fluoride ions and hydrofluoric acid (HF) during combustion. These ions are known to be

aggressive when they contact certain alloy types. HF is commonly used to etch glasses and ceramics

and will similarly attack the protective layer on the surface of passivated alloys. If the fluoride ion

concentration in the combustion products deposited on the surface of an alloy exceeds some critical value, then pitting or crevice attack will result (Fontana, 1987, Zotikov, 1974, and Ricker, 1994). As a

result, the post deployment corrosion damage to aircraft alloys will depend on the corrosivity of the

surface films which contain these ions. Most engineering alloys that exhibit good corrosion resistance do so because they form a

protective surface film. This passive film is usually composed of metallic oxides or hydroxides

(Fontana, 1987). Commercial aluminum alloys and stainless steels are examples of alloys which

contain active elements, but behave in a relatively noble manner because of the formation of oxide

films (Ricker, 1994). These protective surface films may be crystalline in form and may be precipitated from the solution adjacent to the bare metal surface, but it is widely believed that rapidly grown

amorphous films are the most protective (Ricker, 1994). No matter what the structure of the surface
5. COMPATIBILITY WITH METALS 169

Table 28. Average strain to failure at 20 °C (%)

Environment Nitronic 304 Titanium 40 SS 15-3-3-3

Argon 65.0 43.2 5.7

Halon 1301 63.1 44.4 6.9

FC-218 62.5 45.9 6.1

HFC-227ea 63.4 45.0 5.9 HFC- 125 61.5 43.1 6.4

CF3I 63.3 44.0 6.5

Table 29. Standard deviation of the strain to failure measurements at 20 °C (%)

Environment Nitronic 304 Titanium 40 SS 15-3-3-3

Argon 0.9 2.5 0.7

Halon 1301 3.8 2.2 0.4

FC-218 3.7 2.7 0.2

HFC-227ea 2.1 '1.3 3.5 HFC- 125 1.4 3.2 1.0

CF3I 1,1 3.0 0.1

Table 30. Strain to failure ratios at 20 °C [(STF in Agent)/(STF in Ar)]

Environment Nitronic 304 Titanium 40 SS 15-3-3-3

Argon * *

Halon 1301 0.971 1.028 1.227

FC-218 0.962 1.062 1.073

HFC-227ea 0.975 1.041 1.045

HFC- 125 0.947 0.999 1.139

CF3I 0.974 1.020 1.154
170 5. COMPATIBILITY WITH METALS

Table 31. t-Test for significance in STF change at 20 °C

Environment Nitronic 304 Titanium 40 SS 15-3-3-3

Argon * *

Halon 1301 -1.006 0.624 1.211

FC-218 -1.301 1.371 0.389 HFC-227ea -0.861 0.902 0.240 HFC- 125 -1.828 -0.030 0.741

CF3I -0.887 0.439 0.819

Table 32. Average reduction in area at 20 °C (%)

Environment Nitronic .^04 Titanium 40 SS 15-3-3-3

Argon 77.6 75.0 54.8

Halon 1301 78.7 74.3 60.3

FC-218 78.2 74.4 55.4

HFC-227ea 78.1 74.2 46.0

HFC- 125 78.5 74.8 68.0

CF3I 78.0 74.7 65.6

Table 33. Standard deviation of the reduction in area measurements at 20 °C (%)

Environment Nitronic 304 Titanium 40 SS 15-3-3-3

Argon 1.2 0.4 6.9

Halon 1301 1.5 2.0 4.7

FC-218 0.7 0.6 4.9

HFC-227ea 1.3 0.3 23.9 HFC- 125 0.9 1.0 25.4

CF3I 0.4 0.5 3.1
5. COMPATIBILITY WITH METALS 171

Table 34. Reduction in area ratios at 20 °C [(RA in Agent)/(RA in Ar)]

Environment Nitronic 40

304 SS

Titanium

15-3-3-3

Argon * *

Halon 1301 1.015 0.990 1.101

FC-218 1.008 0.991 1.011

HFC-227ea 1.007 0.989 0.839 HFC- 125 1.012 0.997 1.241

1.005 0.996 1.197

Table 35. t-Test for sig nificance in RA change at 20 °C

Environment Nitronic 40

304 SSTitanium

15-3-3-3

Argon * *

Halon 1301 1.433 -0.981 0.463

FC-218 0.740 -0.877 0.049

HFC-227ea 0.660 -1.154 -0.739

HFC-125 1.175 -0.262 1.110

CF3I 0.474 -0.384 0.904

film, halogen 10ns generally tend to destabilize and/or reduce the protective nature of these films

(Ricker, 1994).

Light weight alloys such as aluminum alloys are commonly used for aircraft applications.

Unfortunately, these alloys are also susceptible to pitting corrosion in the presence of halide ions.

Aluminum is a very active metal and alloys of this type will generally corrode as rapidly as mass

transport kinetics will allow if the protective surface film is removed (Ricker, 1994). For this reason aluminum alloys are not used in environments where the passivity cannot be maintained (i.e., pH<4 or pH>10). Other passivated materials such as titanium alloys and stainless steels are also susceptible to passivity breakdown.

In phase 1 of this study, the atmospheric corrosion behavior of representative aircraft alloys

whose surfaces were contaminated with different chemical species expected to be produced by the

combustion of different fire suppressants was investigated. In that investigation, three different surface

pretreatments were used: (1) artificial seawater plus NaF, (2) artificial seawater plus

NaHC03/Na2C03, and (3) artificial seawater plus NaOH. The last two of these were used to represent

surface conditions that may result from the use of NaHCOj as a fire suppressant and, therefore, are not

relevant to this study. The first was used to represent surface conditions that may result from the

combustion of fluoride or other halogen containing fire suppressants. When a surface film on a metal

contains HP, the HF will react with the metal to form metal ions and hydrogen gas. This effectively
172 5. COMPATIBILITY WITH METALS

removes the hydrogen ion and increases the pH of the solution in the surface film. For NaF, the initial pH is higher, but the metal ions produced by corrosion will hydrolyze reducing the pH of the surface

film. The end result is, whether the starting pH is low or high, the pH of the surface film will be

determined by the chemical species present and the hydrolysis reaction. In Phase 1, it was decided to emulate the post deployment conditions by misting the surface of the samples. Because misting with

HP would be a safety hazard and the steady state pH of the surface film is essentially independent of

the initial pH, it was decided to sprinkle the surface with NaF instead of misting with HF. This

resulted in questions about the difference that might result if the initial pH were lower. To resolve

these questions, simple exposure experiments were conducted by dipping samples into different

concentrations of HF and measuring their subsequent weight changes when exposed to humid air.

5.4.1 Materials. The materials selected by the sponsors for this study were designed to represent the

broad range of aircraft structural materials which could be exposed to fire suppressant combustion bi- products. The materials selected were: 6061-T6 aluminum alloy, Inconel alloy 718, Inconel alloy

903, 410 stainless steel, titanium alloy Ti-8-1-1-1, titanium grade 5 alloy, Haynes alloy 188 and a 3501-6/AS4 epoxy-graphite composite. The compositions of these alloys are shown in Table 36.

5.4.2 Exposure Environment. According to the current military specifications, combustion product

layers could remain on component surfaces for up to one month before the surfaces are cleaned. The

effects of humidity variations, corrosion by condensate droplets induced by thermal cycling and

contamination of the surface by salts or particulate matter were not included in this study. Also, the

combustion products themselves may form fine carbonaceous particulates which could then precipitate

onto the metal surfaces and behave as cathodes stimulating corrosion.

5.4.3 Experimental Procedure. The sample geometry used for this analysis was identical to the one shown in Figure 1 (Section 5.2.2). After machining, the surfaces were glass bead blasted to remove

any remaining surface oxides or scale and to provide a consistent smooth surface finish (120 grit,

nominal). Additional preparation prior to the start of these tests consisted of an ultrasonic cleaning in acetone then in alcohol, drying with warm air and immediate weighing using the same procedure used

for the weight loss experiment.

The purpose of these experiments was to evaluate the corrosion behavior of representative aircraft materials after a deployment of a fire suppressant. The results of a study conducted by Linteris et al., revealed that the initial concentration of HF produced by combustion decomposition of these agents

ranges between 1.0 % and 10.0 % by volume (Nyden, 1994 and Linteris, 1994). As a result, one

complete set of samples was immersed in an aqueous solution containing 1.0 volume percent HF for

60 s and a second set was immersed in a similar solution containing 10.0 volume percent HF for the same time period. At the end of this exposure, the samples were extracted and hung on individual

hooks inside a sealed chamber with a constant relative humidity for 30 days. The chamber used for

this exposure was specially designed to maintain a near-saturation relative humidity through the use of

wicks along each side of the chamber. Regular hygrometer measurements indicated that the relative

humidity inside the chamber ranged between 98 % and 100 %. Mass change measurements were

taken at regular intervals during the exposure using the same weighing technique outlined previously

in this report (Section 5.2.2).

5.4.4 Results. The results of the post deployment tests are presented in Figures 19 and 20. Figure 20 is a plot of the mass change per unit exposed area as a function of the exposure time

following a 60 s immersion in 1.0 % by volume HF. In this figure, it can be seen that after approximately four days of exposure at 100 % relative humidity, there is little change in the rate of
5. COMPATIBILITY WITH METALS 173

^ 0

1

1—1 1

-

1

1—1—1

1 1—1—r

—1—!—1—

r

0 AA6Q61

—a - 410 St. Steel . - - Haynes188 . Inconel 718

- - Inconel 903 — - TI-8-1-1-1 - -A • - Ti Grade 5 — 3501-6/AS4 -

1 1 —— 8—

. _ . _ ^— • —

-8

\tr.z...

—

1

1 1 1 1 1 1 1 1 1 I 1 1 1 1 1 1 10 15 20

Post-Dip Exposure Time, Days

25 30

Figure 19. Mass change on exposure to saturated laboratory air following 60 s dip in 1 % HF.
174 5. COMPATIBILITY WITH METALS

60

40

20

T—I—I 1 II

' i "T"""i—r

-20

^—

4-

-40

-60 _l I I I ! I L

T—

r

I

I

I I I I 1—!

r —e— AA 6061

410 St. Steel — ^ - Haynes188 „ Inconel 718

- -- - Inconel 903

Ti-8-1-1-1

Ti Grade 5 - — 3501-6/AS4

4

-i I I [ I I I

'

41

10 15 20

Post-Dip Exposure Time, Days

25 30

Figure 20. Mass change on exposure to saturated laboratory air following 60 s dip in 10.0 % HF.
5. COMPATIBILITY WITH METALS 175

Table 36. Composition of alloys used for post deployment testing (mass fraction in percent)

Titanium Inconel 410 Haynes Al Titdnium Tnrnnpl Element 8-1-1 718 SS 188 6061 Grade 5 903

Al 7.60 0.51 -- - bal 6.30 4.00 C 0.01 0.03 0.14 0.11 - 0.03 0.02

Cr - 18.26 11.98 21.92 0.04 ~ —

Cu - 0.01 0.14 - 0.15 - ~

Fe 0.05 18.21 bal 2.47 0.70 0.18 41.90 Mn - 0.05 0.31 0.86 0.15 - 0.12 Mo 1.00 2.87 0.02 - - - -

Ni - 54.01 0.20 22.53 - - 37.96

Ti bal - - - 0.15 bal -

V 1.00 - - - - 4.00 - Mg - - - - 1.20 - -

Zn 0.25 OBN 0.01 0.02 0.01 Co 0.02 bal 14.70 Nb 4.90 2.88 W 13.83

weight gain for the remainder of the thirty day exposure. The graphite/epoxy composite did continue

to increase in mass, but this could be due to moisture uptake resulting from the high humidity.

Research has shown that for these materials, moisture can be transported into the bulk by the wicking

action of the fibers along the exposed edges and by solid state diffusion mechanisms (Stoudt, 1991).

Figure 20 is a similar plot showing the mass change as a function of exposure time after a 60 s immersion in 10.0 % by volume HF. During the immersion, the dissolution reactions were substantially more aggressive than those observed in the 1.0 % HF solution. As in Figure 19, the rate of mass

change slowed shortly after the coupons were removed from the HF solution. The graphite/epoxy

composite once again increased in mass, apparently the result of moisture uptake. In both cases, the

decrease in the mass change rate can be attributed to the drying out of the corrosion products. Unlike

the tests conducted in Phase I, the samples in this evaluation were not treated with any materials

which could absorb and hold the water necessary to sustain the corrosion reactions. While the air in the exposure chamber was at saturation humidity, apparently, the corrosion products on the surface of

the metal did not absorb water from the air otherwise corrosion and/or mass increases would have

been observed. As a result, there was no liquid phase present to stimulate corrosion reactions. In essence, the layer of the corrosion products on the surface have an equilibrium vapor pressure for water that is greater than that of the air in the chamber so they stayed relatively dry. This may or may

not be the case in actual service, since any kind of re-wetting of the corrosion product layer will re- establish the corrosion reactions (e.g., thermal transients causing condensation - fog, dew, rain, etc.)
176 5. COMPATIBILITY WITH METALS

and actual combustion product films will be composed of many more chemical species many of which

may absorb water.

5.5 Electrochemical Measurements

There are a number of different electrochemical techniques that can be used to measure the corrosion

rates of metals (Bard, 1980). However, because the fire suppression candidates have a high electrical

resistivity, testing in these solutions becomes very difficult. For example, HCFC-124 (CF3CHCIF),

has a resistivity of 1.5 x 10^ ohms cm. One technique that can be employed to enable electrochemical

measurements in these environments is to add a salt to the environment which will dissolve and

provide charge carriers to the electrolyte, but not alter the corrosivity of the environment (Sawyer, 1974). The choice of a salt for supporting the transfer of charge in these agents is complicated by

their low dielectric constants. For example, the dielectric constant of HCFC-124 is about 6.73 (CRC

Handbook, 1990). While the dielectric constant for the other compounds cannot be found in tables,

examination of compounds with similar symmetry and bond strengths indicates that the dielectric constant for CF3I and HFC- 125 should be around 3.5 and 6, respectively, and even higher for FC-218,

HFC-227ea and halon 1301. As a result, simple salts will not dissolve in these compounds. Several

authors have shown (Poumaghi-Azar, 1994, Bond, 1988, Park, 1994 and Kadish, 1984), that quaternary ammonium salts, on the other hand, will dissolve in low dielectric constant media such as dichloromethane and chloroform where the dielectric constants are 9.08 and 4.81, respectively.

The behavior of the candidate compounds can essentially be characterized as aprotonic in nature. As a result, their ability to ionize a supporting electrolyte is significantly reduced. The ability of a salt to be ionized in this type of medium is primarily a function of the anion (Sawyer, 1974). As the size

of the anion increases and its electronegativity decreases, ionization becomes more probable (Sawyer,

1974). Furthermore, dissociation of the supporting salt is limited by the very low dielectric constant of the replacement candidates (Sawyer, 1974). This can be overcome, to a certain extent, by choosing

a salt in which the cation and anion have a small charge to volume ratio. On the other hand, as the

size of the cation and anion increase, their mobility, and hence the solution conductivity, is reduced

(Ue, 1994). For a given charge on an ion, the size effect will have a larger affect on the dissociation constant of a salt than on the ionic mobility in a given medium. The molecular structure of the

supporting salt will also affect its mobility in a solvent. The ionic mobility, and hence the solution

conductivity, will decrease as the molecular interaction increases. Because of their small charge to volume ratio and their anions are of intermediate size, tetrabutylammonium hexafluorophosphate

(TBAPFg) and tetrabutylammonium trifluoromethanesulfonate (TBATFMS) were chosen as supporting

salts (Ue, 1994).

As explained in a previous study, in order to define the potential axis for polarization measurements, a platinum quasi-reference electrode (Pt-QRE) referenced to the half wave potential (Ej/2) of

the ferrocene/ferrocinium (Cp2Fe^''^) reduction-oxidation (redox) couple is necessary (Dante, 1994).

The nonpolar character of this outer sphere redox couple facilitates its dissolution into compounds with

a low dielectric constant. Once the potential scale is established, three electrode slow scan measurements can be made and a

Tafel extrapolation can be used to estimate the corrosion rate of the storage vessel alloys in the halon

replacement candidates.
5. COMPATIBILITY WITH METALS 177

5.5.1 Materials. The alloys examined in this analysis were: 304 stainless steel, Nitronic 40 stainless

steel, aluminum alloy 6061-T6, and Titanium alloy 15-3-3-3. The compositions of these alloys are given in Table 1, Section 5.2.2.

5.5.2 Experimental Procedure. The vapor pressure for HCFC-124, HFC- 125, and CF3I are 3.393 Pa X 10^ Pa, 11.051 Pa x 10^ Pa, and 4.390 Pa x 10^ Pa, respectively, at room temperature.

Therefore, electrochemical measurements of the corrosion rate in the liquid phase of these agents

requires the experiments be conducted in a closed pressure chamber. A stainless steel pressure vessel was modified to allow for electrical feed-throughs, Figure 21. The vessel was constructed of 0.635

cm thick 316 stainless steel and was rated for pressures of ~ 6.9 MPa. Four electrical feed-throughs were mounted through the head of the cell for performing two, three, or four electrode electrochemical

tests. A pressure gauge was mounted on the cell to measure the vapor pressure in the cell. At higher

temperatures, most of the gasses used in this study can reach very high vapor pressures so a rupture

disc was incorporated to prevent accidental over pressurization. Temperature was monitored with the use of a thermocouple mounted in the head of the cell. Electrochemical measurements were made using a commercial potentiostat. A commercial

frequency response analyzer was used to make impedance measurements and the data was collected

using commercially available software.

Before filling the electrochemical cell with one of the halon replacement candidates, 5 x

10"^ mol/L of one of the supporting salts, either TBATFMS or TBAPF^, was placed into the cell. Then,

the cell was closed and purged with dry argon to remove water vapor. After 5 min, the fill valves were closed and the cell was chilled for 30 min to 60 min. Chilling facilitated filling the cell with

agents. The agent was removed from the liquid phase in the storage bottle. Enough agent was

transferred to the electrochemical cell to form a 3 cm deep (250 ml) liquid phase in the bottom of the

cell, i.e., 352 g, 353 g, and 555 g of HCFC-124, HFC-125, and CF3I, respectively. In all tests, a large area platinum-rhodium counter electrode and a platinum wire quasi-reference electrode (Pt-QRE) were

used. The Pt-QRE electrode was placed between the counter and working electrodes approximately

2.5 cm from the working electrode. The electrodes were attached to the electrical feed-throughs using

heat shrink tape. For the working electrodes, electrical contact was made using clips.

The effectiveness of both TBATFMS and TBAPFg in lowering the resistivity of HCFC-124,

HFC-125, and CF3I was measui|ed using electrochemical impedance spectroscopy (EIS). The cell resistance was obtained by noting the value where the curve intersected the real impedance axis on the

Nyquist plot. This value was converted to a resistivity by multiplying by the electrode area and

subsequently dividing by the distance between the working and reference electrode. Unfortunately, it was not possible to get sufficient quantities of either of these salts to dissolve in either HFC-227ea or halon 1301 to enable electrochemical measurements.

Tests were run to elucidate the electron transfer kinetics of Cp2Fe^''''' in HCFC-124, HFC-125 and

CF3I. The tests were run at room temperature. A Pt disk working electrode (area = 0.00317 cm^) was

prepared by encapsulate a platinum wire in glass and polishing the electrode face to a 1 \xm finish. The oxidation and reduction waves of ferrocene in each experiment were measured with respect to the

Pt-QRE. The potential scale was then normalized to the half-wave potential of Cp2Fe^^''' redox couple.

In subsequent tests, the corrosion rate of 304 (area = 0.00785 cm^) and Nitronic 40

(area = 0.01461 cm^) stainless steels, aluminum 6061-T6 (area = 0.01954 cm^), and Ti 15-3-3-3 were

studied in the three agents using slow sweep voltammetry (scan rate = 0.5 mV/s) and Tafel extrapolations. In all cases, wires of each alloy were encapsulated in glass and the electrode surfaces were

polished to a 1 pm finish. All of the alloys except 304 stainless steel were extruded and heat treated

from stock samples. It should be noted that the encapsulation process involved temperature in excess

of 500 °C for short periods of time which may have modified the metallurgy of the wires slightly.
178 5. COMPATIBILITY WITH METALS

Electrode

(Sample)

Figure 21. Electrochemical cell for making measurements in high vapor pressure low

conductivity electrolytes.
5. COMPATIBILITY WITH METALS 179

Table 37. Resistivity and electron transfer rate constants of Cp2Fe and oxidation rates in various

replacement candidates

Agent Supporting Resistivity, p

Salt Ohms-cm cm/s

HCFC-124 TBATFMS 80 0.0020

HCFC-124 TBAPFg 79 0.0013

HFC- 125 - TBATFMS 44 0.0024

HFC- 125 TBAPFg 740 2.5x10-^

CF3I TBATFMS 748 B-exlO"*

CF3I TBAPFg

5.5.3 Agent Resistivity. Electrochemical impedance spectroscopy (EIS) was used to measure the

resistivity of the agents as a function of added salt. Values for the resistivity of the agent/salt

combinations are shown in Table 37. Several conclusions can be drawn from this data. The resistivity

of the agent/supporting salt electrolyte is related to several factors i.e., the ionization of the salt, the

mobility of its ions and the dissociation constant of the salt in the medium which regulates the number

of ions available as charge carriers from the solvated salt. Furthermore the size of the two salts is approximately the same (Ue, 1994) with the TFMS anion being slightly larger. Analysis of Table 37

reveals that the resistivity of TBAPFg and TBATFMS in HCFC-124 is similar. The charge to volume

ratio of the TFMS anions are slightly smaller than that of the PFg" anions and hence dissociate to a

slightly larger extent in HCFC-124. In order for the resistivity of the solutions containing the different

salts to be similar, the mobility of the PF^" ions must be higher than that for the TFMS anions.

The situation in HFC- 125, however, is very different. The resistivity of the TBAPFg electrolyte is an order of magnitude higher than that of the TBATFMS electrolyte. With the verified assumption

that all of the salts are dissolved and knowing that the ions are of the same general size, the difference

in resistivity must be attributed to a difference in the number and mobility of the charge carriers in HFC- 125. The TFMS ion has a structure that more closely resembles the structure of HFC- 125 than

PFg' does, thus facilitating the ability of TBATFMS to move through the medium in response to an

applied field. The charge to volume ratio of the TFMS anions are slightly smaller than that of the

PF^" anions and hence dissociate to a slightly larger extent in this medium. These two effects combine to give a much lower resistivity for TBATFMS in HFC- 125.

The TFMS ions more closely resemble the molecules of HFC- 125 than HCFC-124 implying a weaker intermolecular force and hence a higher mobility in HFC-125. The dielectric constant of

HFC- 125 is slightly lower than that for HCFC-124. Thus, the number of charge carriers would be

greater in HCFC-124. From Table 37, the resistivity of the TBATFMS solution is slightly lower in HFC-125 suggesting that the mobility of the charge carriers is the major factor contributing to the

lower resistivity of TBATFMS in HFC-125. That is, even though the number of charge carriers is lower in HFC-125, they are more mobile than those in HCFC-124 and result in a lower resistivity in HFC-125. Table 37 illustrates the fact that the resistivity is an order of magnitude higher in HFC-125
180 5. COMPATIBILITY WITH METALS

relative to HCFC-124 containing TBAPF^. The number of charge carriers in HCFC-124 would be

greater than that in HFC- 125 because of the higher dielectric constant of HCFC-124. This implies that the mobility of the charge carriers in HCFC-124 would be similar to or slightly higher than in HFC- 125.

HCFC-124 and HFC-125 containing 5 x 10"^ mol/L of TBATFMS have similar resistivities. The

resistivity in CF3I, however is an order of magnitude higher. Since the dielectric constant of CF3I is much lower than that in the other salts, the number of charge carriers in this medium are much less. This fact alone accounts for the much increased resistivity. TBAPF^ does not dissolve in CF3I. The

slightly higher polarity of this molecule compared to TBATFMS is enough to prevent its dissolution.

5.5.4 Defining a Reference Potential. Figure 22 displays a 0.1 V/s CV scan from -2.5 V to 1 V

versus a Pt-QRE in HCFC-124 and HFC-125 containing 5 x

10'"* mol/L of ferrocene and 5 x

10'^

mol/L of TBAPFg. These scans represent a typical CV in the above environments. The current was

measured at a glass encapsulated Pt electrode. The potential axis has been normalized to E,/2(Cp2Fe°^"'), which appeared at 0.234 V and 0.481 V (Pt-QRE) in HCFC-124 and HFC-125

respectively. It can be seen from the figure that a major reduction occurs in both environments at approximately -2.55 V (Cp2Fe^^'^). This suggests that the peak at -2.55 V (Cp2Fe^''^) is a reduction

process associated with the supporting salt. This peak is in the vicinity of the onset of reduction for

the Bu4N'^ cation in dimethylformamide and acetonitrile (Kadish, 1984). At approximately

-1.50 V (Cp2Fe*^''"^) a slight change in the curve can be seen for both environments. In this figure, the

peak height is small, but it varies in magnitude each time the test cell is reassembled. The peak is thought to be a result of water contamination. This is supported by other experiments that demonstrate

that the magnitude of the peak changes when controlled amounts of H2O are injected into the cell. In

this study, there was no evidence of any oxidation occurring at 1 .0 V (Pt-QRE) or 0.77 V (Cp2Fe^^''")

leaving at least a 2.27 V window where no reactions occur on Pt. By averaging the peak separation

between ferrocene and the BU4N"'" and ferrocene and the water peak for scan rates up to 0.5 V/s in each environment, the following results are obtained: E,/2(Cp2Fe^'^) - E(Bu4N+) = 2.55 V ± 0.002 V

and 2.56 V ± 0.014 V, respectively, Ej/2(Cp2Fe^'-*-) - E ^20 = l-^O V and 1.50 V for HCFC-124 and

HFC-125 containing TBAPF^, respectively.

Figure 23 displays a 0.1 V/s scan from -3 V to 1 V (Pt-QRE) in HCFC-124, HFC-125, and CF3I

containing 5x10""^ mol/L of ferrocene and 5 x

10'^ mol/L of TBATFMS. These are typical results for the above environments. As before, the current was measured at a glass encapsulated Pt electrode

and the potential has been normalized to Ej/2(Cp2Fe^^"^) which appeared at 0.19 V, 0.17 V, and

0.35 V (Pt-QRE) for HCFC-124, HFC-125, and CF3I, respectively. At 2.48 V (Cp2Fe°^"') and

2.55 V (Cp2Fe^^"^) for HCFC-124 and HFC-125, respectively, a major reduction peak can be seen. The peak separation between the ferrocene half wave and the major reduction peak is similar to the case in which TBAPF^ was the supporting salt lending further credence to the belief that the major

reduction peak is a result of the reduction of Bu4N"*'. A second peak at 1.503 V (Cp2Fe^^''") and

1.458 V (Cp2Fe*^^+) for HCFC-124 and HFC-125, respectively, is associated with water. This peak is much more pronounced for the curves in Figure 22 than for those is Figure 23. This may be a result of contamination of the TBATFMS. In several cases, the water peak masks the onset of the large

reduction peak of the salt and care must be taken when determining the half wave potential of

ferrocene in the situation where no ferrocene is present. Oxidation of the electrolyte was not seen for

several hundred mV above the half wave potential of ferrocene yielding a usable potential window of

at least 2.0 V anodic of the water peak. Peak separations for scans up to 0.5 V/s were averaged and

the results are as follows: Ej/2(Cp2Fe°^+) - E(Bu4N'-) = 2.49 V ± 0.008 V and 2.55 V ± 0.008 V for
COMPATIBILITY WITH METALS 181

Figure 22. 0.1 V/s voltametric sweep on a Pt-QRE in HCFC-124 and HFC-125 using 5 x lO""^ mol/L of ferrocene and 5 x

10"^ mol/L of TBAPF^ (1 out of 50 data points shown).
182 5. COMPATIBILITY WITH METALS

Figure 23. 0.1 V/s voltametric sweep on a Pt-QRE in HCFC-124, HFC- 125, and CF3I using

5 X

10""^ mol/L of ferrocene and 5 x

10"^ mol/L of TBATFMS (1 out of 50 data

points shown).
5. COMPATIBILITY WITH METALS 183

HCFC-124 and HFC- 125 respectively, Ei/2(Cp2Fe"'+) - Eh2o = 1.50 V ± 0.009 V and

1.47 V ± 0.008 V for HCFC-124 and HFC-125, respectively.

The situation with TBATFMS in CF3I is different from this salt in the other two environments.

As can be seen in Figure 24, only one major reduction peak is observed at 1.095 V (Cp2Fe^^'^). The

peak is very large and the peak maximum is never detected. The addition of water has no affect on

the voltammograms. It can not be ruled out, however, that this large peak is a result of water

contamination. It is known that water is often found as a contaminant of CF3I. The usable window

for CF3I containing TBATFMS is at least 1.5 V. Peak separations for scans up to 0.5 V/s were

averaged yielding Ei/2(Cp2FeO^^) - E,„k„,^„ = 1.103 V ± 0.008 V.

In subsequent experiments studying the corrosion of alloys in HCFC-124, HFC-125, and CF3I,

there was no ferrocene added to the cell. Ej/2(Cp2Fe^^"^) was then defined by noting the position of

the salt reduction, finding the appropriate value for the separation of the reduction peak and the

ferrocene peak listed in Table 38, and realizing that ferrocene is this much more anodic than the

Bu^N"^ reduction peak. In several of these experiments, however, the reduction peak associated with

H2O masked the onset of reduction of the salt. In this case, Ej/2(Cp2Fe^'''^) was defined with reference

to the separation of ferrocene and the water peak. In the case of CF3I, the half wave potential of

ferrocene was measured using the peak separation of the large reduction peak and ferrocene, listed in Table 38, in all cases.

5.5.5 Cp2Fe*^^'*' Electron Kinetics. Cyclic voltammograms of Cp2Fe^''^ oxidation and reduction peaks

on a Pt disk electrode in HCFC-124, HFC-125, and CF3I at 0.1 V/s with TBATFMS as the supporting

salt are shown in Figure 24. Using TBAPF^ as the supporting salt resulted in similar scans. Peak

separation (E^ -E^,) and the separation of the anodic peak potential and the half-wave potential of

Cp2Fe°''"^ (Eg - Ey2) as a function of scan rate were used to estimate the electron transfer rate constant,

k (Bard, 1980). Values for k^ in all of the agent/supporting salt combinations is shown in Table 37.

k^ was calculated assuming a diffusion coefficient of 1 x 10"^ cm^/s and a = 0.5. It should be noted

that no IR compensation was used in the current experiments. The very large uncompensated

resistance would thus result in a lowering of the k^ values as the scan rate increased. Compensation

for IR drop was attempted, but efforts to do so resulted in oscillations at entered resistance values that were much lower than their actual values.

The k^ values found in the literature for Cp2'PeP''^ in acetonitrile, dichloromethane and chloroform

using similar sized electrodes are k° ~ 9 x lO''^, 6 x 10"'^, and 3 x 10'^ cm/s, respectively

(Poumaghi-Azar, 1994 and Kadish, 1984). The k° values of Cp2Fe°'''^ obtained in this study are approximately an order of magnitude lower, in the case of HCFC-124 with both salts and HFC-125

with TBATFMS. It should be mentioned that k*^ values for Cp2Fe°^'^ of the magnitude seen in this study have been observed in electrolytes with similar dielectric constants using electrodes of similar

proportions (Bond, 1988). In the case of HFC-125 with TBAPF^ and CF3I with TBATFMS, the k*^ values are about 1.5 orders of magnitude lower. In a study by Kadish et ai, there appears to be an

inverse relationship between the solution resistivity and the value of k^ that is not related to uncompensated resistance (Kadish, 1984). This trend is verified in the data for both TBATFMS and

TBAPF..

"

It is interesting to note that in HCFC-124, as is the case for the resistivity, the k values appear to be unaffected by the salt used. In HCFC-124 with both salts and HFC-125 with TBATFMS, k^

decreases slightly with scan rate, at higher scan rates, indicating that IR does play a minor role in the

low k^ values obtained in this study. It was also noted that the k*^ values decrease with scan rate to a

greater extent in HFC-125 using TBAPF^ and CF3I using TBATFMS. This implies that the IR effect

is much more noticeable in those two environments as would be expected given the values of the

resistivity in Table 37.
184 5. COMPATIBILITY WITH METALS

Figure 24. Ferrocene oxidation and reduction peaks on a Pt-QRE in HCFC-124, HFC-125, and

CF3I. (1 out of 50 data points shown).
5. COMPATIBILITY WITH METALS 185

Table 38. Reference peak values

Agent Supporting

Salt

E,/2(CP2Fe°'*)- E,^(TBA)

Volts

E,/2(CP2Fe«'^)-E„2o

Volts

HCFC-124 TBATFMS 2.49 ± 0.008 -1.50 ± 0.009

HCFC-124 TBAPF^ 2.55 ± 0.002 -1.50

HFC- 125 TBATFMS 2.55 ± 0.009 1.47 ± 0.008

HFC- 125 TBAPFg 2.56 ± 0.014 -1.50

CF3I TBATFMS 1.103 ± 0.008

CF3I TBAPFg

Irp = large reduction peak for tetrabutylammonium (TBA)

Based on solution resistivity and k values in the various media, it appears that alternative fire suppressant solutions containing 5 x

10"^ mol/L of TBATFMS and 5 x

10""* mol/L of ferrocene are the

most appropriate systems for measuring corrosion rates of storage container alloys.

5.5.6 Corrosion Rates of Storage Vessel Alloys in HCFC-124, HFC-125, and CF3I. Figures 25,

26, and 27 show Tafel plots of 304 stainless steel, Nitronic 40, Al 6061 T6, and Ti 15-3-3-3 in HCFC-124, HFC-125, and CF3I, respectively. In all cases, the free corrosion potential (FCP) was very

difficult to obtain and moved cathodically from a more or less stable value when potentiodynamic

scans were performed. This cathodic movement was observed regardless of the direction of the scan.

Fairly reproducible FCPs and scans were made by preconditioning the working electrode prior to a

scan. A potential of -0.1 V was applied to the working electrode for 10 s prior to each scan. The

potential was then allowed to return to a steady state over approximately 30 min. The sliding FCP

was less severe in HFC-125. Therefore, scans were begun at -250 mV below the FCP and ended at 500 mV over the FCP. In the other two environments, where the FCP is less stable, potentiodynamic

scans were begun at -100 mV below the FCP and terminated at 1.5 V of the FCP.

Table 39 shows the results of Tafel (Bard, 1980) extrapolations of the curves in Figures 25, 26,

and 27. The extrapolation was based primarily on the anodic curves. It can be seen from the results

that the corrosion current density in alternative fire suppressants is extremely low. The values are on

the order of those obtained for passivated metals in aqueous solutions. No evidence of transpassive

behavior was observed for any of the metals in any environment, even at high overpotentials. The low

current densities may be a result of both a protective air formed oxide as well as the low dielectric constant of the environment. The presence of oxidation and reduction precursors and by products, i.e., H^, OH", and metal ions, are highly unfavorable in the low dielectric constant media.

Corrosion rates of the four alloys are reported in the third column of Table 39. The estimated

weight loss values were calculated assuming the current densities in Table 39 and that the major

dissolution phases are Fe/Fe^+ for Nitronic 40 and 304 SS, A1/A1^+ for Al 6061, and Ti/Ti^^ for Ti

15-3-3-3. In the fourth column of Table 39, weight change data from 25 day immersion tests at 20 °C

is presented. All but two of the weight change data are positive indicating that film formation had
186 5. COMPATIBILITY WITH METALS

Figure 25. Potentiodynamic scans of 304SS, Nitronic 40, Al 6061-T6, and Ti 15-3-3-3 in HCFC-124 containing 5x10"^ mol/L of TBATFMS (1 out of 50 data points shown).
COMPATIBILITY WITH METALS 187

0.1

SS 304

Nit 40

AA 6061

Ti 15-3-3-3

—

r

I

To

c

o

o

0.01

0.001

0.0001

•0.8

I I -0.4 0.0 0.4 0.8

Potential (V vs. Ferrocene/Ferrocinium Eq.)

1.2

Figure 26. Potentiodynamic scans of 304SS, Nitronic 40, Al 6061-T6, and Ti 15-3-3-3 in HCFC-125 containing 5 x 10"^ mol/L of TBATFMS (1 out of 50 data points shown).
188 5. COMPATIBILITY WITH METALS

Figure 27. Potentiodynamic scans of 304SS, Nitronic 40, Al 6061-T6, and Ti 15-3-3-3 in CF3I

containing 5x10'^ mol/L of TBATFMS (1 out of 50 data points shown).
5. COMPATIBILITY WITH METALS 189

Table 39. Polarization for various alloys in replacement candidates as compared to weight change

values after a 30 day immersion

Agent Alley ^Corr

(Volts)

ICorr (A/cn?)

Estimated aM/A

(g/m^ day) Measured aM/A

(g/m^ • day)

HCFC-124 Nitronic 40 -0.066 1.18x10"^ -29.5x10"^ —

HCFC-124 304 Stainless Steel -0.314 3.97x10'^ -9.93x1

0"3 —

HCFC-124 Al 6061-T6 -0.401 3.67x10"^ -2.96x10"^ —

HCFC-124 Ti 15-3-3-3 -0.435 2.52x10"^ -5.40x10-^ —

HFC- 125 Nitronic 40 -0.388 4.15x10-^ -10.38x10"^ 2.79x10-^

304 Stainless Steel -U.jUo 4. / jX iU -H .oZXlU O. /DXlU HFC- 125 Al 6061-T6 -0.453 1.50x10-^ -1.21x10"^ 10.21x10"^ HFC- 125 Ti 15-3-3-3 -0.320 5.58x10-^ -11.96x10"^ 0.89x10-^

CF3I Nitronic 40 -0.532 3.72x10"^ -9.30x10"^ -1.01x10"^

CF-,1 304 Stainless Steel -0.460 1.16x10-^ -2.90x10-^ -0.39x10-^

CF3I Al 6061-T6 -0.501 2.11x10'-^ -1.70x10"^ 1.09x10-^

CF3I Ti 15-3-3-3 -0.531 6.63x10"'' -1.42x10"^ 0.58x10"'

occurred. Attempts to remove this film and measure an actual metal dissolution rate were unsuccessful. Therefore, the conclusions that can be drawn from comparisons of columns 3 and 4 in Table 39

are limited. However, it is interesting to note that all but one of the corrosion weight loss data values

obtained using electrochemical techniques are within one order of magnitude of immersion test values.

Furthermore, the absolute value of all but one of the electrochemical data are greater than the absolute

values of the immersion test values. This would indicate that the electrochemical data would allow for a worse case design criteria based on the higher corrosion values. It is believed that the electrochemical data is more realistic in this study because a true dissolution rate was not obtained for the

immersion tests. According to the electrochemical data, it appears that Al 6061-T6 has the best

overall performance in the agents studied. It is interesting to note that the Al alloy has the highest

weight gain in both HFC- 125 and CF3I. This is probably indicative of the formation and growth of a

protective layer.

5.5.7 Corrosion Rates of Metals in Other Halon Replacement Agents. From the data in Table 39,

it appears that, in general, the corrosion rates for the various metals are lower in CF3I than for the

other two replacement environment. The lower dielectric constant for this medium is believed to be

the reason for this lower value. Typical charged species involved in corrosion have a low probability

of being dissolved in a low dielectric constant media because of their low charge to volume ratio resulting in a lower corrosion rate. The other halon replacement agents under consideration, FC-218 and HFC-227ea have lower

dielectric constants than those examined in these electrochemical experiments. In support of this assumption, the salts used in this study did not dissolve in FC-218 or HFC-227ea. Thus, it is proposed that the corrosion rate of storage vessel alloys in these media will be less than those

measured here because the solubility of aggressive species and corrosion products will be reduced.
190 5. COMPATIBILITY WITH METALS

5.6 Conclusions

The primary objective of this investigation was to determine whether any of the candidates selected as

potential replacements for halon 1301 should be eliminated from further consideration because of

unacceptable corrosion rates of the metals used in the fire suppressant storage and distribution systems on aircraft. In Phase II of this study, the remaining candidates from the screening process (Phase I) were examined for longer times and over a broader range of environmental conditions. In addition,

this phase examined both the corrosivity of a representative post deployment deposit on aircraft structural materials and the use of supporting salts which resolved the resistivity of the suppressants

thereby enabling electrochemical measurements of corrosion rates. As in the first phase of this investigation, no single candidate was found to cause serious

corrosion problems with all of the alloys and no serious concerns were raised for all of the alloys in any of the agents. However, some agents did consistently demonstrate poorer performance than others or yielded results which indicate further testing is warranted before they should be placed into service. In particular, CF3I demonstrated higher mass change rates for most of the alloys than the other agents

requiring a mass change scale at least one order of magnitude larger than halon 1301 to plot the

results. The other agents demonstrated mass change rates that essentially the same or less than those

observed for halon 1301. Visual examination of the samples after the exposure tests revealed that AISI 4130 developed small pits during exposure to HFC-227ea, FC-218, and CF3I with the largest pits

being observed in CF3I. Small pits were also observed in Nitronic 40 and Al 6061 during exposure to HFC-227ea. The only agent to demonstrate any conclusive evidence of environmental induced

fracture during the slow strain rate tensile tests was CF3I which caused cracking of two of the three

samples tested at 150 °C.

Electrochemical measurements could only be made in three of the agents: HCFC-124, HFC- 125,

and CF3I. Presumably, the corrosion rates in the others would be lower than these three due the same

factors that inhibit the measurement of corrosion rate (high resistivity and dielectric constant). Of

these three, the lowest corrosion rates were observed in CF3I which is in contrast to the immersion test results. This may be due to the formation of protective surface films that slow corrosion resulting in a lower average corrosion rate over the long exposure times in the other agents or it could be an

indication that the corrosivity observed for CF3I is the result of second phase contaminants present in the agent as supplied and the fact that the electrochemical experiments used smaller volumes of agent

and very small area samples which would result in a much lower probability of contaminating the

surface with an aggressive species such as a halogen acid. If contaminants are responsible for the poor

performance of CF3I in these tests, then removal of these impurities could make this agent acceptable

for use. The results of this investigation demonstrate that it may be possible to use any of these agents in

aircraft fire suppression systems, but before CF3I is used, a better understanding of the factors

influencing its corrosivity will be needed. Also, this conclusion is based on the quality of the agents

as supplied for testing and changes in the impurity content of any of the agents could result in dramatic shifts in the corrosivity. Assuming that the performance of the CF3I was the result of

trapped impurities, then the presence of similar impurities in any of the agents should result in similar

levels of performance.
5. COMPATIBILITY WITH METALS 191

5.7 Acknowledgments

The authors would like to thank Dr. Thomas Moffat and Mr. Carlos Beauchamp of NIST for their

technical assistance and helpful suggestions during the design and development of the electrochemical

techniques. We also would like to thank Mr. George Hicho and Mr. Richard Harris, also of NIST, for

their insightful reviews. We acknowledge Dr. Kumar Jata and Mr. Thomas Naguy of the WrightPatterson Materials Laboratory for their assistance and recommendations in the selection of the alloys

used in this study. The authors especially wish to thank Ms. Joan Teagle, also of NIST, for her

dedication and invaluable help with the editing of this document.

5.8 References

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Procedures, Philadelphia, PA, 1993.

ASTM, Standard Guide for Applying Statistics to Analysis of Corrosion Data, Annual Book of

ASTM Standards, Section 3, Metals Test Methods and Analytical Procedures, Philadelphia, PA, 1993.

ASTM, Standard Practice for Laboratory Immersion Corrosion Testing of Metals, Annual Book of

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Engineermg, 275-324, McGraw-Hill, New York, 1986.

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ASM International, 1987.
192 5. COMPATIBILITY WITH METALS

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M. R., in Evaluation of Alternative In-Flight Fire Suppressants for Full Scale Testing in Simulated

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at Pt, Au and Glassy Carbon Electrodes in Chloroform," Electrochimica Acta 39, 953, 1994.

Ricker, R. E., Stoudt, M. R., Dante, J. F., Fink, J. L., Beauchamp, C. R., and Moffat, T. R, in Evaluation of Alternative In-Flight Fire Suppressants for Full Scale Testing in Simulated Aircraft

Engine Nacelles and Dry Bays, NIST Special Publication 861, NIST, 1994.

Sawyer, D. T., and Roberts, J. L., Experimental Electrochemistry for Chemists, John Wiley and

Sons, Inc., New York, 1974.

Stoudt, M. R., Escalante, E., and Ricker, R. E., "The Influence of Moisture and Pressure on the

Mechanical Properties of a Glass-Epoxy Matrix Composite and a Graphite-Epoxy Matrix Composite,"

Environmental Effects on Advanced Materials, San Diego, CA, 1991.

Stoudt, M. R., Fink, J. L., and Ricker, R. E., "The Environmentally-Induced Failure Susceptibility

of Fire Suppressant Storage Container Alloys In Replacement Candidates for Halon 1301," In Press,

1995.

Stoudt, M. R., and Ricker, R. E., "Characterization of the Hydrogen Induced Cold Cracking

Susceptibility at Simulated Weld Zones in HSLA-100 Steel," NISTIR 5408, NIST, 1994.

Stoudt, M. R., Fink, J. L., and Ricker, R. E., "Evaluation of the Stress Corrosion Cracking

Susceptibility of Fire Suppressant Storage Container Alloys in Replacement Candidates for Halon

1301," Paper presented at the Tri-Service Conference on Corrosion, Orlando, FL, 1994.

Ue, M., "Mobility and Ionic Association of Li and Quaternary Ammonium Salts in Propylene

Carbonate and -Butyrolactone," J. Electrochem. Soc. 141, 3336, 1994.

Zotikov, V. S., Bakhmutova, G. B., Bocharova, N. A., and Semenyuk, E. Y., "Corrosion of Al

and its Alloys in Hydrofluoric Acid," Prot. Met. 10, 2, 154-156, 1974.
5. COMPATIBILITY WITH METALS

Appendix A. Mass Change Measurements at 150 °C

193

Sample Test Environ. AM/Area AM/Area AM/Area AM//^rea 180 AM/Area 365 Number ju uays DU Days g/m 1 zu uays Days g/m^ Uays g/m

g/m g/m

NIT40/13 HFC-227ea 0.11681 0.15692 0.14866 0.14866 —

iNl 14U/14 nru-zz/ea u. 1 /oyo A 1 QO'7Q A A OAAT U.4ZUUJ A 1 yl Ayl A

U.14U4U

Ni IT/1 n/ 1 < nrL-zz /ea A 1 A1QA U. 14jy4 U. 1 jZ 1 J n 1 8878 U. 1 08 /o A 1 ^ 8 1 A

NIT40/16 FC-218 0.17934 0.75866 4.43396 3.26942

MTT/in/l 7 7 1

8

rC-Z 1 o U. lUO jD V.J J 1 ou 4.ZJZZU J.Z84 10

iN 1 1 40/ 1 8 rL--Z 1 o U. I J5U4 U. J jOjZ 7 787A7 J. /8ZD / 7 8A8AQ Z.8U8l)y

INI 14U/iy naion uui u. 1 / lUo n 7*3/170 A •JAC CO A '3AAO/1 ATI A-J 1 NIT40/20 halon 1301 0.12743 0.18288 0.27845 0.34924 0.36694 NIT40/21 halon 1301 0.12153 0.19822 0.26075 0.29143 0.31503

NIT40/22 HFC- 125 0.02124 0.06135 0.06371 0.12389 0.09675 NIT40/23 HFC- 125 0.07197 0.09439 0.11563 0.11563 0.11091 NIT40/24 HFC- 125 0.06253 0.06843 0.09557 0.08023 0.09675

NIT40/25 CF3I 16.16778 16.47808 17.56002 NIT40/26 CF3I 15.40322 16.09344 8.26972 NIT40/27 CF3I 17.85381 18.71158 9.64663

Sample Test Environ. AM/Area AM/Area AM/Area AM/Area 180 AM/Area 365 Number 150 °C 30 Days 60 Days g/m^ 120 Days Days g/m^ Days g/m^

g/m^ g/m^

6061/13 HFC-227ea 0.42492 0.63796 1.00595 1.25245 6061/14 HFC-227ea 0.53056 0.58925 0.73011 0.98834

6061/15 HFC-227ea 0.44957 0.59629 0.52000 0.72658

6061/16 FC-218 0.47774 0.73363 9.49842 7.84688

6061/17 FC-218 0.35097 0.67846 0.49652 0.17607

6061/18 FC-218 0.29228 0.55051 1 .04469 0.55990

6061/19 halon 1301 0.42609 0.97074 1.26419 1.34635 1.24775 6061/20 halon 1301 0.36153 0.56812 0.94609 0.98600 0.90852

6061/21 halon 1301 0.36153 0.52234 0.95665 0.92026 0.87801

6061/22 HFC- 125 0.25472 0.47422 0.93317 0.99656 0.94961

6061/23 HFC- 125 0.23711 0.48243 0.91087 0.84044 0.96839

6061/24 HFC- 125 0.26411 0.51647 0.79232 0.72893 0.89444

6061/25 CF3I 2.17389 5.46523 8.58872 6061/26 CF3I 0.41201 2.32061 4.04376

6061/27 CF3I 0.27584 2.76900 5.07201
194 5. COMPATIBILITY WITH METALS

Sample Test Environ. AM/Area AM/Area AM/Area AM/Area 180 AM/Area 365 Number 150 °C 30 Days 60 Days g/m^ 120 Days Days g/m^ Days g/m^

, 2 g/m^

304/13 HFC-227ea 0.17679 0.22831 0.19318 0.19201 304/14 HFC-227ea 0.14752 0.20957 0.20372 0.18850 —

304/15 HFC-227ea 0. 1 3464 0.17913 0.19552 0.14518 —

304/16 FC-218 0.23182 0.49993 7.36548 4.32727 ... 304/17 FC-218 0.14284 0.33953 7.80453 4.95715 —

304/18 FC-218 0.16625 0.30441 6.66301 4.37644

304/19 halon 1301 0.26811 0.25992 0.30909 0.37348 0.33602

304/20 halon 1301 0.20840 0.22948 0.28099 0.31846 0.32665

304/21 halon IJOl U.l /445 ft TAT CC u.i6iuy 0.33016 0.31612

304/22 HFC- 125 0.04566 0.10303 0.16157 0.24235 0.19904

304/23 HFC- 125 0.09601 0.14752 0.15455 0.15103 0.16625

304/24 HFC- 125 0.07142 0.13230 0.15220 0.16859 0.14401

304/25 CF3I 9.23290 14.76492 32.35966

304/26 CF3I 11.33448 20.14005 24.00251 304/27 CF3I 10.09344 16.99412 15.15245

Sample Test Environ. AM/Area AM/Area AM/Area AM/Area 180 AM/Area 365 Number 150 °C 30 Days 60 Days g/m^ 120 Days Days g/m^ Days g/m^

g/m^ g/m^

Ti/13 HFC-227ea 0.25954 0.37657 0.28503 0.02665

Ti/14 HFC-227ea 0.17728 0.25491 0.19466 0.21088

Ti/15 HFC-227ea 0.20393 0.30357 0.29778 0.29546

Ti/16 FC-218 0.30705 0.88986 0.98603 0.80528

Ti/17 FC-218 0.26418 0.38468 0.89566 0.66392

Ti/18 FC-218 0.22478 0.33717 1.56074 1.32205

Ti/19 halon 1301 0.17728 0.26418 0.34181 0.35571 0.41249

Ti/20 halon 1301 0.19582 0.24796 0.38816 0.44957 0.40322

Ti/21 halon 1301 0.19813 0.26765 0.40438 0.46695 0.44957

Ti/22 HFC- 125 0.08111 0.18307 0.19118 0.21899 0.29199

Ti/23 HFC- 125 0.06836 0.13325 0.14715 0.17959 0.17264

Ti/24 HFC- 125 0.12166 0.20045 0.19118 0.19118 0.19466

Ti/25 CF3I 4.89308 3.14696 6.10969

Ti/26 CF3I 4.51188 3.35089 4.15385

Ti/27 CF3I 4.46322 3.86882 5.35540
5. COMPATIBILITY WITH METALS 195

Sample Test Environ. AM/Area AM/Area

Number 1 JU 60 Days g/m^ iDU uays g/m

4130/13 HFC-227ea -24.06626 -38.42012 4130/14 HFC-227ea 0.32681 -2.49535

A 1 1C\I1 ^ 41 JU/1

J

-U. ioZo /

4130/16 FC-218 3.79199 4.33117

4130/17 FC-218 2.96846 3.53360

41 jU/15 TZC 0 1 C ry^-L 1

0

1U.301 jU

4130/19 halon 1301 1.58098 2.68766

4130/20 halon 1301 2.08241 2.81626

H 1 J\JI £. 1 h,n1nn nOl 1.99982

4130/22 HFC- 125 1.46181 -1.02528 4130/23 HFC- 125 0.85538 0.75981 4130/24 HFC- 125 0.56042 0.46485

4130/25 CF3I -65.49613 -106.61335 4130/26 CF3I -74.17617 -57.24083 4130/27 CF3I -33.08845 -53.94674

Sample

Number

Test Environ. 150 °C AM/Area

90 Days g/m^

321/04

321/05 321/06

halon 1301 halon 1301

halon 1301

1.38949 2.78602

1.49176

321/10

321/11 321/12

HFC- 125 HFC- 125 HFC- 125

0.01411 0.01411 0.00823

Sample Test Environ. AM/Area

Number 150 °C 90 Days g/m^

CDAllO/04 halon 1301 1.52769 CDA 110/05 halon 1301 1.77123 CDA110/06 halon 1301 1.43710

CDAllO/10 HFC- 125 0.05824 CDAl 10/11 HFC- 125 0.12706 CDAl 10/12 HFC- 125 0.02647
196 5. COMPATIBILITY WITH METALS

Sample

Number

Test Environ. 150 °C AM/Area

90 Days g/m^

CDA 172/04 CDA 1 72/05 CDA 172/06

halon 1301

halon 1301

halon 1301

5.89103 2.10562

3.28194

CDA172/10

CDA172/11

CDA 172/ 12

HFC- 125 HFC- 125 HFC- 125

2.47205 2.35912 2.45440

Sample

Number

Test Environ. 150 °C AM/Area

90 Days g/m^

355/04

355/05

355/06

halon 1301 halon 1301

halon 1301

1.04301 1.42683 1.90702

355/10

355/11 355/12

HFC- 125 HFC- 125 HFC- 1250.00459

0.03901 0.01893

Sample

Number

Test Environ. 150 °C AM/Area

90 Days g/m^

1625/04

1625/05

1625/06

halon 1301

halon 1301 halon 1301

0.62422

0.67142 0.68322

1625/10

1625/11 1625/14

HFC- 125 HFC- 125 HFC- 125

-0.01534

-0.03186 -0.07434
5. COMPATIBILITY WITH METALS

Appendix B. Mass Change Measurements at 20 °C

197

Sample

Number

Test Environ. 20 °CAM/Area

30 Days

AM/Area

60 Days

g/m^

AM/Area

120 Days

girt?

AM/Area

180 Days g/m^

AM/Area

365 Days g/m^

XTTX/in/f\l IN114U/U1 NIT40/02

NIT40/03

HFC- 125 HFC- 125

U.Uoj / / 0.03068 0.00236

vVoj 1

1

-0.01652 0.01888

U.l jVzo

0.06371 0.06017

C\ T 1 TIC

0.12271 0.12979

A A AQA1 "3

0.09793

N114U/(J4 NIT40/05

NIT40/06

HrC-zi /ea HFC-227ea

HFC-227ea

U.Ul / /U -0.01062 -0.02950

A A<C<CA"7 U.UooU / 0.03894

0.03422

A AATni

u.uy /yj 0.05545 0.06253

A AAA 1 1 A ACA-T*! u.uyyi 1 U.U5U73

0.07905

Nll4(J/U/ NIT40/08

NIT40/09

nalon IJUl halon 1301

halon 1301

A 1 1 TflA

(J. 1 1 /yy -0.00118 -0.00236

A \ AH AO

U. 14/48

0.02596

0.03068

A y AC\Q

A

U.14yo4

0.05545

0.06725

A T 0 1 AA u.zoiyy

0.10265 0.09557

A 1 CAOO A 1 ^OT^

0.14394

NIT40/10

NIT40/11

NIT40/12

FC-218 FC-218

FC-218

0.01298 0.02950

-0.01416

0.07551 0.05781 0.04484

0.11563 0.02950

0.03422

0.12743

0.08023 0.05781

NIT40/28

NIT40/29

NIT40/30

CF3I

CF3I

CF3I

-0.03068

-0.01180

-0.03658

-0.00708 0.06253 0.02950

0.26547

0.12507 0.09439

Sample

Number

Test Environ. 20 °C

AM/Area

30 Days

g/m^

AM/Area

60 Days

g/m^

AM/Area

120 Days

g/m^

AM/Area

180 Days g/m^

AM/Area

365 Days g/m^

6061/01 6061/02

6061/03

HFC- 125 HFC- 125 HFC- 125

0.30636

0.01409

0.00352

0.42609

0.08686 0.09860

0.80992 0.40614

0.48713

0.96487 0.49065

0.55638

1.19376 0.57986

0.68315

6061/04

6061/05

6061/06

HFC-227ea

HFC-227ea

HFC-227ea

0.05165 0.00587

0.00235

0.36857

0.28993 0.27467

0.50474

0.45074

0.38501

0.55990 0.74419

0.47187

6061/07

6061/08

6061/09

halon 1301 halon 1301 halon 1301

0.11738 0.02348

0.03521

2.53072

0.33219

-0.00352

9.66862

0.02348

1.33109

-8.98664 0.08569 3.35238-9.22493 -0.44957

-2.11285

6061/10

6061/11 6061/12

FC-218

FC-218 FC-218

0.03639

0.05986

0.04343

0.34392

0.56225 0.45426

0.46952

0.82284

0.71485

0.57869 0.91791 0.80288

6061/28

6061/29

6061/30

CF3I

CF3I

CF3I

0.00352

0.03991 0.03287

4.36068

53.50903

78.45355

-4.69052

-6.39019 -17.90990
198 5. COMPATIBILITY WITH METALS

Sample Test Environ. AM/Area AM/Area AM/Area AM/Area AM/Area

Number 20 °C 30 Days 60 Days 120 Days 180 Days 365 Days

^ g/m2 g/m2

304/01 HFC- 125 0.20255 0.21660 0.32548 0.41212 0.07844 0.51398 0.09249

304/02 HFC- 125 -0.01756 -0.01756 0.06088 0.13815 0.08781

304/03 HFC- 125 0.00937 0.01288 0.06791

304/04 HFC-227ea 0.03512 0.07493 0.08781 0.09015 0.08898

304/05 HFC-227ea -0.00351 0.02459 0.02225 0.07259

304/06 HFC-227ea -0.00820 0.04449 0.04566

304/07 halon 1301 0.05503 0.23182 0.13113 0.29153 0.22245 0.19669 0.15572

304/08 halon 1301 0.02107 0.03161 0.07376 0.29504 0.24587

304/09 halon 1301 -0.01873 0.03512 0.10771

304/10 FC-218 0.01054 0.12059 0.04566 0.14635 0.09015

304/11 FC-218 -0.00585 0.05034 0.07142 0.09015

304/12 FC-218 -0.00234 0.03512 0.04566

304/13 CF3I -0.01171 0.31143 0.54559

304/14 CF3I -0.02576 0.27162 0.33485

304/15 CF3I -0.01054 0.18850 0.44959

Sample Test Environ. AM/Area AM/Area AM/Area AM/Area AM/Area

Number 20 °C 30 Days 60 Days 120 Days 180 Days 365 Days

g/m^ g/m^ g/m^ g/m^ g/m2

Ti/01 HFC- 125 0.02665 0.03128 0.07300 0.13788 0.13904 0.22594 0.14368

Ti/02 HFC- 125 0.00579 0.02549 0.09733 0.16106 0.16569

Ti/03 HFC- 125 -0.01275 0.03592 0.13325

Ti/04 HFC-227ea 0.01275 0.05098 0.10892 0.12050 0.11587

Ti/05 HFC-227ea -0.01506 0.08574 0.08458 0.11471

Ti/06 HFC-227ea 0.00463 0.04055 0.09849

Ti/07 halon 1301 0.00927 0.10080 0.10196 0.14020 0.12861 0.18307 0.17496

Ti/08 halon 1301 -0.00811 0.06604 0.06720 0.18539 0.20856

Ti/09 halon 1301 0.03013 0.11239 0.10660

Ti/10 FC-218 0.01854 0.07995 0.09269 0.18886 0.13556

Ti/11 FC-218 0.00000 0.06604 0.09269 0.12398

Ti/12 FC-218 0.02665 0.07995 0.09965

Ti/28 CF3I 0.01738 0.29083 0.35571

Ti/29 CF3I -0.03476 0.06720 0.23405

Ti/30 CF3I -0.01854 0.10080 0.24680
5. COMPATIBILITY WITH METALS 199

Sample Test Environ. AM/Area AM/Area

Number 20 °C 60 Days 150 Days

g/m^

4130/01 HFC- 125 0.30558 0.49553

4130/02 HFC- 125 0.13686 0.96156

4130/03 HFC- 125 1 .44058 2.59917

4 1 30/04 HFC-227ea 0.09675

4 1 30/05 HFC-227ea 0.32209 —

A I "3 A/Ail 41 JU/Uo HrL-zi /ea -1.05241

4130/07 halon 1301 2.14140 -1.28838 4130/08 halon 1301 0.28906 6.95394

halnn 1301 0.41648

4130/10 FC-218 0.48373

4130/11 FC-218 0.57340

4130/12 FC-218 0.25838

4130/30 CF3I 69.90163 S7.2Sm

4130/31 CF3I 44.32047 88.18198

4130/32 CF3I 39.79817 80.18507

Sample

Number

Test Environ. 150 °C AM/Area

90 Days

g/m^

321/01 321/02

321/03

halon 1301

halon 1301

halon 1301

0.06701 0.05172

0.06877

321/07

321/08 321/09

HFC- 125 HFC- 125 HFC- 1250.04702 0.00000

0.03409

Sample

Number

Test Environ. 150 °C AM/Area

90 Days

g/m^

CDAl 10/01 CDAl 10/02 CDAl 10/03halon 1301

halon 1301

halon 1301

0.20589

0.09883 0.07412 CDA11 0/07 CDA 11 0/08 CDAl 10/09

HFC- 125 HFC- 125 HFC- 125

0.11059

-0.00118

-0.00471
200 5. COMPATIBILITY WITH METALS

Sample

Number

Test Environ. 150 °C AM/Area

90 Days

g/m2

CDA 172/01 CDA 172/02 CDA 172/03 CDA172/07 CDA172/08 CDA172/09

halon 1301 halon 1301

halon 1301

-0.05058 0.00118

-0.00588

HFC- 125 HFC- 125 HFC- 125

0.05646

0.00706

-0.04705

Sample

Number

Test Environ. 150 °CAM//Vrea 90 Days

g/m2

355/01 355/02

355/03

halon 1301 halon 1301 halon 1301-0.02295 0.06196

0.02065

355/07 355/08

355/09

HFC- 125 HFC- 125 HFC- 125

-0.07803 0.02065

0.05508

Sample

Number

Test Environ. al50 °C AM/Area

90 Days

g/m^

1625/01 1625/02

1625/03

halon 1301 halon 1301

halon 1301

0.00708

0.00708 0.03009

1625/07

1625/08 1625/09

HFC- 125 HFC- 125 HFC- 125

-0.00354 0.01593

-0.00177
201

6. INTERACTIONS OF AGENTS WITH ELASTOMERS AND

LUBRICANTS

Gregory B. McKenna, Ferenc Horkay, Peter H. Verdier and William K. Waldron Jr. Polymers Division, Materials Science and Engineering Laboratory

Contents

Page

6. INTERACTIONS OF AGENTS WITH ELASTOMERS AND LUBRICANTS 201

6.1 Introduction 202

6.2 Swelling Measurements 203

6.2.1 Theoretical Background 203

6.2.2 Experimental Materials and Methods 204

6.2.3 Results and Discussion of the Swelling Measurements 205

6.3 Mechanical Property Measurements 208

6.3.1 Compression Set 208

6.3.2 Tensile Testing 217

6.3.3 Extrapolating Results to Longer Times 219

6.3.3.1 Compression Set 219

6.3.3.2 Tensile Testing 225

6.4 Testing of Elastomers and Lubricants in CF3I 234

6.4.1 Anomalous Testing Results with CF3I 234

6.4.2 Swelling Measurements in CF3I 235

6.4.3 Durability Test Results in CF3I 235

6.5 Summary 235

6.5.1 Swelling measurements 239

6.5.2 Mechanical Property Measurements 240

6.5.3 Compression Set Measurements 240

6.5.4 Tensile Testing 241

6.6 References 246
202 6. INTERACTIONS OF AGENTS WITH ELASTOMERS

6.1 Introduction

Elastomers and greases used in fire suppression systems serve to prevent container leakage and subsequent suppression system malfunction due to insufficient fire suppressant agent or container pressure. Excessive swelling or mechanical deterioration, particularly of the elastomeric seals, can lead to such

leakage. The goal of the work undertaken in this part of the project was to obtain data on the

compatibility of elastomeric seals and greases with candidate fire suppression agents. In the process a methodology was developed to obtain data that characterizes the swelling of elastomers and resistance

of elastomers to mechanical degradation in the presence of the candidate elastomers.

In the first phase of the project short term exposure experiments were conducted to characterize

the compatibilities of 11 agents with 6 commonly used elastomers and 3 lubricants (McKenna et ai,

1994). Two types of measurement techniques were developed and used:

(i) swelling measurements due to sorption of the agent into the elastomers (or lubricants) at four

different temperatures from 35 °C to 150 °C, and

(ii) durability measurements, which produced data on residual mechanical (rheological) properties

of the elastomers (and lubricants) after exposure to the agents at 150 °C and 5.86 MPa.

Based on the results of the swelling tests a rating system was proposed to characterize the

compatibility of elastomers (and lubricants) with the different agents. In swelling measurements good

compatibility (i.e., an elastomer or lubricant is acceptable for use with the fire suppressant) was

defined as the value of the Flory-Huggins polymer-solvent interaction parameter (%) determined at 35 °C being larger than 1.2. Bad compatibility (% < 0.9) corresponds to excessive swelling. For

values of 0.9 < % < 1.2, the agent was considered to have fair compatibility with the elastomer or

lubricant and represents a marginally acceptable system.

In the measurements of the residual mechanical properties, compatibility ratings were to have

been based on the results of compression set and tensile test data of the elastomers. The important

conclusion to be drawn from the durability testing was that the 150 °C temperature was too high for

reasonable evaluation of changes in the mechanical properties of candidate elastomers and greases and

that the degradation resistance after long term exposure of the elastomeric materials could not be

estimated from the results. Therefore, in the cases of the four agents selected for further testing (HFC227ea, HFC-125, FC-218, and CF3I (not considered in the first phase of the project)) measurements were performed at 50, 75 and 100 °C. The temperature 75 °C was chosen to conduct the long-term

mechanical property measurements. Information from the sponsors indicated that all-metal storage

containers are used for exposures above 70 °C. Further testing of the greases was not considered

necessary. This report summarizes the work performed in FY 1994-95 and is organized as follows:

Section 6.2 describes the results of the swelling measurements for seven sets of crosslinked and

uncrosslinked elastomers in the 3 selected fire suppressant agents. (Note that in the first phase of the

project only six elastomers were investigated. The seventh elastomer, ethylene-propylene-diene

(EPDM) terpolymer was suggested by NIST for completeness of the study because it is a widely-used,

environmentally-inert elastomer.) The results are compared with those obtained for halon 1301 which

serves as a baseline to compare the properties of the fluorocarbon alternatives. In Section 6.3 the

results of the durability measurements are shown. A Kohlrausch-Williams-Watt (KWW) type

stretched exponential function (Williams and Watt, 1970) is used to describe the compression set data. This function is used to predict the changes in the mechanical properties after long exposure times. In Section 6.4 the results obtained for iodotrifluoromethane (CF3I) are presented. This chemical was not

investigated in the first phase of the project. Later it was recommended for inclusion by the sponsor

of the program because it was found to be as effective as halon 1301 in suppressing non-premixed
6. INTERACTIONS OF AGENTS WITH ELASTOMERS. 203

flames. Since testing in this agent was accompanied by several specific problems, the results of these

investigations are discussed separately. In Section 6.5 the results of the thermodynamic (swelling)

investigations in combination with the durability measurements are summarized.

6.2 Swelling Measurements

6.2.1 Theoretical Background. A polymer exposed to a thermodynamically compatible diluent

absorbs solvent molecules. The driving force of the mixing process is mainly entropic. The FloryHuggins lattice theory of polymer solutions (Flory, 1942; Huggins, 1943, Flory 1953) describes the

free energy of mixing and allows the characterization of the affmity of the polymer (elastomer or

lubricant) to the solvent (fire suppressant agent) by a single polymer-solvent interaction parameter %. In a swelling measurement the concentration of the polymer at equilibrium, is measured as a

function of solvent activity (Boyer, 1945; Gee et ai, 1965; Brotzman and Eichinger 1982 and 1983;

Horkay and Znnyi, 1982; McKenna et al., 1989; Horkay et ai, 1989; McKenna et al, 1990; McKenna

and Crissmann 1993). According to the Flory-Huggins theory:

Inflj = hi(p/p'0 = ln(l - Vj) + (1-1/P)V2 + x^l

where p is the equilibrium vapor pressure of the diluent, p° is its saturation vapor pressure, V2 is the

volume fraction of the polymer, P is the degree of polymerization and % is the Flory-Huggins polymersolvent interaction parameter. In general % depends on the polymer volume fraction (Flory, 1970), i.e., X=Xo Xi^2 ••' where and Xi are constants. From Equation (1) it follows that the measurement of the equilibrium polymer volume fraction as

a function of the vapor pressure of the diluent allows the determination of the x parameter, i.e., the

estimation of the solubility of elastomers in the fire suppressant agents. Small x values (% < 0.5)

correspond to solubility across the whole concentration range. In this case uncrosslinked polymers are completely dissolved and crosslinked polymers exhibit large swelling which can lead to seal failure in elastomeric o-ring based joints. High x values are characteristic of polymer/solvent systems with

limited miscibility.

Although the Flory-Huggins theory is formulated in terms of volume fractions (vj and V2), the

theory can also be expressed in terms of weight fractions (wj and w^). This substitution does not change the functional form of Equation (1), i.e.,

]n(p/p°) = hi(l - Wj) + W2 + Xo^2 ^ Xi^2^

where x = Xq + Xi ^2 has been substituted, and the term (1 - 1 / P) has become unity since P is large

for polymeric materials and infinite in crosslinked networks. In this case the x values differ from

those calculated from the equilibrium volume fractions (V| and V2). Throughout the present report

weight fractions are used instead of volume fractions because they are directly measured quantities.
204 6. INTERACTIONS OF AGENTS WITH ELASTOMERS

Table 1 . Elastomers used in swelling experiments

Elastomer Vendor Designation

Silicone Colonial Rubber Si

55% Butadiene-45% Acrylonitnle Goodyear N206

Fluorosilicone Colonial Rubber FSi

Viton E-60 Fluorocarbon Du Pont FKM

Neoprene Colonial Rubber CR

85% Butadiene- 15% Acrylonitnle Goodyear N926

Ethylene-Propylene-Diene Exxon EPDM

Table 2. Lubricants used in swelling experiments

Lubricant Vendor Designation

Krytox 240AC Fluorinated Grease Du Pont 240AC

Braycote 600 Perfluoropolyether Grease, Low Volatility Castrol 600

Braycote 807 Aircraft Grease MIL-G-27617, Type IV Castrol 807

6.2.2 Experimental Materials and Methods. The elastomers and lubricants used in this study are shown in Tables 1 and 2, respectively.^ Both crosslinked and uncrosslinked elastomers were

investigated. The 85 % butadiene- 15 % acrylonitrile crosslinked copolymer was prepared according to a method described earlier (McKenna et al., 1994). The silicone, fluorosilicone, neoprene, and

ethylene-propylene-diene polymers were supplied crosslinked and uncrosslinked by the vendor. The

fluorocarbon polymers were cured at NIST for one hour at 150 °C and post-cured at 240 °C for 24

hours.

The swelling measurements were performed in an apparatus (Figure 1) described earlier (McKenna et al., 1994). The pressure vessels were made of type 304 stainless steel and designed for a maximum working pressure of 5.86 MPa with two view ports 180° apart for viewing and backlighting

purposes.

Certain trade names and company products are mentioned in the text or identified in an illustration in order to specify adequately the experimental procedure and equipment used. In no case does such

identification imply recommendation or endorsement by the National Institute of Standards and

Technology, nor does it imply that the products are necessarily the best available for the purpose.
6. INTERACTIONS OF AGENTS WITH ELASTOMERS.

Table 3. Saturation data for agents

205

Trade p° @ p° @ p° @ 7'c name

Formula Chemical name 35°C - 70 °C 105°C (MPa) (MPa) (MPa) (°C) (MPa)

HFC-227ea C3HF7 heptafluoropropane 0.615 1.49 101.7 2.91 HFC- 125 CHF2CF3 pentafluoroethane 1.78 66.3 3.62

FC-218 C3F8 octafluoropropane 1.15 2.41 71.9 2.68

halon 1301 CFjBr bromotrifluoromethane 2.04 67 3.96

p- critical pressure

p°: saturation vapor pressure

Samples weighing 25-50 mg were placed on quartz pans suspended from quartz springs (see

Figure 1). The sensitivity of the springs was either 1 mg/mm or 1.25 mg/mm. The vessels were

placed in a cylindrical Pyrex jar and completely immersed in silicone bath oil. Two rectangular holes were made so that the view ports were visible. The bath was heated using an immersion heater and

stirred to maintain temperature uniformity. A pressure charging and recovery system was used for

admission of the agent into the vessel and for recovery of it after each measurement.

The mass uptake by elastomers and lubricants was calculated from the displacement of the quartz

spring measured by a cathetometer. Measurements were taken at 35, 70, 105 and 150 °C as a function

of the vapor pressure. For each test temperature the agent vapor pressure was varied between the

saturation pressure at the test temperature and that at the next (lower) temperature. In this way

undesired saturation during cooling, which could ruin the samples, was avoided. Above the critical temperature 7^ of the agent the upper limit of the pressure range was the maximum working pressure

of the apparatus (5.86 MPa). Saturation pressures for each temperature were determined for agents

HFC-125, FC-218, and halon 1301 using thermodynamic properties software (Gallagher et al., 1991),

and for agent HFC-227ea from vendor supplied data (Robin, 1992) and are shown in Table 3. After measurements were performed for all temperatures considered, the vessel was evacuated and

the spring displacements were checked for reversibility. Where appropriate, corrections for buoyancy

effects were made (McKenna et al, 1994).

6.2.3 Results and Discussion of the Swelling Measurements. In Figures 2-5, typical plots of

solvent weight fraction W| vs. vapor pressure p are shown for different elastomer/agent systems. The

fluid uptake of the elastomers considerably increases with the vapor pressure and decreases with

increasing temperature (Figure 5). In Figures 6-8, the experimental data obtained at 35 °C are shown

according to the Flory-Huggins representation. The curves are the least squares fits of Equation (2) to

the data points. In Tables 4 and 5 are listed the values of the Flory-Huggins interaction parameters (x = Xo Xi) determined for the elastomers in the three selected agents and in halon 1301. (Similar

data for the three agents and the six elastomers were given in the 1994 report (McKenna et al. 1994).

In order to check the reproducibility of the swelling tests we repeated these measurements. The

previous values of x are given in parentheses in Tables 4 and 5. The results indicate that the

repeatability of the % data is, generally, better than 5 %.) In Table 6 the % values are shown for the

three lubricants, x nrieasures the solubilities (compatibilities) of the subcritical agents in various

polymers. It is apparent from the data that the values of the interaction parameters for the crosslinked
206 6. INTERACTIONS OF AGENTS WITH ELASTOMERS

Figure 1. Schematic drawing of the experimental apparatus for isopiestic swelling measurements.
6. INTERACTIONS OF AGENTS WITH ELASTOMERS. . . 207

Table 4. Flory-Huggins interaction parameters % for subcritical temperatures studied for different

agent/elastomer systems

u'^CR x'^CR u N926 X N926 II FPDM

£.4. 1 Cd I. J J ^

1 .L.J f)w. 78/ o U.O*T 1 .J I

1 "iA

(1.25'') (1.33^) (0.80) (0.84)

227ea (70) 1.29 1.44 0.85 0.88 1.35 1.37

(1.31) (1.38) (0.85) (0.87)

125 (35) 1.99 2.03 1.22 1.26 1.98 2.02

(4.10'') (1.99) (1.23) (1.28)

218 (35) 1.64 1.70 1.62 1.71 1.53 1.58

(1.63) (1.68) (1.56") (1.71)

218 (70) 1.74 1.77 1.67 1.74 1.62 1.70

(1.75) (1.77) (1.69) (1.74)

halon 1301 (35) 1.98 2.08 1.25 1.12 1.06 1.12

^12.5 <CV < 20% and 0.64 < x < 1.5 (35 °C) ^CV > 20% (35 °C)

"^uncrosslinked polymer

''crosslinked polymer

polymers slightly exceed those of the corresponding uncrosslinked systems, i.e., the crosslinked poly- mer exhibits lower affinity to the solvent than its uncrosslinked counterpart. (Strictly speaking

Equation (2) can only be applied for uncrosslinked polymers. The appropriate expression for a

crosslinked polymer contains an additional term arising from the elasticity of the network (James and

Guth, 1943; Flory and Rehner, 1943; Flory and Erman, 1982). For lightly crosslinked networks at high

polymer concentration (w2 > 0.7), however, this contribution is small in comparison to the mixing

term.)

The rating system for characterizing the compatibility of the agents with the sealing materials was

defined on the basis of the % parameters determined at 35 °C. This temperature was chosen for

several reasons. First, as shown in Figure 5, swelling is largest at 35 °C and gradually decreases with

increasing temperature. Second, it is near the initial storage temperature of 25 °C. A good, bad, or

fair rating is given for each agent/crosslinked elastomer system in Table 7. The X values shown in Tables 4-6 were calculated using a commercial software package.

The curve fitter (SigmaPlot, Jandel, 1992) uses the Marquardt-Levenberg algorithm (Press et al., 1986)

to find the parameters which minimize the variance s^ (In p I p°), i.e., the squared differences between

the observed and predicted values of the dependent variable. The standard error or experimental

standard deviation of the % parameter, s ( x ), and the coefficient of variation, CV = 100s(x)/x, were
208 6. INTERACTIONS OF AGENTS WITH ELASTOMERS

Table 5. Flory-Huggins interaction parameters x for subcritical temperatures studied for different

agent/elastomer systems

Agent ( 1 C) U bl U IN 206 X N206 u rSi X FSi u FKM X FKM

227ea (35) 0.89 0.91 1.33 1.43 0.98 1.04 1.00 1.07

CO 8Q^ n 34*"^ n 4o'''> (0 971 CI nil

227ea (70) 0.92 0.97 1.50 1.50 1.08 1.15 1.10 1.12

(0.94) (0.96) (1.48) (1.50) (1.10) (1.13) (1.09) (1.13)

125 (35) 1.26 1.32 1.69 1.75 1.34 1.36 1.17 1.24

(1.29) (1.30) (1.69") (1.72) (1.33) (1.35) (1.18) (1.21)

218 (35) 1.55 1.63 1.75 1.85 1.53 1.64 1.59 1.71 (1.55) (1.62) (1.77") (1.84) (1.56") (1.64) (1.65) (1.65)

218 (70) 1.60 1.71 1.90 1.99 1.66 1.75 1.68 1.63

(1.64) (1.69) (2.04) (1.99) (1.67) (1.74) (1.66) (1.67)

1301 (35) 0.92 0.97 2.15 2.16 1.04 1.11 1.69 1.71

'12.5 < CV < 20% and 0.64 < x < 15 (35 °C) "CV > 20% (35 °C)

'^uncrosslinked polymer

''crosslinked polymer

also calculated for each % value. Thus, the standard uncertainty of the parameter / is m(x) = s(x)

and the percentage of relative standard uncertainty is shown in Tables 4 and 5 with superscripts.

Estimated values of % for which CV > 20 % are also shown with superscripts in Tables 4-6.

6.3 Mechanical Property Measurements

In order to characterize the long-term compatibility of the elastomers with the fire suppressant agents,

mechanical property measurements (compression set and tensile testing) were performed after exposure

times of 1,2, 4, 8, 16, 32, 48 and 74 weeks. Exposure to the agents was carried out at 75 °C and

5.86 MPa, in pressure vessels similar to those used for the swelling measurements. Mechanical

property measurements were also made after short term (1,2 and 4 weeks) exposures at 50, 100, and

150 °C.

6.3.1 Compression Set. Measurements were carried out using standard test methods (ASTM, 1989

and 1990b). The o-rings used in this study were Parker No. 2-214 with a nominal cross section

diameter (o-ring thickness) of 3.2 mm. The Parker compounds used are listed in Table 8. Specimens

were prepared by cutting 52 mm sections from sample o-rings. The original thickness of each

specimen was measured with a hand micrometer. Two specimens of each compound were placed in a
6. INfTERACTIONS OF AGENTS WITH ELASTOMERS. 209

Table 6. Flory-Huggins interaction parameters % for the subcritical temperatures studied for each

agent/lubricant system considered

Agent (T °C) Krytox 240AC Braycote 600 Braycote 807

227ea (35) 1.14^^ 0.93 0.93*'

227ea (70)

' 1.23 0.97 1.02

125 (35) 1.18 1.37 1.35

218 (35) 1.83'' 1.45 1.76

218(70) 1.86 1.47 1.79

Halon (35) 1.36 1.50 1.57

'12.5 <CV< 20% and 0.64 < % < 1.5 (35 °C)

^CV > 20% (35 °C)

compression set fixture. Each fixture consisted of two 18 cm diameter and 12.7 mm thick plates and

seven spacers 2.65 ± 0.01 mm thick, 75 % of the measured mean value of the o-ring thicknesses.

After removing from the fixture the specimens to be tested were kept at room temperature for a minimum of 30 minutes before measuring. Measurements were taken at four marked locations

equally distributed around the circumference in the axial direction, both on the unexposed samples and

after each exposure time. The average value was used for subsequent calculations.

I The compression set C is defined as • .

rv YV • C = '- X 100, (3)

where is the original average thickness of the o-ring, Wj is the o-ring thickness after exposure and

is the thickness of the fixture spacers.

In Table 9 are displayed the compression set results after each exposure time at 75 °C for all the

elastomers in each fluorocarbon agent. As specified in Option 1 of the standards for compression set (ASTM, 1989 and 1990b), the value of compression set for each elastomer/agent system reported in Table 9 is the average of the compression set calculated from measurements of two samples. The

sample standard deviation of the mean (sdm) of each pair of measurements was also calculated. The

values of sdm obtained were 2 or less except for the fluorosilicone samples, for which the maximum

value obtained was 4.4.
210 6. INTERACTIONS OF AGENTS WITH ELASTOMERS

^1

0.3

0.2 -

o xSi

• xN206

V xFSi

xFKM

xCR

xN926

A xEPDM

0.1

0.0

0.1

8

¥

i

0.2

Ob

§8

0.3

O

o

0.4

o

V

o

V

0.5 0.6

p/MPa

Figure 2. Weight fraction Wj of HFC-227ea as a function of pressure for various crosslinked

elastomers at 35 °C.
6. INTERACTIONS OF AGENTS WITH ELASTOMERS. 211

^1

0.2

0.1

O xSi ^

• xN206 i

V xFSi

T xFKM J

^

xCR o

xN926 *8 ^

ov

A xEPDM V

®

[}3

6

1.0 1.2 1.4 1.6

p/MPa

Figure 3. Weight fraction Wj of HFC- 125 as a function of pressure for various crosslinked

elastomers at 35 °C. .

:
212 6. INTERACTIONS OF AGENTS WITH ELASTOMERS

W

O xSi

• xN206

V xFSi

xFKM

xCR

xN926

A xEPDM

i

0.4 0.6 0.8 1.0 1.2

p/MPa

Figure 4. Weight fraction Wj of FC-218 as a function of pressure for various crosslinked

elastomers at 35 °C.
6. INTERACTIONS OF AGENTS WITH ELASTOMERS. 213

^1

0.3

0.2

0.1

0.0

35 C

1 1 1 open symbols: uSi

filled symbols: xSi

V

¥ 70 c

-

1

f

¥ . -

i

i

1

105 °C

^

9

9 •

1 1 1 0.0 0.5 1.0 1.5 2.0 2.5

p/MPa

Figure 5. Weight fraction Wj of HFC-227ea as a function of pressure in uncrosslinked and

crosslinked silicone at various temperatures.
214 6. INTERACTIONS OF AGENTS WITH ELASTOMERS .

In (p/p )

0,6 0.7 0.8 0.9 1.0

*2

Figure 6. Solvent activity in various crosslinked elastomers in HFC-227ea as a function of

polymer weight fraction W2 at 35 °C. The curves show the least squares fits according to Equation (2).
6. INTERACTIONS OF AGENTS WITH ELASTOMERS. 215

Figure 7. Solvent activity in various crosslinked elastomers in HFC- 125 as a function of

polymer weight fraction W2 at 35 °C. The curves show the least squares fits according to Equation (2).
6 6. INTERACTIONS OF AGENTS WITH ELASTOMERS .

Figure 8. Solvent activity in various crosslinked elastomers in FC-218 as a function of polymer

weight fraction at 35 °C. The curves show the least squares fits according to Equation (2).
6. INTERACTIONS OF AGENTS WITH ELASTOMERS. 217

Table 7. Compatibility of crosslinked elastomers based on swelling measurements in various

fluorocarbon agents and halon 1301 at 35 °C.

Agent Si N206 FSi FKM CR N926 EPDM

HFC-227ea fair good fair fair good bad good

HFC- 125 good good good fair good good good

FC-218 good good good good good good good

Halon 1301 fair good fair good good good fair

Table 8. Elastomers used in durability experiments

Elastomer Vendor Designation

Silicone Parker S604-70^

Nitrile (standard industrial) Parker N674-70

Fluorosilicone Parker LI 120-70

Fluorocarbon Parker VI 164-75

Neoprene Parker CI 185-70

Nitrile (low temperature industrial) Parker N103-70

Ethylene-propylene-diene Parker E0540-80

^The number following the dash in the designation (compound number) represents the Shore

hardness of the elastomer.

6.3.2 Tensile Testing. Tensile tests were conducted using standard test methods (ASTM, 1990a and

1990b). The exposure was made in the same pressure vessels used for the compression set measurements. Ultimate elongation measurements were taken on sample o-rings after each exposure time.

Elongation to failure and tensile strength were determined by stretching each sample at constant speed

until it broke. The thickness of each specimen, W, was measured at four points equally distributed

around the circumference in the radial and axial directions using a dial indicator with a contact force

of 0.2 N. The average of all eight measurements was used for calculation. The internal diameter of

the o-ring was measured using a stepped cone with diametric intervals of 0.50 mm.

An Instron type TT-B instrument was used for the testing. Grips for testing o-rings were

designed and built from low carbon steel. The grips consisted of ball-bearing spools 8.9 mm in diameter. Stresses were minimized by lubricating the contact surfaces of the spools with Super-Lube
218 6. INTERACTIONS OF AGENTS WITH ELASTOMERS

Table 9. Compression set after 1, 2, 4, 8, 15, 32, 48 and 74 week exposures at 5.86 MPa and 75 °C'

in the three fluorocarbon agents

Elastomer

(agent) 1 2 4

weeks

8 15 32 48 74

oOU4- /u (,// / ) 0 0 1 /

1

0

7^2j 7C

O0U4- l\) \ ILJ) 7 0

J 1 "X^ JJ JO

O0U4-/U (Zlo) 0 1

A

10 1ios 17J 1 Tfi JO 4U 41

INC)/4-/U (ZZ/) 1 0 1Q 2o 20 J 1 IKJj Jo J DU

INO/4-/U ylZj) 1

J

ZV 48 SO OD oU

INO/4-/U (/lo; 1 e

1

J

1 S15 Tfl jU jy 00 70

LI 120-70(227) 4 IZ Af\ 4U A s 4j SADo A1ol 04

I 1 1 in '7n/i'^c\ L1120-70(125) t <ID 23 /1

1

41 A702 00 71/

1

7>l /4 71

L1120-70(218) 24 J / A 1 4 / 77/2 74 / / 77 7Qly

V 1 104-/0(22/) 1 1 -11 1

1

-11 1

1

- 1

1

-4 11 3J 0

V

1

1d4- /D(lZD) -0Q -0 n

U 0

1 A14 1 fi lo 2U

V

1

1d4- / j(Z1 0) 2

A4cJ inlU 14. I'f 14. 171 /

Cll85-70(227) 12 15 22 35 52 60 66 74 C 11 85-70(125) 9 16 34 46 57 70 77 88

C1185-70(218) 11 20 30 35 50 65 76 85 N 103-70 (227) 20 24 30 37 40 47 53 65

N103-70 (125) 21 26 41 49 60 73 80 89

N103-70 (218) 28 37 49 50 63 75 84 92

E0540-80(227) 12 14 17 15 21 24 26 43

£0540-80(125) 10 11 15 16 21 27 30 49

£0540-80(218) 9 11 17 15 24 25 30 46
6. INTERACTIONS OF AGENTS WITH ELASTOMERS. 219

o-ring lubricant. The crosshead speed was set to 50 cm/min, and a load cell with full scale load

ranges of 2, 5, 10, 20, and 50 kg was used.

The ultimate elongation (U.E.) is expressed as a percentage of the original inside circumference as follows:

U.E. =

+ G-C ^ C

where D is the distance between centers of the spool grips at the time of rupture, G is the circumfer- ence of a spool, and C is the inside circumference of the specimen.

The tensile strength T.S. is the tensile stress at rupture and is calculated as F/A, where F is the

breaking force and A is twice the cross-sectional area, calculated from the thickness W as The median of three measurements was used. Tables 10 and 11 show the percent change in ultimate elongation and tensile strength for all the elastomers in each agent after different exposure

times at 5.86 MPa and 75 °C.

6.3.3 Extrapolating Results to Longer Times. As discussed in Section 6.1, a major goal of this work is prediction of elastomer-agent compatibility at long times based on shorter-time measurements.

We have employed two approaches here. In the first, we measure changes in mechanical properties as

a function of time at fixed temperature, and fit the compression set C to a stretched exponential

function C^{ 1 - exp[-(t/T)'^] }, where C^, x, and (3 are adjustable parameters. If the fit is sufficiently

good, the fitted functions may be used to predict results at longer times. In the second (activated

process model), we seek to predict long-time behavior at lower temperatures from short-time behavior

at higher temperatures by fitting the time required to achieve a given change in mechanical properties

to the form t'-exp(-AH/RT), where t' and AH/R are adjustable parameters and T is absolute

temperature.

6.3.3.1 Compression Set. Figures 9-11 show log-log plots of the compression set vs. exposure

time data in Table 9 for elastomers in HFC-227ea, HFC-125, and FC-218 at 75 °C. (The data for

elastomer VI 164-75 are not shown for agents HFC-227ea and HFC-125 because the compression set values were small or negative due to swelling.) The lines show the result of fitting the measured

values of compression set for exposure times from 1 to 48 weeks to a KWW type stretched exponential function

C=CJl-exp[-(?/T)P]} (5)

where C is the compression set, t is the exposure time, and T and |3 are the fitted parameters. The

values of C^,t and (3 so obtained are given in Table 12. The data obtained at exposure times of 74

weeks have not been used in the fitting. These 74-week values may be compared with the results predicted from measurements at shorter times to assess the reliability of the fit. It can be seen from

Table 9 and Fig. 9-11 that the compression set for N674-70 in agents HFC-227ea and HFC-125, and

for E0540-80 in all three agents, is markedly higher after 74 weeks exposure than would be predicted

from extrapolations based on the behavior at shorter times.

Table 13 shows the values of compression set obtained for elastomers in HFC-227ea, HFC-125,

and FC-218 after short-term exposures at 50, 100 and 150 °C. For the activated process model, times

tjQ and t2o required to obtain compression sets of 10 % and 20 %, respectively, were estimated for
220 6. INTERACTIONS OF AGENTS WITH ELASTOMERS

Table 10. Percent change in tensile strength after 1, 2, 4, 8, 15, 32, 48 and 74 week exposures at 5.86 MPa and 75 °C in the three fluorocarbon agents

Elastomer

(.agent; ] 2 4

weeks

g 15 32 4Kto 74

5604-70 (227) 14.4 10.9 21.5 0.3 -15.9 -6.7 -11.5 -4.6

S604-70 (125) 11.8 26.4 17.8 -8.9 -5.0 1.8 -10.3 -1.7

S604-70 (218) 12.7 14.4 12.7 -2.6 -5.8 -16.8 -0.9 -13.3

N674-70 (227) 4.6 10.7 -0.1 -65.5 -10.6 -33.4 -21.2 -42.4

N674-70 (125) 16.4 28.8 23.0 -68.6 -0.1 -25.8 -17.9 -41.8

N674-70 (218) 33.8 30.3 35.3 -66.5 -1.6 1.5 -1.6 -13.7

L 11 20-70(227) 17.7 -51.7 13.7 11.0 16.9 13.0 -0.1 -4.9

L 11 20-70(1 25) 28.5 7.4 40.9 -9.7 -0.1 -3.5 -0.7 -12.3

L1120-70{218) 14.1 30.2 20.8 -11.0 -4.9 1.2 -19.4 -7.8 V 1164-75(227) -6.0 -7.3 1.2 -7.7 -20.4 -44.4 -25.0 -42.6

VI 164-75(125) -38.7 -50.8 -71.9 -52.5 -75.6 -86.4 -72.5 -85.0

VI 164-75(218) 12.3 28.6 10.0 -14.7 -28.6 -23.8 -15.6 -29.8

C1185-70(227) 13.8 18.9 24.2 -20.1 -23.0 -20.7 -13.4 -25.3

CI 185-70(125) 18.9 20.6 15.5 -4.8 -20.7 -34.0 -20.7 -39.7

C1185-70(218) -7.3 39.0 7.0 -12.5 -13.7 -17.0 -39.1 -44.7

N103-70 (227) 40.7 29.3 38.8 34.8 16.1 -1.2 -6.4 3.0

N 103-70 (125) 33.0 51.3 46.4 27.3 15.8 -7.7 -0.3 -28.0

N103-70 (218) 55.0 59.2 40.2 16.2 10.3 3.0 1.1 -8.2

£0540-80(227) 13.0 13.0 19.9 -3.2 -4.5 -5.9 -47.7 -17.1

£0540-80(125) 13.7 19.9 9.1 -9.1 -6.8 -11.6 0.0 -7.7

£0540-80(218) 31.0 31.4 31.4 -3.2 -3.6 -14.7 -6.8 -17.1
6. INTERACTIONS OF AGENTS WITH ELASTOMERS. 221

Table 11. Percent change in ultimate elongation after 1, 2, 4, 8, 15, 32, 48 and 74 week exposures at 5.86 MPa and 75 °C in the three fluorocarbon agents

Elastomer weeks

(agent) 12 15 4 8 32 48 74

oC5U^- /U (ZZ / ) t 9 J./. 1 s -/.J 1 0 1 fi -J.O 1 fi -J.O -

1 n Q 1U,8 J 0 J.Z OOW- IK] (IZj) 1 0 "X 0 7 c - /.J n ^ -U.J fi, < O.J 7 n "7 0 -J.O aoW- /U (Zlo) 0 0 -1 /./ - 1 1 .J - 7 n /.u "5 fi -J.O 1 fi 1 -18.J 1 -1.0 1 n fi -1U.8

INO /'- /U 1

)

-J.U - 14 fi 14.0 1 0 T -10.

J

9fi T -Z8. J 90 9 -zy.z -44 AA n.U

IN0/4-/U (1^0) O.O z. / -D.J fi 1 -o.l -18.4 1 fi 4 19 9 -jZ.Z 'ifi -jB.j <;7 fl -J /.8

INO IH- lU \Alo) 1 B 1.0 -o.fi T/

fi 7 -o. / - 1 9 7 IZ.J 9n 9 -ZU.Z -Z4.4 94 4 -jO. 1 -jU.j

S 1 O.l -01.J

4 A 7 S II

1 Q I .y 9n n zu.u O.J 1 fi J.O

111 on on/ 1 o<^ -zu.o 1 J.O fi 1 -8. i J.

8

fi 1 8.1 1 n n lU.U "3 1 J.l

Til on on/o 1 fi^ -iU.U 1 0 94 4 94 4 - ID.

J

1 Q ifi 8 -J8.0 -11 T. 1 J.l

\/11£>4 OC/OOON 0.1 1 fi 1 ^ 1 1 J. 1 9 1 fi Z 1 .0 I 1 9 I I .z -y.j 1 1 1.1 fi o -o.y

Vn64-75(125) -29.6 -43.6 -59.8 -29.1 -56.4 -67.0 -50.8 -67.0

VI 164-75(218) 16.8 24.6 15.1 18.4 7.8 18.4 18.4 7.8 0 185-70(227) 1.6 8.1 12.1 -17.0 -19.4 -11.7 -19.8 -24.7

C1185-7(X125) 10.9 10.9 1.6 -3.6 -17.0 -22.3 -31.5 -45.7

C1185-7(X218) -8.9 10.1 -4.9 -6.5 -6.5 -8.9 -42.7 -44.9

Nl 03-70 (227) 18.0 2.9 5.7 13.9 3.3 -13.9 -46.9 -27.3

N103-70 (125) 5.7 16.7 3.3 0.0 -13.9 -27.8 -20.4 -60.8 N103-70 (218) 13.9 2.9 -12.2 -5.7 -16.3 -30.6 -46.9 -58.0

£0540-80(227) 5.0 -1.5 6.4 -1.5 -5.0 1.5 -12.9 -14.4

£0540-80(125) 1.5 -3.0 -7.9 -7.9 -5.0 -2.0 -7.9 -11.4

£0540-80(218) 5.0 9.9 0.0 -5,0 -1.5 -5.0 -7.9 -11.4
222 6. INTERACTIONS OF AGENTS WITH ELASTOMERS

c

<DOi_

0)Q.

c5

(0co

(0

COs

Q.

E

oO

S604-70

N674-70

L1 120-70

C1 185-70

N 103-70

E0540-80

fit S604-70

fit N674-70

fit L1 120-70

fit C1 185-70

fit N 103-70

fit E0540-70

HFC-227ea

^

10 -

—\

r 10 100

Exposure time, weeks

Figure 9. Compression set for various elastomers vs. exposure time to agent HFC-227ea at 5.86 MPa and 75 °C.
6. INTERACTIONS OF AGENTS WITH ELASTOMERS.

• S604-70

N674-70

L11 20-70

T C11 85-70

N 103-70

E0540-80

fit S604-70

fit N674-70

fit L1 120-70

fit C1 185-70

fit N 103-70

fit E0540-70

HFC-125

10 100

Exposure time, weeks

Figure 10. Compression set for various elastomers vs. exposure time to agent HFC-125 at 5.86 MPa and 75 °C.
224 6. INTERACTIONS OF AGENTS WITH ELASTOMERS

• SS604-70

N674-70

L1 120-70

V1 164-75

C1 185-70

N 103-70

0 E0540-80

fit S604-70

fit N674-70

fit L1 120-70

fit V1 164-75

fit C1 185-70

fit N 103-70

fit E0540-80

FC-218

10 100

Exposure time, weel<s

Figure 11. Compression set for various elastomers vs. exposure time to agent FC-218 at 5.86 MPa and 75 °C.
6. INTERACTIONS OF AGENTS WITH ELASTOMERS. 225

Table 12. Parameters C^, x (weeks), and (3 obtained by fitting compression set C vs exposure time t (weeks) to the KWW stretched exponential form of Equation (5).

elastomer

5604-70 N674-70 LI 120-70 VI 164-75 C 11 85-70 N103-70 E540-80

34.824 37.683 59.348 68.107 84.104 90.513

T 20.109 4.216 9.088 11.206 52.968 7245.738

aP 0.507 0.505 0.999 0.796 0.324 0.222

35.548 67.433 72.261 15.828 77.22 94.216 100

T 5.67 10.394 4.51 12.011 9.581 14.397 953.701

P 1.243 0.719 1.016 2.643 0.81 0.519 0.342

218 c« 39.854 88.16 77.387 15.803 100 100 42.19

T 4.6 24.164 3.483 8.102 28.241 13.98 32.139

P

1.077 0.542 0.859 1.036 0.589 0.413 0.409

CF3I 25.866 80.446 58.51 79.036 89.573 40.121

T/ 11 . 1 06 9.654 5.712 6.04 5.85 8.594

P 0.867 1.359 0.715 1.056 1.061 0.927

each elastomer-agent pair from the compression sets measured at exposure times of 1, 2, and 4 weeks,

and also at 6 weeks for 150 °C exposures. The compression set vs exposure time data for each agent,

elastomer, and temperature were fitted to Equation (5) by unweighted least squares. Values of tjo and

t2o were calculated from the parameters so determined; they are given in Table 14. For a few of the

systems, the data did not allow meaningful estimates of tjg and/or t2o to be obtained; these are shown

as dashes in Table 14. In other cases, especially at 150 °C, the estimates of tjg and/or t2o are too short

to be used with the activated process model. Where the data permitted, values of tjQ and X^q were

fitted separately by least squares to the form In(tj^) = ln(t'j,) -i- (AH/R)j,/T, where n is 10 or 20, T is absolute temperature in Kelvins, and ln(t'^) and (AH/R)^, are the fitted parameters, shown in Table 15. From the values of ln(t'^) and (AH/R),^ , estimates of t„ at other temperatures may be calculated.

Table 15 shows estimated values of t^, in weeks at 70 °C obtained in this way.

6.3.3.2 Tensile Testing. The percent change in ultimate elongation vs. exposure time data in Table 10 are shown in Figures 12-14 for elastomers in HFC-227ea, HFC- 125, and FC-218 at 75 °C.

Since many of the values obtained were negative, a linear scale has been used for the change in ultimate elongation. The scatter in the data is much too large to allow meaningful fitting to stretched

exponential or other relaxation functions.

Tables 16 and 17 show the percent change in ultimate elongation and tensile strength vs. exposure

time data for all the elastomer-agent pairs for short-term exposures at 50, 100 and 150 °C. As with

the longer-term data at 75 °C, there is too much scatter in the data to allow meaningful fits to activation process models.
226

Table 13

6. INTERACTIONS OF AGENTS WITH ELASTOMERS

Compression set after short-term exposures at 5.86 MPa and 50, 100 and 150 °C.

Exposure temperature 50 °C 100 °C 150 °C

Exposure time, weeks 1 2 4 1 2 4 1 2 4 6

Agent Elastomer

HFC-227ea S604-70 4 7 9 17 23 30 25 34 48 57

N674-70 7 10 17 34 38 47 77 84 93 97

LI 120-70 -2 1 9 39 46 52 51 56 72 83

VI 164-75 - -13 -12 -4 -2 0 21 33 60 —

CI 185-70 9 12 15 37 42 50 66 75 81 84 N 103-70 16 19 26 47 52 57 79 87 91 94

E0540-80 8 10 13 19 25 27 — — — —

HFC- 125 S604-70 3 5 9 16 21 24 33 36 50 58

N674-70 3 6 12 29 36 56 70 84 91 96

L 11 20-70 9 14 22 53 57 66 65 65 80 90

VII 64-75 — -12 -12 4 6 15 35 46 72 90

CI 185-70 5 7 12 34 40 55 68 70 80 98

N103-70 17 20 26 37 42 66 73 83 92 83

E0540-80 8 11 13 18 22 26 — — — —

FC-218 S604-70 3 6 10 22 24 27 41 48 50 52

N674-70 11 14 18 39 49 60 77 88 92 95

LI 120-70 3 7 18 57 60 61 68 74 76 79

VI 164-75 6 8 4 7 7 25 39 56 58

CI 185-70 9 13 15 44 52 67 67 79 83 82 N 103-70 10 15 18 58 64 77 78 87 90 95

E0540-80 7 13 15 22 24 28
6. INTERACTIONS OF AGENTS WITH ELASTOMERS. 227

Table 14. Times tjQ and t2Q, in weeks, to obtain compression sets of 10 % and 20 % at the exposure

temperatures shown

ILALtVJ^UlC F Y n*~i c 1 1 rp t p mnp nr 1 1 rp LCI ll^Jvl uLUl t 50 °C 75 °C 100 °c 150 °C

Agent Elastomer ho '20 Mo ^0 ^0 ho ^0

HFC-227ea 5604-70 6.9 — 2.3 6.9 0.3 1.4 0.2 0.7

N674-70 1.9 5 0.7 2.2 0 0.1 0 0

L1120-70 6.2 13.1 1.9 4.1 0 0.1 0 0.1

VI 164-75 0 — — — — — 0.5 1 C11 85-70 1.2 10.3 0.7 3.4 0 0.1 0 0

N103-70 0.4 2.1 0.1 1 0 0 0 0

E0540-80 2 13.2 0.6 11.8 0.4 1.1 ~ —

HFC- 125 S604-70 4.5 12.1 2.3 7.5 0.3 1.8 0.1 0.4

N674-70 3.3 7 1 2.5 0.2 0.6 0 0

L1120-70 1.2 3.4 0.6 1.5 0 0 0 0

VI 164-75 0 -- -- -- 2.7 5.7 0.3 0.6 C 11 85-70 3.1 8.6 1.1 2.2 0.1 0.3 0 0

N103-70 0.2 1.8 0.3 1.1 0.1 0.4 0 0.1

E0540-80 1.8 1.2 9.4 0.2 1.4

FC-218 5604-70 3.8 8.6 1.1 8.9 0 0.6 0.1 0.2

N674-70 0.8 5.6 0.6 2 0 0.2 0 0

LI 120-70 2.3 4.9 0.3 0.8 0 0 0 0

VI 164-75 6.4 31.1 0.3 0.8

CI 185-70 1.3 0.9 2 0 0.1 0 0

N103-70 1 0.1 0.5 0 0 0 0

E0540-80 0 1.4 5.9 0 0.6
228 6. INTERACTIONS OF AGENTS WITH ELASTOMERS

Table 15. Activated process parameters and predicted values at 70 °C for compression sets C of

10 % and 20 %. Values of (AH/R)jo and (AH/R)2o are in Kelvins; times are in weeks

HFC- 125

FC-218

Elastomer

C = 10 C = 20

ln(t',o) (AH/R),o tio(70°C) ln(t'2o) (AH/R)20 t2o(70°C)

I S604-70 -14.28 5171 2.2 -10.73 4310 6.2

N674-70 -13.26 4498 0.9 -24.11 8421 1.5

L 11 20-70 -14.84 5386 2.3 • :9.07 6926 3 V 11 64-75 " - - - ~ -

C1185-70 -8.35 2770 0.8 -32.36 11359 2.1

N 103-70 -18.05 5496 0.1 -9.34 3256 1.2

E0540-80 -12.06 4083 0.9 -15.26 5890 6.7

5604-70 -16.82 5979 1.8 -12.82 5028 6.3

N674-70 -19.81 6820 1.1 -16.28 5921 2.6

LI 120-70 -8.72 2867 0.7 -10.45 3773 1.7

VI 164-75 -18.72 7357 15.1 -17.54 7196 30.9

C1185-70 -26.32 8963 0.8 -22.11 7883 2.4

N103-70 -5.34 1297 0.2 -11.63 3997 1

E0540-80 -13.88 4740 0.9 -26.13 9878 14.2

S604-70 -15.81 5542 1.4 -15.6 5865 4.5

N674-70 -3.64 1088 0.6 -23.98 8388 1.6

LI 120-70 -25.86 8627 0.5 -24.05 8285 1.1

VI 164-75 -10.55 4010 3.1 -12.28 5080 12.5

CI 185-70 -5.16 1743 0.9 -39.25 13915 3.7

N103-70 -30.96 10002 0.2

E0540-80 -32.78 12031 9.8
6. INTERACTIONS OF AGENTS WITH ELASTOMERS. 229

HFC-227ea

S604-70 - N674-70

—A- L11 20-70

V1 1 64-75

C1 185-70

N 103-70

E0540-80

20 -

% 0

-

-20 -

-40 -

-60 -

-80

1 10 100

Exposure time, weeks

Figure 12. Percent change in ultimate elongation for various elastomers vs. exposure time to agent HFC-227ea at 5.86 MPa and 75 °C.
230 6. INTERACTIONS OF AGENTS WITH ELASTOMERS

S604-70

N674-70

—Ar- L11 20-70 — V1 164-75

C11 85-70

N 103-70

E0540-80

40

-80 -H \ 1

1 10 100

Exposure time, weeks

Figure 13. Percent change in ultimate elongation for various elastomers vs. exposure time to agent HFC-125 at 5.86 MPa and 75 °C.
6. INTERACTIONS OF AGENTS WITH ELASTOMERS. 231

FC-218

-•- S604-70

N674-70 —A- L11 20-70

V11 64-75

C1 185-70

- N 103-70

E0540-80

-80 -^n \

1 10 100

Exposure time, weeks

Figure 14. Percent change in ultimate elongation for various elastomers vs. exposure time to agent FC-218 at 5.86 MPa and 75 °C.
232 6. INTERACTIONS OF AGENTS WITH ELASTOMERS . . .

Table 16. Percent change in ultimate elongation after short-term exposures at 5.86 MPa and 50, 100

and 150 °C

Exposure temperature 50 °C 100 °c 150 °C

Exposure time, weeks 1 2 4 1 2 4 1 2 4 6

Agent Elastomer

HFC-227ea 8604-70 -J.o 1 .u % 7 8 ^ 1

1

D.J - 7 S 1 .J - iniU.Df\ 11 .oa fi 7 O.Z

N674-70 9 1 -80. 7/ -7"^ z.^.z7 -Of, 7 -40 4 JL.y .8fl 7 -inn

Li! 20-70 1 Q 14 n -171 / .08 D.^ .41 4 .41 4

VI 164-75 0.

1

7 a u.o 1I J.U "K n .1-1.11 -J J.

7

.41 -t 1 .z7 -OZ.J -OO.J

CI 185-70 4 0 I .it 4 0 -7.7 -2.4 -27 5 -36 7 OJ xj.y

N103-70 f\ 1 0.

1

-6D.JS / .o -70 Q -inn

E0540-80 J.J 1 <

X .J J.v/ 1 ^

HFC- 125 S604-70 /.u -J.o - 1 .0 -iU.o 1 c\ a -o.D ft 4 7 1 -J.J 1 n A

N674-70 9 7 Z. / 8 7 -o. / 7n ^ 1

7

- 1 Z.J 4ft Q 74 A 08 7

L11 20-70 -u.o -IJ.O -U.O 414 r>

1 .U J J.U -o.O ft A 77 4 -jO.J - /j.l

VI 164-75 zz.u Q ^ V .J ^8 1 -JO. 1 -Dj. 1 -Jo. 1

ft -07.0 AO A - /U.J -o87/ . 7/

CI 185-70 1 1 .u ± fl *f .U 7 8 Z.o a Q 1 Q A 9'^ ft 41 n -OJ.2' -00.

U

N 103-70 90zu.un 7fi n 1 J .u 7Z.7Q 1 8 fl -J J. 1 .fxCi ft -7Q n -04 4 -inn

E0540-80 5.0 1.5 3.5 5.0 -1.5 -3.0

FC-218 3604-70 3.2 3.2 3.2 -15.6 -3.8 8.6 1.3 -6.2 -9.3 -4.4

N674-70 6.9 1.8 8.7 -6.3 -14.2 -32.2 -37.7 -59.1 -73.7 -81.6

L II 20-70 -8.1 -9.4 -12.5 -11.3 -19.4 -11.3 -47.9 -31.0 -43.2 -43.2

VI 164-75 6.1 7.8 2.2 13.0 -19.0 -50.9 -52.8 -72.2 -76.2

CI 185-70 11.0 6.9 -3.6 5.3 0.0 -11.7 -25.0 -41.4 -48.8 -58.6

N 103-70 22.0 -9.4 3.7 -4.1 -22.0 -33.5 -50.8 -72.9 -83.5 -89.0

E0540-80 48.0 1.5 0.0 6.4 -6.4 0.0
6. INTERACTIONS OF AGENTS WITH ELASTOMERS. . . 233

Table 17. Percent change in tensile strength after short-term exposures at 5.86 MPa and 50, 100 and

150 °C

Exposure temperature 50 °C 100 °C 150 °C

Exposure time, weeks 1 2 4 1 2 4 1 2 4 6

Agent Elastomer

HFC-227ea S604-70 13.0 19.0 15.0 56.0 27.0 32.0 -4.6 -12.9 -15.8 -26.6

N674-70 -1.5 -4.4 -1.5 9.2 9.2 -29.3 -8.6 -63.7 - -92.1

LI 120-70 19.0 18.0 16.0 24.0 -20.1 13.0 131.0 -9.5 -39.9 -49.3

V1164-75 ~ -10.1 -10.4 4.5 15.0 -26.5 -51.7 -42.9 -55.0 -56.9

C1185-70 8.7 -5.9 -1.2 26.0 6.9 -15.5 -19.6 -53.8 -74.3 -90.5

N103-70 22.0 32.0 35.0 50.0 22.0 8.4 4.5 -51.3 -88.2 ~

E0540-80 - - -0.8 29.0 -4.5 4.6 - - - ~

HFC- 125 S604-70 7.8 6.7 15.0 1.3 -3.2 1.3 -3.1 -6.7 -16.7 -32.3

N674-70 7.5 9.0 6.1 -7.7 4.6 -36.4 -32.8 -51.5 -72.8 -87.9

L 11 20-70 7.4 2.1 16.0 9.9 2.0 -2.0 -10.8 -25.9 -52.0 -76.1

V1164-75 - 0.8 -7.7 -75.2 -80.5 -76.9 -68.5 -70.0 -54.5 -70.3

C1185-70 5.5 3.6 0.7 0.0 -10.0 -17.5 -22.9 -32.3 -60.9 -82.6

N 103-70 31.0 39.0 29.0 20.0 12.0 1.1 -44.8 -52.7 -72.4 -

E0540-80 -1.4 1.5 -3.2 -10.4

FC-218 S604-70 15.0 25.0 7.0 39.0 43.0 67.0 4.5 -0.9 1.5 -3.1

N674-70 2.9 1.5 7.4 38.0 47.0 35.0 -16.0 -32.8 -42.0 -29.6

LI 120-70 13.0 21.0 3.8 27.0 26.0 42.0 -40.0 -18.1 -23.6 -28.3 V 11 64-75 -3.1 -4.5 -7.7 6.1 -29.8 -59.1 -59.5 -67.9 -63.2

CI 185-70 7.2 -1.0 -1.2 26.0 29.0 14.0 -18.1 -33.2 -34.9 -45.0 N103-70 15.0 -98.2 15.0 -83.9 48.0 44.0 -26.7 -45.5 -52.6 -55.0

E0540-80 2.3 30.0 17.0 27.0
234 6. INTERACTIONS OF AGENTS WITH ELASTOMERS

6.4 Testing of Elastomers and Lubricants in CF3I -

In this chapter the results of swelling measurements and mechanical property measurements obtained

for CF3I are reported. In the first phase of the project this agent was not investigated. It was included

later in the testing programs because it was found to be very effective in suppressing non premixed

flames. During the tests in this agent several anomalies were observed (the saturation vapor pressure

at the test temperatures was not reproducible, oil-like residue was found in the pressure vessels after exposure, etc.). The details of the observed anomalies are described in 6.4.1.

6.4.1 Anomalous Testing Results with CF3I. In the first lot of CF3I both swelling and mechanical

properties measurements were discontinued because of chemical purity problems. During the swelling

measurements a slightly yellow solid residue developed on the Pyrex windows at 75 °C which

changed from yellow to orange to a brownish red at 150 °C. Most of the elastomers were destroyed

and even the stainless steel fixtures used for the compression set measurements were severely corroded

after 1 week exposure at 100 °C. A high viscosity oil-like residue was found in the bottom of the

pressure vessels. It was concluded CF3I was degraded creating HI, HP and I2. Testing with a new lot of CF3I (Deepwater Lot #224940901) has been repeated, but our observations indicate that the

saturation vapor pressure of this chemical increases after exposures at elevated temperatures. This seems to be a consequence of chemical degradation, but to a much lesser degree than in the first lot. We found that during the exposure at a given temperature the vapor pressure of the agent was

constant. Then the CF3I was removed from the pressure vessel to a storage vessel so that the samples

(o-rings) could be withdrawn for mechanical testing. When the pressure vessel was refilled with the same amount of CF3I, the vapor pressure was significantly higher than originally, i.e., the mass of

CF3I required to attain the original pressure was reduced. The details of our observations are outlined

below.

The saturation vapor pressure of the new supply of CF3I was observed to increase after 1 week

exposure at both 75 °C and 100 °C. After filling a completely dark pressure vessel with approximately 0.9 kg of unused CF3I, the initial vapor pressure of the agent was measured as 1 .66 MPa at 76.9 °C. Approximately 0.9 kg of agent was recovered and stored in a previously empty and

evacuated steel tank with brass fittings for approximately 3 hours. The pressure vessel was then

refilled with approximately 0.3 kg of the previously used CF3I and the vapor pressure increased to 2.02 MPa at 76.2 °C. After another week of exposure, the vapor pressure was 2.01 MPa at 76.2 °C

and then approximately 0.3 kg was recovered from the pressure vessel. For the next time interval of

2 weeks, approximately 0.3 kg was filled and recovered from the pressure vessel and the vapor

pressure was 1.94 MPa at 76.2 °C. It is possible that saturation was not achieved during this exposure. A black oil was removed from the bottom of the pressure vessel and 0.8 kg of unused (i.e., fresh) CF3I was pumped into the vessel for the next exposure. The vapor pressure was measured as

1 .68 MPa at 76.0 °C and approximately 0.6 kg was recovered.

Similarly for the 100 °C exposures, the vapor pressure increased from 2.46 MPa at 99.0 °C to 2.79 MPa at 98.8 °C after recovery in the same steel tank, but the filling mass was approximately

constant. Note that the quantity of filling mass must be considered carefully because any mass in excess of the amount necessary for saturation will not increase the vapor pressure. Also, this pressure

vessel did have glass windows (although "UV free" PMMA sheets were used for filtering purposes)

and consequently some amount of blue and ultraviolet light may have reached the agent.
6. INTERACTIONS OF AGENTS WITH ELASTOMERS. 235

These observations during exposures at 75 and 100 °C indicate possible degradation of the CF3I,

and raise a flag about the chemical stability of the material. All swelling and mechanical measure- ments were made with Deepwater Lot #224940901.

6.4.2 Swelling Measurements in CF3I. The experimental method has been described in 6.2.2. All seven sets of crosslinked and uncrosslinked elastomers and lubricants were exposed to CF3I at temperatures 35, 70, and 105 °C.

In Figure 1 5 the solvent uptake of the elastomers is shown as a function of the vapor pressure at 35 °C. In Figure 16 the same data are displayed according to the Flory-Huggins representation. The

curves in Figure 16 show the fit of equation (2) to the experimental points. The values of the %

parameters for all the elastomers and lubricants are displayed in Table 18.

6.4.3 Durability Test Results in CF3I. Compression set and tensile measurements were made on

samples exposed to CF3I at 5.86 MPa and 75 °C, using the methods described in 6.3. The results are given in Tables 19-21.

Figure 17 shows log-log plots of the compression set vs. exposure time data in Table 19 for five elastomers in CF3I. (The data for elastomers N674-70 and VI 164-75 are not shown because the

compression set values were small or negative due to swelling.) The lines show the result of fitting

the measured values of compression set for exposure times from 1 to 32 weeks to eq. 5 by unweighted

nonlinear least squares. As with the other three agents, the data obtained at the longest exposure time

(58 weeks) have not been used in the fitting to allow comparison of the values measured at 58 weeks

with the results predicted from measurements at shorter times. The values of C^, x and P so obtained

are given in Table 12. It is seen from Table 19 and Fig. 17 that the compression set for S604-70 and

E0540-80 in CF3I is markedly higher after 58 weeks exposure than would be predicted from

extrapolations based on the behavior at shorter times. Figure 18 shows the data of Table 20, percent

change in ultimate elongation vs. exposure time for all the elastomers.

The results of tensile tests of elastomers exposed to CF3I at 5.86 MPa and 50 and 100 °C for 1, 2, and 4 weeks are shown in Tables 22 and 23. As with the other three agents, the results of tensile

tests of elastomers exposed to CF3I exhibited too much scatter to allow meaningful fitting for the

extrapolation of either time (relaxation functions) or temperature (activated process models).

6.5 Summary

In order to characterize the compatibility of elastomeric o-rings in the fire suppressant storage system,

swelling and mechanical property (compression set and tensile test) measurements were conducted and

data were generated on the changes in the properties of seven sets of crosslinked and uncrosslinked

elastomers (silicone, fluorosilicone, neoprene, fluorocarbon, two different butadiene-acrylonitrile

copolymers, ethylene-propylene-diene terpolymer) after exposure to HFC-227, HFC- 125 and FC-218.

In CF3I swelling measurements were conducted for the seven sets of elastomers plus three lubricants

(Krytox 240AC, Braycote 600 and Braycote 807). Exposures at 75 °C and 5.86 MPa were carried out

for 74 weeks (in the case of the CF3I for 58 weeks) to provide estimates of the long term resistance of

elastomers to in-service conditions.

The following sections summarize the swelling and mechanical property estimates of elastomer

performance in the fire suppressant agents. In addition, swelling performance of the lubricants is summarized. The important conclusion to be derived from these summaries and the detailed data of
236

Table 18.

6. INTERACTIONS OF AGENTS WITH ELASTOMERS

Flory-Huggins interaction parameters % for elastomers and lubricants in CF3I

Elastomer/ X (35 °C) X (70 °C) compatibility

Lubricant

uSi 1.06 1.29 fair

xSi 1.11 1.40 fair UN206 0.94 1.20 fair xN206 1.00 1.25 fair

uFSi 1.20 1.41 good

xFSi 1.28 1.48 good

uFKM 1.43 1.67 good

xFKM 1.47 1.72 good

uCR 1.35 1.48 good

xCR 1.40 1.53 good

uN926 1.03 1.15 fair xN926 0.95 1.21 fair uEPDM 1.08 1.18 fair xEPDM 1.13 1.25 fair

Krytox 240AC 2.11 2.12 good

Braycote 600 2.24 2.34 good

Braycote807 1.84 1.86 good
6. INTERACTIONS OF AGENTS WITH ELASTOMERS. 237

^1

0.5

0.4

0.3

0.2

0.1

o

1 xSi

xN206

V xFSi yv J. 1—J ± T xFKM

xCR 9

xN926

- A xEPDM A

—

• A V

o

A V

V T

V T

1 1 0.0

0.35 0.40 0.45 0.50

p/MPa

Figure 15. Weight fraction Wj of CF3I as a function of pressure for various crosslinked

elastomers at 35 °C.
238 6. INTERACTIONS OF AGENTS WITH ELASTOMERS

0,

In (p/p )

0.0

0.1

-0.2

-0.3

\

'

s,

N

\

\

\

's ^

*. \ ^ \ \ \

\\ \ \

rnLj^

T

\ \

o xSi

\

^^^^^

\ \ \ .

\ " \

• xN206

\ \ \ \

\ \ \ \ \

\ \

^.

\ 'i

\

"

\ \ \ V xFSi

\ \ \

^.

\ \ \

'

\

'

1 xFKM \ \ \ \ ^ 9¥

\ \ \

xCR

\ \ \

\ \ \

\. \

\ \ \

\ \ \

\ xN926 \. \ \ \ \ \ \ w

\\ \ w

A xEPDM

t

\

\ mi

\

'\

'\

'\

1 1

\ \ \ \ \

\

\ \ \ \ \ \

'

\ \ \ \ w

^'V \ \\

0.5 0.6 0.7 0.8 0.9 1.0

w.

Figure 16. Solvent activity in various crosslinked elastomers in CF3I as a function of polymer

weight fraction W2 at 35 °C. The curves show the least squares fits according to Equation (2).
6. INTERACTIONS OF AGENTS WITH ELASTOMERS. . . 239

Table 19. Compression set (C) after 1, 2, 4, 8, 16, 32 and 58 week exposures at 5.86 MPa and

75 °C in CF3I

Elastomer weeks

1 2 4 8 15 32 58

S604-70 2 6 8 14 18 24 33

N674-70 -3 12 24 35 70 80 81

Lll 20-70 13 20 36 40 48 58 65

V1164-75 -11 -5 -6 -6 -2 -1 2

CI 185-70 9.1 20 43 50 76 78 81

N 1 03-70 9 27 47 58 88 88 92

E0540-80 2 11 17 22 33 39 57

this and the prior reports (McKenna et al, 1994) is that for each of the final four agents, there exists at least one compatible elastomer and grease. It is up to the design engineer to make the appropriate

selection.

6.5.1 Swelling measurements. The swelling measurements were used to make compatibility ratings

based on the x values determined for each polymer-agent system at 35 °C. Small % values correspond

to relatively high solubility or, for present purposes, bad compatibility. The results of the swelling

measurements are summarized below for the four agents. An important finding is that for each agent

there are several elastomers that show good compatibility, hence providing design flexibility for agent

storage vessels.

HFC-227ea

55 % butadiene-45 % acrylonitrile (N206), neoprene and ethylene-propylene-diene have good

compatibility, and silicone, fluorosilicone and fluorocarbon show fair compatibility. 85 % butadiene15 % acrylonitrile (N926) has bad compatibility. Lubricants Braycote 600, Braycote 807, and Krytox

240AC are fairly compatible.

HFC- 125

Sihcon, 55 % butadiene-45 % acrylonitrile (N206), fluorosilicone, neoprene, 85 % butadiene15 % acrylonitrile (N926) and ethylene-propylene-diene have good compatibility, and fluorocarbon has

fair compatibihty. Lubricants Braycote 600 and 807 have good compatibility, and Krytox 240AC is fairly compatible.

FC-218

All the elastomers have good compatibility. Lubricants Braycote 600, Braycote 807, and Krytox

240AC have good compatibility.
240 6. INTERACTIONS OF AGENTS WITH ELASTOMERS

Table 20. Percent change in ultimate elongation after 1, 2, 4, 8, 16, 32 and 58 week exposures at 5.86 MPa and 75 °C in CF3I

Elastomer

1 2 4

weeks

8 15 32 58

5604-70 0.0 10.2 17.2 6.5 3L7 -61.3 -17.2

N674-70 1.8 -2.4 -8.1 -8.1 -51.5 -56.6 -78.6

LI 120-70 -20.6 -32.5 -21.9 -36.9 -21.9 3.8 -11.3

VI 164-75 18.4 14.5 21.8 21.8 -10.1 11.2 -8.9

CI 185-70 1.6 1.6 -27.5 -1.2 -14.2 -18.2 -32.3

N103-70 14.7 9.0 -2.4 3.3 -27.3 -67.8 -78.4

E0540-80 5.0 5.0 1.5 -5.0 1.5 12.9 11.4

CF3I

Fluorosilicone, fluorocarbon and neoprene have good compatibility, and silicone 55 % butadiene

45 % acrylonitrile (N206), 85 % butadiene- 15 % acrylonitrile (N926) and ethylene-propylene-diene

have fair compatibility. Lubricants Krytox 240AC, Braycote 600 and Braycote 807 have good

compatibility.

6.5.2 Mechanical Property Measurements. Two types of experiments were performed to produce

data on residual mechanical properties of the elastomers: compression set and tensile test measurements. In order to study the effect of temperature short term (1,2 and 4 weeks) exposures to the

four agents were performed at 50, 75 and 100 °C. An activation energy analysis was used to determine the temperature dependencies of the compression set data. The results of long-term

exposures for 8, 16, 32, 48, and 74 weeks (58 weeks for CF3I) carried out at 75 °C and 5.86 MPa

have been analyzed using a KWW stretched exponential creep function. This analysis allows

extrapolation to longer exposure times.

6.5.3 Compression Set Measurements. The compression set measurements showed significant

differentiation between the results obtained for the elastomers in each agent after exposure at 5.86

MPa and 75 °C.

HFC-227ea

After 74 weeks exposure the compression set was small for silicone (30 %), moderate for

ethylene-propylene-diene and standard nitrile (43 % and 50 %), and large for fluorosilicone, lowtemperature nitrile and neoprene (64, 65 and 74 %). In the case of fluorocarbon swelling was

observed.
6. INTERACTIONS OF AGENTS WITH ELASTOMERS. 241

Table 21. Percent change in tensile strength after 1, 2, 4, 8, 16, 32 and 58 week exposures at 5.86

MPa and 75 °C in CF3I

Elastomer

1 2 4

weeks

8 15 32 58

S604-70 -1.1 -0.9 5.8 -9.4 1.8 -74.2 -34.6

N674-70 4.5 3.0 3.0 -68.8 -48.5 -34.3 -69.4

LI 120-70 -28.4 -40.9 -30.9 -44.1 -33.5 -6.4 -15.8

VI 164-75 -20.4 -26.2 -22.8 -22.8 -50.0 -30.7 -49.2

CI 185-70 -1.2 0.2 -32.0 -2.0 -30.3 -23.6 -45.5 N 103-70 25.4 12.4 17.9 23.5 -14.5 -51.3 -60.0

E0540-80 L5 -5.5 -0.8 -13.6 -21.4 -13.6 -7.4

HFC- 125

After 74 weeks exposure the compression set was small for fluorocarbon (20 %), moderate for

silicone and ethylene-propylene-diene (39 % and 49 %), and large for fluorosilicone, standard nitrile,

neoprene and low-temperature nitrile (73, 80, 88 and 89 %).

FC-218

After 74 weeks exposure the compression set was small for fluorocarbon (22 %), moderate for

silicone and ethylene-propylene-diene (41 % and 46 %), and large for fluorosilicone, standard nitrile,

neoprene and low-temperature nitrile (79, 79, 85 and 92 %).

CF3I

After 58 weeks exposure the compression set was moderate for silicone and ethylene-propylenediene (33 % and 57 %) and large for fluorosilicone, neoprene and low-temperature nitrile (65, 81 and

92 %) In the cases of standard nitrile and fluorocarbon swelling was observed.

As a note of caution, the silicone, standard nitrile and ethylene-propylene-diene elastomers exhibited

abrupt increases in compression set between 48 and 74 week exposures in some of the agents. The

reader is referred to Tables 9 and 19 and Figures 9-11 and 17 for details. This indicates the difficulty

of extrapolating the results to longer times than the measurements.

6.5.4 Tensile Testing. The ultimate elongation data showed very large scatter with the exposure time

at all temperatures. As a result, such standard test results are not very amenable to reliable interpretation of the agent-elastomer compatibility.
242

Table 22.

6. INTERACTIONS OF AGENTS WITH ELASTOMERS . .

Percent change in ultimate elongation after short-term exposures to CF3I at 50 and 100 °<

Exposure

ICllipcI cL- ture

50 °C 100 °C

Exposure

time,

W ^wfVo

1 L

A

H'

11 L

A

ddaiuiiier

14.0 17.0 12.0 24.0 6.5 28.0

N674-70 4.8 4.8 6.9 -10.2 -16.3 -46.1

LI 120-70 12.0 16.0 1.9 12.0 12.0 12.0

V 11 64-70 13.0 17.0 11.0 7.3 7.8

CI 185-70 11.0 9.3 9.3 -27.5 -11.7 -27.5

N103-70 8.2 18.0 16.0 -2.4 -8.2 -41.2

E0540-80 7.9 9.4 11.0 15.0 15.0 7.9
6. INTERACTIONS OF AGENTS WITH ELASTOMERS. 243

100

c

U

a>Q.«

cg

'w

(i) Q.

E

oo

10

• S604-70

L1 120-70

A C1 185-70

N 103-70

E0540-80

fit S604-70

fit L1 120-70

fit C1 185-70

fit N 103-70

fit E0540-80

CF3I

10

Exposure time, weeks

100

Figure 17. Compression set for various elastomers vs. exposure time to agent CF3I at 5.86 MPa

and 75 °C.
244

Table 23.

6. INTERACTIONS OF AGENTS WITH ELASTOMERS .

Percent change in tensile strength after short-term exposures to CF3I at 50 and 100 °C.

Exposure

temperature

50 °C 100 °C

Exposure

time,

weeks

1 2 4 1 2 4

Elastomer

Sd04-70 25.0 27.0 21.0 3.9 -10.3 4.2

N674-70 6.0 3.0 7.5 4.6 -1.6 -26.6

LI 120-70 15.0 13.0 3.6 -10.1 -9.5 -3.6

VI 164-70 -9.1 -4.5 -26.2 -32.1 -29.8

CI 185-70 5.8 7.2 5.2 -6.9 -10.8 -27.6

N103-70 7.5 16.0 22.0 18.0 16.0 -8.8

E0540-80 9.1 12.0 -0.8 -5.5
6. INTERACTIONS OF AGENTS WITH ELASTOMERS. . . 245

CF3I

S604-70

N674-70

L11 20-70

V1 164-75

C1 185-70

N 103-70

E0540-80

q

COco

<D

E

3

c

oc

c

CDO

0) CL

-100

10 100

Exposure time, weeks

Figure 18. Percent change in ultimate elongation for various elastomers vs. exposure time to agent CF3I at 5.86 MPa and 75 °C.
246 6. INTERACTIONS OF AGENTS WITH ELASTOMERS

6.6 References

ASTM, Designation D 395-89, "Standard Test Methods for Rubber Property - Compression Set,"

Annual Book of ASTM Standards, 09.01, 24 (1989).

ASTM, Designation D 412-87, "Standard Test Methods for Rubber Properties in Tensions,"

Annual Book of ASTM Standards, 09.02, 100 (1990a).

ASTM, Designation D 1414-90, "Standard Test Methods for Rubber 0-Rings," Annual Book of

ASTM Standards, 09.02, 181 (1990b).

Boyer, R.F., "Deswelling of Gels by High Polymer Solutions," J.Chem.Phys., 13, 363 (1945).

Brotzman, R.W. and Eichinger, B.E., "Volume Dependence of the Elastic Equation of State. 3. Bulk-Cured Poly(dimethylsiloxane)," Macromolecules 15, 531 (1982).

Brotzman, R.W. and Eichinger, B.E., "Swelling of Model Poly(dimethylsiloxane) Networks,"

Macromolecules, 16, 1131 (1983).

Flory, P.J., "Thermodynamics of High Polymer Solutions," / Chem. Phys., 10, 51 (1942).

Flory, P..I., Principles of Polymer Chemistry, Cornell University Press, Ithaca, New York (1953).

Flory, P.J., "Thermodynamics of Polymer Solutions," Disc. Faraday. Soc, 49, 7 (1970).

Flory, P.J. and Erman, B., "Theory of Elasticity of Polymer Networks 3.,"Macromolecules 15,

800 (1982).

FI017, P.J. and Rehner, J., Jr., "Statistical Mechanics of Cross-Linked Polymer Networks,"

J.Chem.Phys., 11, 521 (1943).

Gallagher J., McLinden, M., Huber, M., and Ely, J., NIST Standard Reference Database 23:

Thermodynamic Properties of Refrigerants and Refrigerant Mixtures Database (REFPROP), Version

3.0, U.S. Department of Commerce, Washington, DC (1991).

Gee, G., Herbert, J.B.M. and Roberts, R.C., "The Vapor Pressure of a Swollen Cross-Linked

Elastomer," Polymer 6, 51 (1965).

Horkay, F. and Zrinyi, M., "Studies on Mechanical and Swelling Behavior of Polymer Networks

on the Basis of the Scaling Concept. 4. Extension of the Scaling Approach to Gels Swollen to Equilibrium in a Diluent of Arbitrary Activity," Macromolecules 15, 1306 (1982).

Horkay, F, Hecht, A.-M. and Geissler, E., "The Effects of Cross-Linking on the Equation of

State of a Polymer Solution," J. Chem. Phys., 91, 2706 (1989).

Huggins, M.L., "Some Properties of Solutions of Long-Chain Compounds," J. Phys. Chem., 46,

151 (1942).

James, H.M. and Guth, E., "Theory of Elastic Properties of Rubber," J. Chem. Phys., 11,455

(1943).

Jandel Scientific, SigmaPlot Scientific Graphing Software, Version 5.0, (1992).

McKenna, G.B., Waldron, W.K. and Horkay, F., "Elastomer Seal Compatibility," in Evaluation

of Alternative In-Flight Fire Suppressants for Full-Scale Testing in Simulated Aircraft Engine Nacelles

and Dry Bays, Eds. W.L. Grosshandler, R.G. Gann and W.M. Pitts, U.S. Department of Commerce,

NIST SP 861, pp 729-763 (1994).

McKenna, G.B. and Crissmann, J.M., "Thermodynamics of Crosslinked Polymer Networks: The

Anomalous Peak in Swelling Activity Measurements," Macromolecules 1992, ed by J. Kahovec, VSP,

Utrecht, The Netherlands, pp. 67-81 (1993).

McKenna, G.B., Flynn, K.M. and Chen, Y., "Experiments on the Elasticity of Dry and Swollen

Networks: Implications for the Frenkel-Flory-Rehner Hypothesis," Macromolecules 22, 4507 (1989).
6. INTERACTIONS OF AGENTS WITH ELASTOMERS. 247

McKenna, G.B., Flynn, K.M. and Chen, Y., "Swelling in Crosslinked Natural Rubber: Experimental Evidence of the Crosslink Density Dependence of %," Polymer 31,1937 (1990).

Press, W.H., Flannery, B.P., Teukolsky, S.A., and Vetterling, W.T., "Numerical Recipes,"

Cambridge University Press (1986).

Robin, M.L., Thermodynamic and Transport Properties of FM-200, Great Lakes Corporation

(1992).

Williams, G., and Watts, D.C. Trans. Faraday Sac. 66, 80 (1970).
1I
7. AGENT STABILITY UNDER STORAGE

Richard H. Harris, Jr. Building and Fire Research Laboratory

Contents

Page

7. AGENT STABILITY UNDER STORAGE 249

7.1 Introduction 250

7.2 Experimental 251

7.2.1 Test Matrix 251

7.2.2 Materials 251

7.2.3 Preparation of Cylinders for Testing 252

7.2.4 Spectral Analysis Equipment and Procedure 256

7.2.5 Use of the FTIR Spectra to Determine Agent Degradation 256

7.2.6 Determination of Uncertainty 266

7.3 Results 269

7.3.1 FC-218 269

7.3.2 HFC- 125 269

7.3.3 HFC-227ea 281

7.3.4 Summary of Fluorocarbon Area Data 281

7.3.5 lodotrifluoromethane (CF3I) 282

7.3.5.1 Changes in the Agent Peak Area of CF3I in the 53.3 Pa Spectra for the

Different Conditions 290

7.3.5.2 Changes in the Impurity Peak Areas Appearing in the 5330 Pa Spectra .... 304

7.3.6 Summary of the CF3I Area Data 305

7.4 Conclusions 338

7.5 Acknowledgments 338

7.6 References 338

Appendix A. Initial and Final FTIR Spectra for HFC- 125 340

Appendix B. Initial and Final FTIR Spectra of HFC-227ea 349

Appendix C. Initial and Final FTIR Spectra of CF3I Tested in the Dry Condition without

Copper 362

Appendix D. Initial and Final FTIR Spectra of CF3I Tested in the Dry Condition with

Copper 378

Appendix E. Initial and Final FTIR Spectra of CF3I Tested in the Moist Condition without

Copper 390

Appendix F. Initial and final FTIR Spectra of CF3I Tested in the Moist Condition with

Copper 397
250

7.1 Introduction

7. AGENT STABILITY UNDER STORAGE

1

Significant losses in fire suppression effectiveness and increases in toxicity are possible if a fire extinguishing agent degrades during multi-year storage. Halon 1301 is known to be stable in metal

containers for many years, and any trace degradation products do not affect its fire suppression

effectiveness. For candidate replacement agents, comparable data are needed, reflecting the storage

conditions of elevated temperature and pressure. The storage environment fosters conditions which may have an adverse effect on the stability of

halon replacements. Stored chemicals may engage in oxidation-reduction reactions, hydrolysis, and

other corrosive interactions with metal cylinders. They are also subject to unimolecular decomposition

and attack by reactive impurities in the agent. Water and oxygen, for example, will sorb to surfaces of

cylinders and transfer lines and can never be completely excluded. These sources of instability, along

with the possibility of catalytic interactions with the cylinder walls, can promote the evolution of

undesirable products and a concomitant loss of fire suppression effectiveness. Toxicity and corrosive- ness are particularly important concerns with respect to halogenated compounds, due to the tendency

to liberate hydrogen halide in the process of degradation.

Section 6 of NIST Special Publication 861 (Peacock et ai, 1994) detailed a screening test for the

stability of thirteen agents and eight metals, with the purity of the agent determined by infrared

spectral analysis. This test involved storing the agents in PTFE-lined cylinders for one month at a temperature of 150 °C (300 °F). The infrared spectra of the original and "aged" samples were

compared. The results of the aging study in combination with other studies of the thirteen agents

resulted in the selection of four agents for continued study. These agents were HFC- 125, FC-218,

HFC-227ea, and CF3I. Also, in Section 7 of the above publication (Ricker et ai, 1994) three of the

eight metals studied were chosen as being the best with respect to corrosion. Additionally, the

sponsors requested a titanium alloy as a candidate storage metal and copper as an additive for CF3I be

investigated in this study.

Although the screening test was an appropriate method for short-term evaluation, the promising

candidate agents need a more rigorous examination. A more comprehensive, long-term evaluation is needed to give stability and degradation information as a function of time and temperature. The

information obtained in this study, in combination with information obtained in the studies described

in Sections 5 and 6, can then be used in selecting the most thermally stable agents and/or appropriate

storage materials.

In this study, samples of the four candidate agents with each of the four storage metal candidates

and the copper additive in CF3I were evaluated in pressurized cylinders at various temperatures and

conditions. In order to allow for potential interactions analogous to actual storage conditions, a measured amount of metal (with separate tests for each candidate cylinder metal) was introduced into the containers prior to the experiments. The vessel and its contents were stored at ambient conditions

or in ovens at elevated temperatures for as many as 52 weeks as the project time allowed. After

specified aging times the cylinders were removed from their respective environments, cooled to ambient conditions, and an infrared (IR) spectrum of the aged sample was collected. Degradation of

the sample would be indicated by a systematic decrease in the absorbance of peaks attributable to the

agent and/or the appearance of new peaks in the IR spectrum of the aged agent.

This report gives the details of the test procedure and a comparison of agent absorbance band

areas from low concentration spectra. Also, presented are a comparison of absorbance bands in high

density spectra for impurities present in the agents or produced as a result of degradation. These data

provide a quantification of any degradation of the agents during long-term storage.
7. AGENT STABILITY UNDER STORAGE 251

7.2 Experimental

7.2.1 Test Matrix. Table 1 shows the test matrix. This matrix was designed to utilize oven space,

increase the number of exposure temperatures, and expand the exposure conditions to include moisture

and addition of copper to CF3I. The sponsors designed this test matrix to include all agents, all metals, and the various exposure temperatures. The resulting test matrix required 68 sample cylinders.

This test procedure was agreed to by NIST

Since the agents might interact with the metal storage cylinder during long term storage, it was

important to evaluate the selected agents in the best metals found in the previous screening study.

Containers constructed of each metal would have been ideal for this study, but the cost and availability

of cylinders necessitated using metal coupons inside an inert cylinder. In this approach, an amount of

metal with the surface area roughly equal to the storage cylinder was introduced into PTFE-lined

stainless steel cylinders.

The manufacturers of organoiodide compounds have routinely stored these compounds in the

presence of copper (Dierdorf, 1995.) Iodine, if formed during storage, will react with copper to produce cuprous iodide, which is virtually insoluble in most solvents. The sponsors requested that copper be tested as an inhibitor in the CF3I agent. With this in mind, the test matrix incorporated the

addition of copper coupons in certain cylinders, in addition to the test metals.

Since CF3I appeared to be the least stable agent in the previous screening study, the addition of

small amounts of water to some of the test cylinders was incorporated into the test matrix for this agent. In any realistic filling method of storage containers, the presence of residual water in lines and

other components of a filling apparatus can be expected. The sponsors requested that the study

include the addition of microliter quantities of water into certain cylinders to determine what affect, if any, water has on the degradation of CF3I.

The temperature dependence of the degradation was also incorporated into the test matrix. This

temperature dependence can be dramatic, especially for catalytic degradation. If by-products do form

then knowledge of the amounts produced from degradation at different temperatures can be used

extrapolate to longer storage times.

7.2.2 Materials. The four agents that were tested for their long-term storage stability are listed in Table 2. They were typical of production grade rather than ultra-pure research samples. All agents were used as received from the manufacturers. Different lots of FC-218 (R-218'),

HFC-227ea (FM-200), and CF3I (Triodide) were used in the study. The lots of each agent were compared to each other by comparing their high density Fourier transform infrared (FTIR) spectra

(5330 Pa). No differences were found in the lots of FC-218 or HFC-227ea. However, COj was found

in the lot of HFC- 125. COj and CF3H were found in the lots of CF3I. More detail on the amounts of

impurities in the lots will be given in 7.2.5. The five metals that were exposed to each of the agents were the same as those used in the study

described in Section 5, Table 1. They were Nitronic 40 (N40), Ti-15-3-3-3 (Ti), C4130 alloy steel (C4130), Inconel 625 (1625), and CDA 110. The metals were supplied by Metal Samples Co., Inc. as coupons 10.2 cm long, 0.8 cm wide, and 0.2 cm thick. In this study, the C4130 alloy steel coupons

had a small hole drilled into one end. They were then sent to Walter Kidde Aerospace for phosphate

'Certain trade names and company products are mentioned in the text or identified in an illustration in order to specify adequately the experimental procedure and equipment used. In no case does such

identification imply recommendation or endorsement by the National Institute of Standards and

Technology, nor does it imply that the products are necessarily the best available for the purpose.
252 7. AGENT STABILITY UNDER STORAGE

Table 1. Test matrix

Agent/Test Metal/Temperature

Condition

Blank Nitronic-40 Ti-15-3-3-3 C4130 Inconel 625

FC-218 IT IT IT IT IT

HFC- 125 2T 2T 2T 2T 2T

HFC-227ea 3T 3T 3T 3T 3T

CFjI/a, 3T* 3T* 3T* 3T* 3T*

CFjI/b, 3T* 2T 2T 2T 2T

2T* IT IT IT IT

CF3l/b2 2T* IT IT IT IT

Key:

a = w/o Cu T, = 23 °C IT = T,

b = w/Cu ^2 = 100 °C 2T = T„ T4

1 = dry T3 = 125 °C 2f* = T„ T4

2 = moist T4 = 150 °C 3T

3T* = T„

= T„ T2,T,

Table 2. Agents used in agent stability study

Agent Supplier Lot Numbers

R-218 (FC-218) 3M Chemolite Center 10 L- 1 2677, L- 13201

HFC- 125 Allied Signal Chemicals 835

FM-200 (HFC-227ea) Great Lakes Chemical Corp. 92-002-356, 93-200-278

Triodide (CF3I) Deepwater 224940901,226941712,

226941891

treatment prior to being introduced into the cylinders. The phosphate treatment renders the C4130

alloy steel less susceptible to rusting.

The storage cylinders were constructed of stainless steel, lined with polytetrafluoroethylene

(PTFE), and had a 1000 ml capacity. All new cylinders were assembled with high temperature/high

pressure stainless steel valves and end plugs, then heated to 150 °C (300 °F) for at least 48 hours with

the valves completely open. For re-used cylinders, the heating time was cut to 24 hours.

7.2.3 Preparation of Cylinders for Testing. Each agent/metal combination was exposed to the

various temperatures specified in the test matrix. Testing conditions dictated that the maximum

exposure temperature be no higher than 150 °C, and that the room temperature fill pressure be
AGENT STABILITY UNDER STORAGE

Relief

Valve

Regulated

Nitrogen

Storage Cylinder

Pressure

Gage

Agent

Supply

^

House

Vacuum

Figure 1 . Experimental apparatus for agent filling study.
254 7. AGENT STABILITY UNDER STORAGE

4.13 MPa (600 psia). In order to assure that the storage cylinder pressure not exceed the pressure

rating on the cylinder or valve when exposed to the elevated temperature, a conservatively low amount

of agent was placed in each cylinder. The cylinders were filled with agent up to a maximum pressure

of the saturation vapor pressure of the agent at room temperature. Nitrogen was then added to achieve

the final pressure. Since small amounts of additional liquid would vaporize at the elevated temperature only a single phase would exist anyway, though a significant pressure increase would be observed.

The effect of the system pressure on degradation was not explored, but is thought to be small.

To assure that the metal coupons were free of oil and markings, they were immersed and stirred

in dichloromethane solvent. The coupons were removed from the solvent, stripped with clean paper

towels, and dried at room temperature. The phosphate-treated C4130 alloy steel coupons were used as received from Walter Kidde Aerospace. After cleaning, care was taken to handle the coupons with

gloved hands only.

A filling procedure was designed to insure low levels of contaminants (including room air) and a reasonably accurate measurement of the amounts of material introduced into the cylinder. Highly

accurate measurement of amounts of agent and/or nitrogen were not required in this method, since it is the change in spectra that was desired. Figure 1 is a schematic of the filling apparatus. It consisted of

a three-valve manifold, an absolute pressure gage, and filling lines, all constructed from stainless steel and capable of handling high pressures. The lines to the agent supply tank and the house vacuum

were not high pressure rated and were isolated by valves. The following steps outline the filling

procedure.

1. Thirty pieces of each metal were weighed and placed into their respective cylinders. The number

of coupons was fixed at 30 for each cylinder because that amount is approximately the surface area of the interior of the cylinder. Overlapping of the coupons in the cylinder effectively

decreased the exposed surface area, but this decrease, though not calculated, should be similar for

all tests, except those containing the additional copper coupons. The end plug threads were wrapped with PTFE tape and tightened. The cylinder valves were opened and the cylinders were

heated for at least 12 hours at 150 °C. Immediately, the cylinder valves were closed. When the

cylinders were cool enough to be handled with gloves, they were attached to the sampling system

used for the FTIR analyses shown in Figure 2. The cylinders were purged with Nj, then

evacuated to an absolute pressure of less than 1.3 Pa (0.01 torr). Each evacuated cylinder was

weighed.

2. The evacuated storage cylinder was attached to the filling apparatus.

3. Certain cylinders required the addition of water (Table 1). A fitting was designed with a rubber

septum. The fitting was attached to the cylinder with Nj purging the connection. After the fitting was tightened, the cylinder valve was opened. A syringe containing 100 |iL of distilled HjO was

inserted through the septum and injected. The cylinder valve was closed and the "evacuated"

cylinder was then attached to the filling apparatus.

4. The filling apparatus was evacuated with the house vacuum which was capable of reduced

pressures near 34 Kpa. Then the valve to the house vacuum was closed.

5. The filling apparatus was pressurized with nitrogen in excess of 1.4 Mpa. The pressure was

released through the relief valve.
7. AGENT STABILITY UNDER STORAGE 255

6. The filling apparatus was then evacuated again with the house vacuum. The residual room air in the apparatus was less than 1 % of the original amount. , . . .,

-

7. The agent inlet valve was opened slowly until the pressure in the filling apparatus reached the

vapor pressure of the agent at ambient temperature (the agent inlet line was pre-charged with

agent). The agent inlet valve was closed. This dilutes any residual air again. The house vacuum

was opened to remove the agent and any remaining residual air. The agent inlet valve was

opened slowly until the pressure again reached the agent's vapor pressure. The valve to the

evacuated storage cylinder was slowly opened and agent flowed in. Both the inlet valve and the

cylinder valve must be opened slowly to avoid the possibility of drawing liquid agent into the

apparatus. The system was allowed to equilibrate for one to two minutes, the pressure recorded,

and the agent inlet valve closed.

8. The mass of agent in the storage cylinder was determined by removing the first cylinder filled in a series for a given agent and weighing it. The cylinder was then reattached to the filling

apparatus. With the cylinder valve closed, the filling apparatus was again evacuated and purged

with Nj. Then the cylinder valve was opened and the cylinder was filled to nominally 4.2 Mpa

with Nj. This pressure was slightly higher than the prescribed target pressure of 4.13 Mpa. This

allowed a small amount of the agent/Nj mixture to be removed from the cylinder for the initial FTIR analysis. The total pressure was recorded and the valve to the storage cylinder closed.

9. Proceeding with this method for each cylinder was time consuming and introduced the chance of

contamination from air. Since the exact amount of agent in each cylinder at the start of aging was not considered critical, each succeeding storage cylinder filled in a series was attached only once and filled with both agent and Nj. The agent pressure was recorded for all cylinders prior to introducing the N2.

10. The high pressure gas was bled from the filling apparatus and the cylinder disconnected and

weighed. At this point, a new cylinder was connected and filled starting with Step 1 above.

11. After FTIR analyses were completed for a series of storage cylinders, they were placed in their

proper storage environment as specified in the test matrix. Ambient cylinders were placed on

laboratory benches where the temperature remains at 23 ± 1 °C. Cylinders at elevated temperature conditions were placed in laboratory/production ovens with forced convection airflow. The

ovens were capable of maintaining a temperature uniformity of ±1 °C at 100 °C and ± 2 °C at 200 °C.

12. When a series of aged cylinders required analysis, they were removed from their respective

storage environment. Cylinders at elevated temperature storage were cooled to ambient temperature (at least 2 hours for those at 150 °C). While the cylinders were cooling, the FTIR was

prepared for the analyses. After analyses, the cylinders were weighed and the weights recorded.

Before being put back into storage, the cylinders were checked for leaks around the valves and

endcaps.

13. The cylinders remain in their respective environments pending the desire for longer-term results.

Tables 3-9 list for each cylinder the metal mass, the agent pressure, initial agent/N2 pressure, and

the mass before and after storage for all of the agents as well as for all the different environments of
256 7. AGENT STABILITY UNDER STORAGE

the CF3I. As mentioned above (steps 8 and 9), only the first cylinder in a given series of agent was

actually weighed to see if the agent weight was approximating what was expected from the ideal gas

law. For the remaining cylinders in a series the agent mass can be estimated from the agent density at the recorded agent pressure and ambient temperature (along with the fixed 1000 Ml volume of the

cylinders. The error in this calculated value is the combination of the uncertainty in the agent pressure

reading which is estimated to be ± 0.02 Mpa (± 2.5 psia), deviations of the agent temperature from

ambient temperature (at least two minutes were allowed for the system to equilibrate), and the error associated with the equation of state. Here, generalized compressibility charts were used (Balzhiser

et al., 1972). Where the tables have a mass value in parentheses the weight was calculated from the

measured vapor pressure. The mass before storage was the cylinder weight after the initial spectrum

had been taken and the cylinder was ready to go into storage. The mass after storage was the weight

of the cylinder after the final spectra were obtained. The difference in the above two masses is thus

the amount of agent/Nj lost as a result of analyses and leakage. Leakage was a problem over the course of the study for roughly 17 of the 68 cylinders in the study. The leakage resulted primarily

from cylinders that were reused. The overall validity of the test results was not affected.

7.2.4 Spectral Analysis Equipment and Procedure. The IR analyses were done with a Galaxy

Series 7000 FTIR Spectrometer fitted with a KBr beamsplitter and a narrow band mercury cadmium

telluride (MCT) detector. The spectral range was from 600 to 4000 cm"' with a resolution of 1 cm''. A variable path 20-meter gas cell with a volume of 5.4 L and a path length of 10 m was used. A

diagram of the sampling system used for the FTIR analyses is shown in Figure 2. All cylinders were

analyzed after cooling to ambient temperature. The gas cell was at 105 ± 1 °C for all analyses. The

cylinders were connected to the inlet of the gas cell on the FTIR spectrometer. The lines and gas cell were purged with gaseous from a liquid nitrogen cylinder. A vacuum of less than 1.3 Pa

(0.01 torr) was drawn on the gas cell and inlet line up to the cylinder valve. A positive shutoff toggle

valve and a micro-metering valve were placed in line between the cylinder valve and the gas cell to accurately control the flow of the agent into the gas cell. The micro-metering valve allowed precise

filling of the gas cell and provided reproducible filling of the gas cell. The agent/Nj was introduced

into the gas cell to an absolute pressure of 53.3 ±0.1 Pa (0.400 ± .001 torr). A spectrum was taken,

then the cell was emptied. Two more replicates were done in the same manner. After the third

replicate additional agent/Nj was added to fill the gas cell to 5333 ± 13 Pa (40.0 ± 0.1 torr). A

spectrum was taken. The cylinder valve was closed tightly. Then the gas cell and lines were emptied

with vacuum and N2 purging. The cylinders were then removed, the end caps fastened tightly, and

weighed. Each cylinder was then checked for leaks before being put back into its aging temperature

environment.

The fluorocarbon agents were analyzed every 8 weeks, unless the results indicated no changes were occurring. The CF3I samples were analyzed every 4 weeks, unless the results indicated no

changes were occurring. Samples that were not changing (e.g., those at ambient conditions) were

analyzed every 16 or 8 weeks for fluorocarbon and CF3I cylinders, respectively.

7.2.5 Use of the FTIR Spectra to Determine Agent Degradation. I he FTIR spectra of the agents were used as both quantitative and qualitative tools to determine if there was degradation of the agent.

As a quantitative tool, an absorbance band for each agent was selected in the 53.3 Pa spectra and the area under this absorbance band was determined for each 4 or 8 week interval. Chemical analysis by

IR spectroscopy is based on the assumption that the IR spectrum of a compound is sufficiently unique

to identify it. An IR spectrum is obtained by measuring the ratio of the intensity of IR radiation

which passes through the sample, /, to the intensity of the incident radiation, as a
7. AGENT STABILITY UNDER STORAGE 257

Table 3. Agent and metal amounts in storage cylinders for FC-218

Metal Temp.

(°C)

Metal mass

(±0.1 g)

Agent Agent/N,

(± 0.02 Mpa) (g)

Pr<* ceil (± 0.02 Mpa)

Mass

Dciurc

Storage (± 0.1 g)

Mass

after storage (± 0.1 g)

Blank 150 0.85 79.9 4.20 122.2 95.6 N40 150 344.4 0.84 (82,3) 4.20 120.5 107.5

Ti 150 188.3 0.83 (81.3) 4.24 113.8 76.9

C4130 150 304.5 0.76" (74.5) 4.20 113.0 106.2

1625 150 326.9 0.83 71.9 4.24 118.1 108.7

Storage tank was almost empty when this cylinder was filled.

Table 4. Agent and metal amounts in storage cylinders for HFC- 125

Metal

Temp.

(°C)

Metal mass

(±0.1 g)

Agent Agent/N,

Pressure

(± 0.02 Mpa)

Mass

(g)

Pressure

(± 0.02 Mpa)

Mass

before

storage (± 0.1 g)

Mass

after storage (± 0.1 g)

Blank 23 1.28 80.2 4.24 119.7 116.0 N40 23 345.1 1.28 (76.8) 4.24 115.5 109.5

Ti 23 187.5 1.28 (76.8) 4.24 119.2 114.4

€41 30 23 304.4 1.28 (76.8) 4.20 121.2 115.2

1625 23 327.0 1.31 78.6 4.27 121.7 114.7

Blank 150 1.28 (76.8) 4.27 119.8 111.9 N40 150 343.3 1.28 (76.8) 1.28 117.9 94.6

Ti 150 188.0 1.28 (76.8) 4.27 121.0 105.4 C4130 150 304.2 1.28 78.8 4.24 122.3 114.1

1625 150 327.0 1.28 (76.8) 4.27 118.2 109.0
258 7. AGENT STABILITY UNDER STORAGE

Table 5. Agent and metal amounts in storage cylinders for HFC-227ea

Metal

Temp.

(°C)

ivieiai mass

(±0.1 g)

Agent Agent/Nj

Pressure

(± 0.02 Mpa)

Mass

(g)

Pressure

(± 0.02 Mpa)

Mass

before

storage (± 0.1 g)

Masj

after storage (± 0.1 g)

DlaUK Rl^lnl- 0.41 31.8 4.20 77.3 74.6

N40 23 343.7 0.41 (31.8) 4.20 73.7 71.4

Ti 23 187.3 0.41 (31.8) 4.20 76.0 73.2

€41 30 23 304.3 0.41 (31.8) 4.20 72.1 68.9

1625 23 327.0 0.41 30.6 4.20 72.0 67.4

BianK \ Zj 0.41 32.9 4.20 80.7 74.7

N40 125 344.5 0.41 (31.8) 4.20 74.3 63.1

Ti 125 187.7 0.41 (31.8) 4.20 74.9 64.4

C4130 125 303.9 0.41 (31.8) 4.20 74.6 67.4

1625 125 327.2 0.41 (31.8) 4.20 73.1 68.1

DlallK 1 sn 0.41 (31.8) 4.20 76.5 71.7

N40 150 344.6 0.41 (31.8) 4.20 72.9 66.1

Ti 150 187.3 0.41 (31.8) 4.20 78.1 70.3

C4130 150 304.4 0.41 (31.8) 4.20 73.9 66.3

1625 150 327.0 0.43 (33.4) 4.20 58.7 42.9
7. AGENT STABILITY UNDER STORAGE 259

Table 6. Agent and metal amounts in storage cylinders for dry CF3I samples w/o copper

Metal

Temp.

(°C)

Metal mass

(±0.01 g)

Agent Agent/Nj

Pressure (± 0.02 Mpa)Mass

(g)

Pressure

(± 0.02 Mpa)

Mass

before

storage (± 0.1 g)

Mass

after storage (± 0.1 g)

Blank 23 -- 0.45 37.4 4.24 85.4 70.4 N40 23 345.7 0.45 (40.2) 4.24 82.5 71.4

Ti 23 188.3 0.45 (40.2) 4.24 80.1 72.0

C4130 23 304.3 0,45 36,9 4.20 81.4 72.9

1625 23 326.8 0.48 36.8 4.20 79.9 72.4

Blank 100 -- 0.48 36.9 4.20 83.3 35.3 N40 100 343.7 0.48 (42.9) 4.20 80.9 --

Ti 100 187.3 0.48 (42.9) 4.20 78.1 70.2

C4130 100 304.3 0.45 (40.2) 4.20 79.1 54.2

1625 100 327.5 0.45 (40.2) 4.20 78.0 71.5

Blank 150 0.47 40.1 4.20 87.3 33.8 N40 150 345.1 0.47 (42.0) 4.20 80.9 55.8

Ti 150 187.5 0.47 (42.0) 4.20 84.8 30.1

C4130 150 304.5 0.45 (40.2) 4.20 79.1 32.8

1625 150 327.0 0.48 (42.9) 4.20 80.4 66.3
260 7. AGENT STABILITY UNDER STORAGE

Table 7. Agent and metal amounts in storage cylinders for dry CF3I samples with copper

Metal

Temp.

(°C) Cu(CDAl 10)/ metal mass

(±0.1 g)

Agent Agent/Nj

Pressure

(± 0.02 Mpa)

Mass

(g)

Pressure

(± 0.02 Mpa)

Mass

before

storage (± 0.1 g)

Mass

after storage (± 0.1 g)

Blank 23 353.9/-- 0.45 (40.2) 4.24 80.4 72.3 N40 23 354.5/342.6 0.45 (40.2) 4.24 77.7 68.1

Ti 23 354.5/187.1 0.45 (40.2) 4.24 78.7 70.6 C4I30 23 354.6/304.2 0.45 (40.2) 4.20 74.8 68.6

1625 23 354.1/327.0 0.48 (42.9) 4.20 75.2 25.7

Blank 100 355.1/- 0.48 (42.9) 4.20 80.5

Blank 150 352.9/- 0.47 (42.0) 4.20 84.0 71.8 N40 150 352.7/343.5 0.47 (42.0) 4.20 80.2 60.0

Ti 150 352.5/188.2 0.47 (42.0) 4.20 78.5 67.6 C4130 150 354.1/304.2 0.45 (40.2) 4.20 74.8 49.2

1625 150 353.6/327.0 0.48 (42.9) 4.20 77.1 65.3

Table 8. Agent and metal amounts in storage cylinders for moist CF3I samples w/o copper

Metal

Temp.

(°C)

Metal mass

(±0.01 g)

Agent Agent/Nj

Pressure

(± 0.02 Mpa)

Mass

(g)

Pressure (± 0.02 Mpa)

Mass

before

storage (± 0.1 g)

Mass

after storage (± 0.1 g)

Blank 100 0.48 (42.9) 4.20 81.8 73.4

Blank 150 0.47 39.1 4.20 86.0 60.1 N40 150 345.5 0.47 (42.0) 4.20 82.4 63.9

Ti 150 188.1 0.47 (42.0) 4.20 80.6 60.7

C4130 150 304.1 0.45 (40.2) 4.20 77.1 63.7

1625 150 327.0 0.48 (42.9) 4.20 80.7 70.5
7. AGENT STABILITY UNDER STORAGE 261

Table 9. Agent and metal amounts in storage cylinders for moist CF3I samples with copper

Metal

Temp.

(°C) metal mass

(±0.1 g)

Agent Agent/Nj

Pressure (± 0.02 Mpa)

Mass

(g)

Pressure (± 0.02 Mpa)

Mass

before

storage (± 0.1 g)

Mass

after storage (± 0.1 g)

Blank 100 354.2/-- 0.48 (42.9) 4.20 79.1 70.7

Blank 150 354.4/- 0.47 (42.0) 4.20 80.0 54.9 N40 150 353.7/344.4 0.47 (42.0) 4.20 76.6 64.9

Ti 150 352.6/187.4 0.47 (42.0) 4.20 75.4 64.5

C4130 150 353.5/304.3 0.45 (40.2) 4.20 73.0 59.1

1625 150 352.7/319.1 0.45 (40.2) 4.20 73.9 61.9

function of frequency (v). Concentrations of individual components are quantified by applications of

Beer's law

q-m]c,. ^(v) = -iog«^. (i:

where subscript r indicates known values obtained from a reference sample. IR spectroscopy is extremely versatile in the sense that almost all compounds, with the notable exception of single atoms

and homonuclear diatomics, are IR active. In principle, IR analysis makes it possible to monitor the

degradation of each candidate and identify the corresponding products from a simple before and after comparison of the spectra. In practice, IR analysis can suffer from overlapping peaks in a spectrum

similar to matrix effects with other analytical methods. In addition, quantitative determination of

degradation products require either pre-existing spectra for comparison (libraries of spectra for

common materials are commercially available) or available reference samples. The criteria for

choosing an agent absorbance band for integration were the following: an absorbance band representing the weakest bond in the molecule; a small band with maximum absorbance less than 0.6 so that if a reduction in area occurred as a result of degradation the change would be detected; assumed to obey

Beer's law at lower concentrations if the absorbance is less than 0.6 at higher concentrations; and well

resolved with respect to the baseline noise.

FC-218 is a symmetrical molecule with 2 C-C and 8 C-F bonds. The well resolved peak at 731 cm ' was chosen for integration. Both HFC- 125 and HFC-227ea contain C-C, C-H and C-F

bonds. The C-H stretch absorbance band near 3000 cm"' for each of these agents was chosen. All of

these chosen agent absorbance bands satisfied the above criteria. The spectra for these agents are shown in Figures 3-5 v/ith the absorbance band that was integrated noted.

For CF3I, the ideal absorbance band to have integrated would have been that for the C-I stretch. The fundamental vibrational frequency for the C-I stretch from the literature (McGee, 1952) is 539 cm"'. However, the current detector is only sensitive from 4,000 to 600 wavenurnbers. As shown

in the lower spectrum in Figure 6, a small, but well resolved peak appears at 2255 cm"'. This peak
262 7. AGENT STABILITY UNDER STORAGE

7

8

6 9®

60

2 3 4

10

11 12

1 . Agent Storage Cylinder

2. Cylinder Valve

3. Toggle Valve

4. Micro-Metering Valve

5. Toggle Valve

6. Nj Purge Line

7. Variable Path Length Gas Cell @ 10 m Path Length

8. Pressure Gauge

9. Gas Cell Pressure Gauge Valve

10. Gas Cell Filling Valve

1 1 . Main Vacuum Valve

1 2. Two-Stage Vacuum Pump

Figure 2. Diagram of the equipment for the FTIR spectral analyses.
AGENT STABILITY UNDER STORAGE 263

A

b 1.0-

3000 2500 2000 1500 1000

Wavenumbers (cm-l)

Figure 3. Agent absorbance band in the 53.3 Pa IR spectrum of FC-218 that was integrated.
264 7. AGENT STABILITY UNDER STORAGE

3500 3000 2500 2000 1500 1000

Wavenumbers (cm-1)

Figure 4. Agent absortance band in the 53.3 Pa IR spectrum and CO2 absorbance band in the

5330 Pa IR spectrum of HFC- 125 that were integrated.
AGENT STABILITY UNDER STORAGE 265

A

b

s

o

rb

a

n

c

e

1.0

0.8

0.6-

0.4

0.2-

0.0^

Peak for integration

53.3 Pa spectrum

J t 3000 2500 2000 1500 1000

Wavenumbers (cm-l)

Figure 5. Agent absorbance band in the 53.3 Pa IR spectrum of HFC-227 that was integrated.
266 7. AGENT STABILITY UNDER STORAGE

corresponds to a combination mode of two larger peaks that correspond to the fundamental stretching

frequency of C-F bonds and are labelled peaks 1 and 2 in Figure 6. The small peak thus results from

both the stretching and bending of symmetrical C-F bonds. If only a small fraction of CF3I molecules

in the cylinder are degrading, then a reduction in area under this small band will be more easily

detected because the low intensity band is more likely to follow Beer's law. Therefore, we used the

2255 cm ' absorbance band.

As a qualitative tool, higher concentration spectra were used. If degradation of the agent were to occur, then the degradation products should begin to form and impurity bands in the aged IR spectrum

from each cylinder should appear. Since the amount of degradation products might be small in comparison to the amount of agent, aged spectra at 5330 Pa (or 100 times as dense as the quantitative

spectra) were obtained for each cylinder at each testing interval. The baselines of these higher

concentration, "qualitative spectra" were examined closely to determine if new absorbance bands were

appearing.

In addition to any new bands appearing, bands from impurities already present in the agents could

be determined from the higher concentration spectrum. Impurity bands from COj were found in all of

the initial spectra at 5330 Pa for both the HFC- 125 and CF3I samples. COj might have been present

in FC-218 and HFC-227ea, however, this could not be determined because both agents have strongly

absorbing bands in the 670 and 2360 wavenumber regions which obscures the absorption spectra of

CO2. Also detected in the initial spectra for the CF3I cylinders was CF3H. Since different lots of some agents were used. Table 10 shows the initial areas for the impurity absorbance bands for each lot of agent determined from the initial 5330 Pa FTIR spectrum.

The only agent to show a new absorbance band appearing in the spectra at elevated temperatures was CF3I. This new absorbance band will be discussed in 7.3.5.2. Figures 4 and 6 also show the absorbance bands in the higher concentration spectra that were

integrated.

As the study progressed, it became obvious that the 5330 Pa spectra for samples stored at 150 °C

were changing, especially for the CF3I samples. A software program (Mattson Instruments, Inc., 1992)

generates a correlation coefficient comparison (Mendenhall et ai, 1992) report of sample data (aged

spectrum) and reference data (initial spectrum). This spectral comparison takes into account changes

that are occurring in a spectrum as a result of changes in existing peaks and the appearance of new

absorbance peaks. The more a spectral comparison decreases from 1.0(X) (the value obtained when a spectrum is compared to itself) represents the magnitude of the changes taking place. Thus the value

of the spectral comparison for an aged spectrum compared to its initial spectrum allowed a quantitative

value to be assigned to the changing spectra as a function of temperatures and conditions of aging.

7,2.6 Determination of Uncertainty. A Type A evaluation of the standard uncertainty (Taylor,

1995) was used in this study. The samples kept at ambient conditions (23 °C) were not changing

with time. This was evident by seeing no pattern in the measured integrated peak areas as a function

of time. For each agent and metal combination at 23 °C, the peak area measurements obtained over

the study period were averaged and a standard deviation was determined. As a worst-case scenario,
AGENT STABILITY UNDER STORAGE 267

Figure 6. Agent absorbance band in the 53.3 Pa IR spectrum and impurity absorbance bands in

the 5330 Pa IR spectrum of CF3I that were integrated.
268 7. AGENT STABILITY UNDER STORAGE

Table 10. Impurity levels in the lots of agents at the start of the study

Agent Lot numbers

Integrated area

Impurity compound

CO2 CF3H

FC-218 L- 12677 ND

HFC- 125 835 0.01 ND

HFC-227ea 92-002-356 ND ND

HFC-227ea 93-200-278 ND ND

CF3I 224940901" 0.09 0.01

CF3I 226941712"^ 0.18 0.04

CF3I 226941891' 0.01 ND

Not detected

Storage tank was roughly half full when analyzed

Storage tank was full when analyzed

Storage tank was nearly empty when analyzed.

the largest standard deviation for a given agent and condition was used for the uncertainty. Agents

FC-218 and CF3I in the moist condition were exceptions. Neither of these samples had ambient

cylinders specified in the test matrix. The uncertainty reported for FC-218 was made from the 150 °C

peak area data as stated above, based on the fact that no changes in the 5330 Pa spectra were

occurring. The uncertainty for the moist CF3I samples was assigned the same as that for the dry

condition. Table 1 1 lists the wavenumber ranges over which integration was performed for each agent

and impurity band along with the uncertainty in the area measurements and the number of observations (N). Note the relatively high uncertainties for the FC-218 and HFC-125. These were the two

strongest absorbance bands in the 53.3 Pa spectra of these agents.

The uncertainty in the spectral comparison values was determined in the same manner as described above for the change in area. For each agent and metal combination at 23 °C, each aged

spectrum was compared to its respective initial spectrum. These values were averaged and a standard

deviation was determined. Once again the largest standard deviation was chosen as the standard

uncertainty. The spectral comparisons and uncertainties are reported below in the tables of agent

absorbance band areas for each agent and metal combination at 23 °C, along with the number of

observations (N).

In some cases the impurity bands for COj and CF3H in the initial spectra were very small. To

ascertain whether these bands could be measured reliably, a measure of the uncertainty in the detection

of these peaks was determined. Examination of the baseline of a typical FTIR spectrum in the

expanded axis mode reveals a baseline that is sinusoidal in nature. One of these very small "sinusoidal" peaks where no absorbance was occurring as a result of agent, and near the impurity absorbance

band of interest was integrated. One of the smallest areas for absorbance reported (0.01) was three

times larger than one of the largest baseline absorbance areas. The changes in the agent absorbance

bands and impurity bands are described below for the individual agents.
7. AGENT STABILITY UNDER STORAGE 269

Table 1 1 . Wavenumber ranges for integrated areas of agent and extraneous peaks

Agent

^^avpnnmhpr rj^ncrp for integrated area (cm ')

Twiic^i TriPi^rtJ^i idiiiiy ntv in111 rnpficiirf^H iiicdoUiCLi area/N

FC-218 710 - 750 0.15/6

HFC- 125 2963 - 3038 (Agent)

667 - 671 (CO2)

0.042/4

0.0023/4

HFC-227ea 2966 - 3006 (Agent) 0.0095/4

CF3I

2219 - 2274 (Agent)

667 - 671 (CO2)

698 - 701 (CF3H)

944 - 953 (F-alkene)

0.011/9

0.0048/9

0.0011/9

0.011/9

Because of the large number of samples, the time required to obtain all materials and equipment,

and the time required for initial analyses, the introduction of cylinders into the test matrix took several

months. Thus, not all metals in a given agent or test condition of CF3I received the same time of

aging. In the tables of integrated data area, the last column to have data reported was the extent of

aging. Because the agents, especially CF3I, were in short supply at the beginning of this study and

other studies had to get underway, it was not possible to retain a particular lot of any agent. When a

series of cylinders was filled, the agent storage tank was passed to other researchers. By the time that agent was needed again in this study, all of the lot was sometimes used. Therefore, getting all of the

cylinders started required the use of different lot numbers of agents. These factors added to the

uncertainty of the data analyses. Operator experience and care in performing the analyses also affected

the spectral comparisons. For this reason, an attempt to establish Type B evaluations of standard

uncertainty (Taylor, 1995) were not done.

7.3 Results

7.3.1 FC-218. At the onset of this investigation, FC-218 was considered to be the most stable of the

four agents. Therefore, the test matrix (Table 1) specified this agent to be tested only at 150 °C.

Figures 7-11 show the overlaid spectra of the initial and final aged spectra at 5330 Pa for each metcil and blank. Visual examination of these spectra indicated no changes occurred. Table 12 shows the

integrated area for the agent absorbance band with time and the spectral comparison for the samples.

The initial area value for C4130 is an outlier known to result from improper filling of the IR gas cell. Figure 12 shows the integrated areas in Table 12 graphically represented. The areas tended to drift upward over the 48 week period. This was the only agent for which the integrated area for the

absorbance band of an agent showed the slight upward drift. The data for this agent had more scatter than any of the other agents. However, the spectral comparison values in the last column of Table 12

indicate the spectra are unchanged.

7.3.2 HFC-125. Because of the weaker C-H bond in this molecule, the test matrix (Table 1) specified

two test temperatures. A typical initial and aged spectra at 5330 Pa and at 23 and 150 °C for the
270 7. AGENT STABILITY UNDER STORAGE

Table 12. Integrated peak areas for a C-F absorbance band in the 53.3 Pa spectra for FC-218 at 150 °C and the spectral comparisons

Integrated area from 710 to 750 cm"' (± 0.15) Spectral

companson

(± 0.00070)^

Week number

Metal 0 8 16 24 32 40 48

Blank 6.2 6.3 6.4 6.4 6.5 6.5 0.996

N40 6.6 6.8 NM 6.9 6.8 7.0 7.0 0.996

Ti 6.6 6.7 NM 6.8 6.7 6.8 6.7 0.996

C4130 5.7'-' 6.5 6.4 6.5 6.5 6.5 NM 0.999

1625 6.5 6.5 6.7 6.6 6.6 6.7 6.9 0.998

Number of observations was 5 Not measured

Outlier from improper gas cell filling.

blank for HFC- 125 are shown in Figures 13 and 14, respectively. Visual examination of the initial and final spectra for the cylinder at 23 °C (Figure 13) showed no change in the absorbance and no new peaks. Visual examination of the initial and final spectra for the cylinder tested at 150 °C

(Figure 14) did show a small increase in the absorbance band for COj in the 670 cm"' region. The

samples containing each of the four metals gave similar results. Their spectra are shown in Appendix A for reference.

Table 13 lists the integrated areas under the C-H stretch absorbance band with time and

temperature for each metal and temperature. Figure 15 is a graphic representation of this data. In all cases the peak areas randomly varied from 2.7 to 2.9. The graphs showed some scatter of the areas, but no upward or downward drift with time. All of the areas reported for the aged samples were

within three standard deviations of its initial area. At the 99 % confidence interval, the spectral

comparisons did not change significantly either.
7. AGENT STABILITY UNDER STORAGE 271

Figure 7. Initial (lower) and 48 week (upper) 5330 Pa spectra for the blank for FC-218 at 150 °C.
272 7. AGENT STABILITY UNDER STORAGE

3500 3000 2500 2000 1500 1000

Wavenumbers (cm-1

)

Figure 8. Initial (lower) and 48 week (upper) 5330 Pa spectra for nitronic 40 in FC-218 at 150 °C.
7. AGENT STABILITY UNDER STORAGE 273

3500 3000 2500 2000 1500 1000

Wavenumbers (cm-1

)

Figure 9. Initial (lower) and 48 week (upper) 5330 Pa spectra for Ti- 15-3-3-3 in FC-218 at 150 °C.
274 7. AGENT STABILITY UNDER STORAGE

Figure 10. Initial (lower) and 40 week (upper) 5330 Pa spectra for C4130 in FC-218 at 150 °C.
. AGENT STABILITY UNDER STORAGE

Figure 11. Initial (lower) and 48 week (upper) 5330 Pa spectra for 1625 in FC-218 at 150 °C.
276 7. AGENT STABILITY UNDER STORAGE

I I I I I

I 0 8 16 24 32 40 48

Time (weeks)

Figure 12. Integrated peak areas for all samples of FC-218 at 150 °C plotted as a function of

time.
AGENT STABILITY UNDER STORAGE 277

Figure 13. Initial (lower) and 48 week (upper) 5330 Pa spectra for the blank for HFC-125 at 23 °C.
278 7. AGENT STABILITY UNDER STORAGE

b

3000 2500 2000 1500 1000

Wavenumbers (cm-1

)

Figure 14. Initial (lower) and 48 week (upper) 5330 Pa spectra for the blank in HFC- 125 at 150 °C.
7. AGENT STABILITY UNDER STORAGE 279

Table 13. Integrated peak areas for the C-H stretch absorbance band in the 53.3 Pa spectra for

HFC- 125 at 23 and 150 °C and the spectral comparisons

Integrated Area from 2963 to 3038 cm ' (± 0.042)

Spectral

Comparison

(± 0.00095r

Week Number

Metal

Temp.

(°C) 0 8 16 24 32 40 48

Blank 23 2.8 NM" NM 2.7 NM 2.8 2.7 0.996

N40 23 2.7 NM NM 2.7 NM 2.7 2.7 0.997

Ti 23 2.8 NM NM 2.8 NM 2.8 2.8 0.996

C4130 23 2.8 2.8 2.8 2.9 2.8 NM NM 0.998

1625 23 2.8 NM 2.8 2.8 2.8 2.8 2.9 0.998

Blank 150 2.8 2.8 NM 2.8 2.7 2.8 2.8 0.995

N40 150 2.8 2.8 NM 2.8 2.7 2.8 2.8 0.995

Ti 150 2.8 2.8 NM 2.8 2.7 2.8 2.8 0.996

C4130 150 2.8 2.9 2.8 2.8 2.9 2.9 NM 0.998

1625 150 2.8 2.9 2.8 2.8 2.8 2.8 2.9 0.998

Number of observations was 3

Not measured
280 7. AGENT STABILITY UNDER STORAGE

Figure 15. Integrated peak areas for all samples of HFC-125 at 23 °C and 150 °C plotted as a

function of time.
7. AGENT STABILITY UNDER STORAGE 281

In the 5330 Pa spectrum of HFC- 125 a small amount of COj (<100 \iL/L) was detected in all of

the initial analyses. Table 14 shows the integrated areas under the absorbance band at 670 cm"' for

COj for all samples and temperatures. Figure 16 is the graphic representation of this data. At 23 °C

no changes in the amount of COj were seen. At 150 °C the COj content in the samples increased

anywhere from three and a half to a twelve times after 32 weeks. Beyond 32 weeks the COj

concentration remained constant. These data confirm the visual observations of an increase in the COj

content with time and temperature. The COj could be from outgassing or from agent decomposition

in air, which would also yield HjO, HF, HI, or other compounds. These would be measurable if they

didn't sorb to or react with the cylinder walls or metal coupons. There were no hints from the spectra

of the above compounds forming, so the COj increase was probably from outgassing. It should be

noted that the increase in this peak is not sufficient to change the spectral comparisons.

There are no other apparent changes in the initial and final spectra of any of the cylinders at 23

or 150 °C.

7.3.3 HFC-227ea, HFC-227ea is a three carbon chain molecule with C-C and C-F bonds and a C-H

bond. Of the three fluorocarbon agents this agent was thought to be the most likely to undergo

degradation at elevated temperatures. The test matrix (Table 1) specifies that this agent was to be

tested at 23, 125, and 150 °C to see if thermal degradation would result. Figures 17-19 show the

initial and final spectra at 5330 Pa for the blank at the three temperatures, respectively. Visual

examination of the baselines and comparison of the spectra at 5330 Pa show no additional absorbance

bands appearing in any of the spectra at any temperature. This is true for all of the metal samples at

all three temperatures. The initial and final spectra for all of these samples are shown in Appendix B

for reference.

Table 15 shows the integrated peak areas for the C-H absorbance band and the spectral comparisons for each metal and blank at each temperature. The integrated area for 1625 at 150 °C is half that for the other cylinders. This is a result of incomplete filling at the outset. Figure 20 is the graphical

representation of the data in Table 15. As with the HFC- 125, there was little if any change in the

peak areas with time or temperature. Also the data show no tendencies to drift either up or down. All

of the reported aged data areas fall within two standard deviations of their respective means. The

spectral comparisons are all 0.997 or greater. The blank, Nitronic 40, and Ti at 125 °C all had

spectral comparisons higher than their respective 23 °C controls. The 125 °C samples were prepared

and analyzed later in the study. By this time the analytical technique had become more efficient and

quite routine, leading to better spectral comparisons.

7.3.4 Summary of Fluorocarbon Area Data. The above graphs of the agent peak areas suggested

that no changes were occurring in the agents. Table 16 is a summary of the area data from Tables 12,

13 and 15. This table shows only the initial and final peak areas along with the number of weeks

each sample was aged. The last column shows whether the final data areas changed significantly at the 99 % confidence interval. None of the final peak areas were significantly less for any agent. All

of the final areas for the FC-218 were actually higher than their original peak areas. For a peak to increase, the change has to be in the analysis. As such, the "increase" may set a level of uncertainty.

Table 17 summarizes the spectral comparisons for all agents tested at 150 °C. Since there were

no ambient controls for FC-218, the spectral comparisons listed for this agent are simply the final spectra compared to the initial spectra. These spectral comparisons suggest that little change in any of

the samples occurred. For the other two agents, the spectral comparisons are amazingly similar. With

the exception of the blank and Nitronic 40 in HFC-125, which showed a slight decrease, but not

statistically less at the 99 % confidence interval, there were no changes in the spectral comparisons for
282 7. AGENT STABILITY UNDER STORAGE

Table 14. Integrated peak areas for an absorbance band of COj in the 5330 Pa spectra of

HFC- 125 at 23 and 150 °C

Integrated Area from 667 to 671 cm"' (± 0.0023)

Week Number

Metal Temp.

(°C) 0 8 16 24 32 40 48

Blank 23 0.01 NM- NM 0.01 NM 0.01 0.01 N40 23 0.01 NM NM 0.01 NM 0.01 0.01

Ti 23 0.01 NM NM 0.01 NM 0.01 0.01 C4130 23 0,02 0.01 0.01 0.01 0.01 NM NM

1625 23 0.0! NM 0.01 0.01 0.01 0.01 0.01

Blank 150 0.01 0.06 NM 0.08 0.09 0.09 0.09 N40 150 0.01 0.08 NM 0.10 0.12 0.12 0.12

Ti 150 0.01 0.06 NM 0.07 0.08 0.08 0.08 C4130 150 0.01 0.06 0.08 0.09 0.09 0.10 NM

1625 150 0.02 0.04 0.06 0.07 0.07 0.08 0.09

Not measured

samples aged at elevated temperatures. The summary data presented in Tables 16 and 17 indicate the

fluorocarbon agents are stable at elevated temperatures for as long as 48 weeks.

7.3.5 lodotrifluoromethane (CF3I). This agent contains C-F bonds and a C-I bond. The C-I bond in

this molecule is very weak, about 223 kJ/mole (Felder, 1992). This suggests that degradation,

particularly at elevated temperatures, may be a factor. Additionally, the presence of moisture in the

storage vessels at elevated temperatures was investigated, since water was hypothesized to be an

accelerator of degradation. Finally, the possibility that the addition of copper to the cylinders might

inhibit degradation was investigated. Therefore, this agent was tested not only at three temperatures,

but also at four different conditions of copper and moisture content (Table 1). The initial spectroscopic analyses of the three lots of CF3I that were tested indicated that they

contained COj and CF^H. These could have been left from the synthesis or they may be an indication

that some degradation of the agent has already occurred. The amount of these impurities depended in large part to the amount of agent still remaining in the storage container when the test cylinders were

filled. Since the boiling points of both COj and CF3H are lower than that of CF3I, the vapor that is removed from a storage tank is richer in these impurities. Table 10 shows the starting impurity levels of CO2 and CF3H for each of the lots used. Notice in the footnotes the amount of agent remaining in the storage tanks when they were analyzed. Lot number 226941712 was analyzed before any of the

agent had been used and the CO2 and CF3H were the highest. Lot number 226941891 was nearly

empty when it was analyzed and it contained little CO2 and undetectable CF3H (probably present, but

below the detection limit of the method.) In the forthcoming graphical presentations of the data for

these impurities, the initial areas vary up and down the y-axis, generally, as a function of which lot of

CF3I was used and when the cylinder was filled (higher areas as a result of fuller tanks.)
Figure 16. Integrated peak areas for the COj absorbance band at 670 cm"' for all samples of

HFC-125 at 23 °C and 150 °C plotted as a function of time.
284 7. AGENT STABILITY UNDER STORAGE

Figure 17. Initial (lower) and 48 week (upper) 5330 Pa spectra of the blank in HFC-227ea at 23 °C.
AGENT STABILITY UNDER STORAGE

3500 3000 2500 2000 1500 1000

Wavenumbers (cm-1

)

Figure 18. Initial (lower) and 40 week (upper) 5330 Pa spectra for the blank in HFC-227ea at 125 °C.
286 7. AGENT STABILITY UNDER STORAGE

Figure 19. Initial (lower) and 48 week (upper) 5330 Pa spectra of the blank in HFC-227ea at 150 °C.
7. AGENT STABILITY UNDER STORAGE 287

Table 15. Integrated peak areas for the C-H stretch absorbance band in the 53.3 Pa spectra of

HFC-227ea at 23, 125, and 150 °C and the spectral comparisons

Integrated area from 2966 to 3006 cm"' (± 0.0095)

Spectral

comparison

(± 0.00044)^

Week number

Metal

Temp.

( C,) U

oO

1 AId z4 32 4U A Q 4o

Blank 23 0.22 NM NM 0.21 NM 0.20 0.21 0.997

N40 23 0.21 NM NM 0.20 NM 0.19 0.20 0 QQ7

Ti 23 0.21 NM NM 0.20 NM 0.20 0.20 0 007

C4130 23 0.19 0.19 0.19 0.18 0.20 NM NM 0 000

1625 23 0.18 NM 0.19 0.18 0.19 0.19 0.19 0 000

Blank 125 0.21 0.22 0.21 0.21 0.23 0.20 NM 0.999

N40 125 0.19 0.20 0.20 0.20 0.21 0.18 X TX /f NM 0xj.yyy 000

Ti 125 0.19 0.19 0.19 0.19 0.20 0.18 XTX /r NM 0 000

C4130 125 0.20 0.20 0.19 0.20 0.21 0.20 XTX JT NM 0 009

1625 125 0.20 0.20 0.20 0.20 0.20 0.20 NM 0 000

Blank 150 0.20 0.20 NM 0.20 0.19 0.19 0.19 0.997

N40 150 0.20 0.20 NM 0.21 0.18 0.20 0.19 0.997

Ti 150 0.23 0.23 NM 0.23 0.21 0.23 0.22 0.997

C4130 150 0.20 0.19 0.19 0.19 0.21 0.20 NM 0.999

1625 150 0.10^ 0.11 0.12 0.10 0.11 0.10 0.11 0.999

Number of observations was three

Not measured

Improper filling of cylinder at outset of test.
288 7. AGENT STABILITY UNDER STORAGE

o

OJ

I O

Li_ X

Co

CD

(DU

Co

-Q

i_O

m<

c

o

0) <

o

Q_

"D

0)

0.25

0.20

0.15 h

0.10

0

0.25 r

0.20

0.15

0.10

0

0.25 r

0.20 -

0.15

0.10

All samples at 150 °C

8 16 24 32 40 48 56

o Blank

V N40

Ti A C4130

o 1625

All samples at 125 °C

8 16 24 32 40 48 56

All samples at 23 °C

0 8 16 24 32 40 48 56

Time (weeks)

Figure 20. Integrated peak areas for all samples of HFC-227ea at 23 °C, 125 °C, and 150 °C

plotted as a function of time.
7. AGENT STABILITY UNDER STORAGE 289

Table 16. Summary of the initial and final peak areas for the fluorocarbon agents at all temperatures

Number of weeks Temp (°C) Initial Final

area ± (0.15)

Significant @ 3x

uncertainty

48

48

48

40

48

150

150

150

150

150

6.2

6.6

6.6

5.7"

6.5

6.5

7.0

6.7

6.5

6.9

area ± (0.042)

48

48

48

32

48

23

23

23

23

23

2.8

2.7

2.8

2.8

2.8

2.7

2.7

2.8

2.8

2.9

no

no

no

no

no

48

48

48

40

48

150

150

150

150

150

2.8

2.8

2.8

2.8

2.8

2.8

2.8

2.8

2.9

2.9

48

48

40

32

48

40

40

40

40

40

48

48

48

40

48

23

23

23

23

23

125

125

125

125

125

150

150

150

150

150

area ± (0.0095)

0.22

0.21

0.21

0.19

0.18

0.21

0,19

0.19

0.20

0.20

0.20

0.20

0.23

0.20

0,10"

0.21

0.20

0.20

0.20

0.19

0.20

0.18

0.18

0.20

0.20

0.19

0.19

0.22

0.20

0.11

no

no

no

no

Improperly filled gas cell during analysis Cylinder did not till completely at ouLset of test.
290 7. AGENT STABILITY UNDER STORAGE

Table 17. Spectral comparisons of the initial and aged fluorocarbon agents

Agent FC-218 HFC- 125 (± 0.00095) HFC-227ea (± 0.00044)

Temp. (°C) 150 23 150 23 125 150

Metal

Weeks

tested

Blank 48 0.996 0.996 0.995 0.997 0.997 0.997

N40 48 0.996 0.997 0.995 0.997 0.997 0.997

Ti 48 0.996 0.996 0.996 0.997 0.997 0.997

C4130 40 0.999 0.998 0.998 0.999 0.999 0.999

1625 48 0.998 0.998 0.998 0.999 0.999 0.999

As mentioned earlier, an absorbance band in the 5330 Pa spectra of CF3I tested at 100 and

150 °C began to appear around 950 cm ' after 4 weeks of aging. This absorbance was looked for in reference spectra of CjF^, COFj, and CH3F, but does not appear in any of these compounds. This

absorbance band is in the region of C=C bond stretch frequencies. The absorbance band has not been

definitely assigned, but may correspond to the presence of ethylene or a fluorinated alkene. Since the

corresponding C-H stretch band for ethylene is absent (or below the detection limit at this concentration), it is suspected that the absorbance band might correspond to a fluorinated alkene. Reference

spectra for double bonded fluorinated compounds such as C2F4, C2H2F2 (the 1,1 compound), and C2H3F

did not have an absorbance band in this region of the spectrum. Not all of the spectra for the

fluorinated alkenes are present in our library, so positive identification still remains unresolved. This

fluorinated alkene peak and the increase in the impurity peaks are shown in Figure 21. The lower

spectrum is the initial, 5330 Pa spectrum of a moist sample without copper and containing phosphatetreated C4130. This is one of the samples that initially had no detectable CF3H or COj. The upper

spectrum is that of the corresponding aged sample after 20 weeks at 150 °C. The small peak at 670 cm"' is an absorbance band of CO2, the peaks at 700, 1400, and 3000 cm"' are from CF3H, and the

absorbance band at 950 cm ' for the fluorinated alkene. This spectrum illustrates quite well the

changes that occurred in the CF3I at elevated temperatures.

The spectrum in Figure 22 shows a spectrum that results from the subtraction of the spectra

shown in Figure 21. This subtracted spectrum not only shows the peaks mentioned in Figure 21, but

also shows the characteristic peak at 2150 cm ' for carbon monoxide and the other characteristic peak

at 2360 cm ' for CO2. The peaks in the 1200 cm"' region correspond to agent peaks that did not completely subtract out because of their high intensity.

The following data for CF3I are presented for the dry and moist conditions and with or without

the presence of copper. Changes in peak areas for agent, CO2, CF3H, and fluorinated alkene are the

order of presentation for each test condition.

7.3.5.1 Changes in the Agent Peak Area of CF3I in the 53.3 Pa Spectra for the Different

Conditions. Tables 18-21 show the integrated areas for the absorbance band in the agent and the

spectral comparisons. The data presented in these tables are an average of the three replicate gas cell

fillings. Figures 23-26 show the respective graphical representations of the data. Figure 23 shows the
7. AGENT STABILITY UNDER STORAGE 291

data for the dry condition without copper. At 23 °C, with the exception of the outlier for Nitronic 40

(a result of improper filling of the gas cell), the data are closely bunched and remain relatively constant. The same was true for the data at 100 °C. The data for 150 °C were more spread out but

generally seem to equilibrate after a slight decrease.

Figure 24 shows the peak area data for the dry condition with copper. At 23 °C, the areas tended

to drift downward slightly, then equilibrated. Overall, there is not much change in the data for any of

the metals over the course of the study. The blank at 100 °C was scattered but also unchanged. The

areas at 150 °C were drifting downward consistently. This suggests that copper is not inhibiting the

degradation of CF3I at the elevated temperature.

Figure 25 shows the peak area data for the moist condition without copper. There were no 23 °C

controls for this condition. The areas for the blank at 100 °C remained fairly constant over the study.

The peak areas at 150 °C drifted downward slightly, but reached a minimum at about 32 weeks and

started to equilibrate. All of the HjO may have been consumed by this time. The areas at 150 °C

were more scattered than the previous areas at 150 °C for the dry condition.

Figure 26 shows the peak area data for the moist condition with copper. Once again there were

no controls at 23 °C for this condition. The areas were changing similarly to the moist condition

without copper shown in Figure 25. The above graphs suggest that small decreases in agent peak area were occurring, especially at 150 °C and in the moist condition more so than the dry condition. It appears that copper is not

inhibiting the degradation, HjO causes accelerated degradation, and both combined cause about the same amount, all at 150 °C. The other metals seem to have no effect. Since the samples at 100 °C

showed no effect, it may be that at < 100 °C there is little loss of CF3I.
292 7. AGENT STABILITY UNDER STORAGE

20 week spectrum at 5330 Pa

Figure 21. Illustration of the changes taking place in CF3I samples at elevated temperatures.
'. AGENT STABILITY UNDER STORAGE 293

0.6-

0.5-

A

b

s

o

rb

a

n

c

e

0.4

0.3

0.2

0.1-

0.0

CF3H

C02

CO

Agent

CF3H

C02

CF3IF-alkene \

3500 3000 2500 2000 1500 1000

Wavenumbers (cm-1)

Figure 22. The subtracted spectrum at 5330 Pa for CF3I in the presence of C4130 alloy and

moisture at 150 °C after 20 weeks.
294 7. AGENT STABILITY UNDER STORAGE

Table 18. Integrated peak areas for an absorbance band of CF3I in the 53.3 Pa spectra tested in the dry condition without copper and the spectral comparisons

Integrated area from 2219 to 2274 cm"' (± 0.011)

Week number

Metal

Temp.

(°C) 0 4 8 16 20 24 28 32 36

Blank 23 0.25 0.25 0.24 0.24 0.25 0.22 0.24 NM

N40 23 0.19 0.27 0.25 0.25 0.25 0.24 NM 0.26 NM

Ti 23 0.26 0.25 0.25 0.25 0.23 0.23 NM 0.24 NM

C4130 23 0.24 0.25 0.26 0.24 0.22 0.24 NM 0.23 NM

1625 23 0.27 0.26 0.24 0.24 0.24 0.24 NM 0.24 0.24

Blank 100 0.25 0.24 0.25 0.25 0.24 0.25 0.24 0.25 0.21

N40 100 0.24 0.25 0.25 0.24 0.24 0.24 0.23 0.25 b

Ti 100 0.25 0.25 0.24 0.24 0.24 0.23 0.24 0.25 0.24

C4130 100 0.25 0.25 0.24 0.25 0.24 0.23 NM 0.25 0.24

1625 100 0.25 0.26 0.25 0.25 0.25 0.22 NM 0.23 0.24

Blank 150 0.28 0.28 0.28 0.27 0.26 0.25 0.25 0.25 0.25

N40 150 0.27 0.26 0.27 0.25 0.25 0.24 0.24 0.23 0.24

Ti 150 0.28 0.27 0.27 0.26 0.25 0.25 0.26 0.25 0.25

C4130 150 0.29= 0.24 0.24 0.24 0.26 0.21 NM 0.21 0.23

1625 150 0.25 0.25 0.24 0.23 0.22 0.23 NM 0.22 0.22
7. AGENT STABILITY UNDER STORAGE 295

Table 18 (cont.)- Integrated peak areas for an absorbance band of CF3I in the 53.3 Pa spectra

tested in the dry condition without copper and the spectral comparisons

Integrated area from 2219 to 2274

(± 0.011)

cm

Spectral

comparison

(0.0018)'*

Week number

Metal

Temp.

(°C) 40 44 48 52

Blank 23 0.23 NM 0.23 NM 0.996

N40 23 0.25 NM 0.24 NM 0 99R

Ti 23 0.24 NM 0.23 NM

C4130 23 0.25 NM NM NM 0 QQQ

1625 23 0.25 0.24 NM NM

Blank 100 0.25 0.25 NM NM 0.999

N40 100 __b __b __b __b 0\j*yy 9971

Ti 100 0.24 0.24 NM NM 0 998

C4130 100 0.24 NM NM NM 0 999

1625 100 0.24 NM NM NM 0 999

Blank 150 0.25 0.25 0.24 0.25 0.991

N40 150 0.24 0.23 0.23 0.25 0.995

Ti 150 0.26 0.25 0.25 0.25 0.996

C4130 150 0.23 NM NM NM 0.994

1625 150 0.22 0.22 NM NM 0.988

Not measured

Cylinder emptied after 32 weeks as a result of leak

Suspected to be an outlier because of improper gas cell filling

Number of observations was eight.
296 7. AGENT STABILITY UNDER STORAGE

0.3

0.2

0.1

(L—o

A-" VX~'^"~'"~--v---^

A

1 1

All samples at 150 ''C

1 1 1 I* 1 1 1 1 1 1 1 0 4 8 12 16 20 24 28 32 36 40 44 48 52

0.3 r

All samples at 100 °C

J I I I I L J I L 0 4 8 12 16 20 24 28 32 36 40 44 48 52

0.3 r

0.2

0.

(>^^^ —

V

0 Blank

V N40

Ti A C4130

1

All samples at 23 °C

1 1 1 1

O

1

1625

0 4 8 12 16 20 24 28 32 36 40

Time (weeks)

44 48 52

Figure 23. Integrated agent peak areas for all samples of CF3I tested in the dry condition without copper at 23 °C, 100 °C, and 150 °C plotted as a function of time.
7. AGENT STABILITY UNDER STORAGE 297

Table 19. Integrated peak areas for an absorbance band of CF3I in the 53.3 Pa spectra tested in the dry condition with copper and the spectral comparisons

Integrated area from 2219 to 2274 cm"' (±0.011)

Week number

Metal

Temp.

CO 0 4 8 16 20 24 28 32 36

CDA 110

Blank 23 0.26 U.ZJ 0.24 0.2!) A0.23 A ^ /I 0.24 JNM U.24 NM

CDA 110

N40 23 0.26 0.25 0.24 0.23 0.23 0.24 NM 0.24 NM

LUA i lU

Ti 23 0.27 0.26 0.25 0.24 0.24 0.24 NM 0.24 NM

CDA 110

C4130 23 0.23 0.25 0.24 0.23 0.23 0.24 NM 0.21 NM

CDA 110

1625 23 0.26 0.25 0.25 0.23 0.24 0.23 NM 0.24 0.24

CDA 110

Blank 100 0.25 0.25 0.23 0.23 0.25 0.23 0.25 0.26 __b

TDA 1 1 0

Blank 150 0.29 0.26 0.26 0.26 0.25 0.24 0.25 0.23 0.24

CDA 110

N40 150 0.26 0.27 0.27 0.25 0.24 0.23 0.23 0.23 0.23

CDA 110

Ti 150 0.28 0.26 0.27 0.26 0.24 0.24 0.24 0.24 0.24

CDA 110

C4130 150 0.24 0.25 0.25 0.25 0.23 0.22 NM 0.23 0.25

CDA 110

1625 150 0.25 0.24 0.25 0.22 0.22 0.23 NM 0.23 0.23
298 7. AGENT STABILITY UNDER STORAGE

Table 19 (cont.). Integrated peak areas for an absorbance band of CF3I in the 53.3 Pa spectra

tested in the dry condition with copper and the spectral comparisons

Integrated area from 2219 to 2274

(±0.011)

cm"'

Week number

Spectral

comparison

(±0.0018)'

j

Metal

Temp.

(°C) 40 44 48 52

CDA 110

Blank 23 0.25 NM 0.23 NM 0.997

CDA 110

N40 23 0.23 NM 0.24 NM 0.997

CDA 110

Ti 23 0.24 NM 0.25 NM 0.997

CDA 110

C4130 23 0.24 NM NM NM 0.999

CDA 110

0.24 0.23 NM NM 0 999

CDA 110

Blank 100 __b __b __b 0.999

CDA 110

Blank

-

150 0.24 0.23 0.22 0.23 0.980

CDA 110

N40 150 0.21 0.22 0.20 0.19 0.875

CDA 110

Ti 150 0.23 0.24 0.23 0.23 0.985

CDA 110

C4130 150 0.22 NM NM NM 0.997

CDA 110

1625 150 0.22 0.23 NM NM 0.994

Not measured

Cylinder emptied after 32 weeks as a result of leak

Number of measurements was eight.
AGENT STABILITY UNDER STORAGE

0.3

0.2 -

0.1

0.2

0.1

S 0.3 r

J I L

All samples at 150 °C

J I \ \ I 1_ J I I 0 4 8 12 16 20 24 28 32 36 40 44 48 52

0.3 r

O Blank

V N40

Ti A C4130

o 1625

Blank at 100 °C

J L J \ \ L 0 4 8 12 16 20 24 28 32 36 40 44 48 52

16 20 24 28 32 36 40 44 48 52

Time (weeks)

Figure 24. Integrated agent peak areas for all samples of CF3I tested in the dry condition with

copper at 23 °C, 100 °C, and 150 °C plotted as a function of time.
300 7. AGENT STABILITY UNDER STORAGE

Table 20. Integrated peak areas for an absorbance band in the 53.3 Pa spectra of CF3I tested in the moist condition without copper and the spectral comparisons

Integrated area from 2219 to 2274 cm"' (±0.011)

Week number

Metal

Temp.

(°C) 0 4 8 16 20 24 28 32 36

Blank 100 0.25 0.24 0.24 0.24 0.23 0.23 0.24 0.22 0.23

Blank 150 0.27 0.25 0.27 0.25 0.24 0.24 0.24 0.24 0.23

N40 150 0.27 0.25 0.27 0.25 0.24 0.24 0.25 0.24 0.20

Ti 150 0.28 0.25 0.27 0.27 0.24 0.24 0.24 0.24 0.24

C4130 150 0.24 0.25 0.23 0.23 0.22 0.21 NM'' 0.20 0.22

1625 150 0.25 0.24 0.24 0.22 0.22 0.22 NM 0.21 0.21

Table 20 (cont.). Integrated peak areas for an absorbance band in the 53.3 Pa spectra of CF3I

tested in the moist condition without copper and the spectral comparisons

Integrated area from 2219 to 2274 cm '

(±0.011)

Spectral

Comparison

(±0.0018)"

Week number

Metal

Temp.

(°C) 40 44 48 52

Blank 100 0.23 0.24 NM NM 0.999

Blank 150 0.24 0.23 0.24 0.23 0.984

N40 150 0.25 0.25 0.25 0.23 0.992

Ti 150 0.24 0.24 0.23 0.23 0.990

C4130 150 0.21 NM NM NM 0.990

1625 150 0.22 0.21 NM NM 0.983

Not measured

Number of observations was eight.
AGENT STABILITY UNDER STORAGE

c<

XIco

CD

0)u

co

i_o

XI<

co

o

CL

X>

(D

0 4 8 12 16 20 24 28 32 36 40 44 48 52

0.3 r

0.2

0.1

O Blank

V N40

Ti A C4130

o 1625

_J I

Blank at 100 °C

0 4 8 12 16 20 24 28 32 36 40 44 48 52

Time (weeks)

Figure 25. Integrated agent peak areas for all samples of CF3I tested in the moist condition

without copper at 100 °C and 150 °C plotted as a function of time.
302 7. AGENT STABILITY UNDER STORAGE

Table 21. Integrated peak areas for an absorbance band in the 53.3 Pa spectra of CF3I tested in the moist condition without copper and the spectral comparisons

Integrated area from 2219 to 2274 cm"' (±0.011)

Week number

Metal Temp.(°C) 0 4 8 16 20 24 28 32 36 CDA 110 Blank

100 0.25 0.24 0.23 0.24 0.22 0.22 0.23 0.23 0.23

CDA 110 Blank150 0.27 0.25 0.26 0.25 0.25 0.22 0.23 0.22 0.22

CDA 110 N40150 0.27 0.25 0.26 0.24 0.25 0.21 0.22 0.22 0.21

CDA 110

Ti

150 0.28 0.26 0.27 0.25 0.26 0.23 0.23 0.23 0.24

CDA 110 C4130

150 0.23 0.24 0.22 0.23 0.22 0.23 NM" 0.21 0.23

CDA 110 1625150 0.23 0.23 0.23 0.23 0.21 0.22 NM 0.22 0.22

Table 21 (cont.). Integrated peak areas for an absorbance band in the 53.3 Pa spectra of CF3I

tested in the moist condition without copper and the spectral comparisons

Integrated area from 2219 to 2274 cm '

(±0.011)

Week number

Spectral

comparison

(±0.0018)' Metal Temp.

(°C) 40 44 48 52

CDA 110 Blank

100 0.23 0.24 NM NM 0.998

CDA 110 Blank150 0.23 0.22 0.22 0.22 0.981

CDA 110 N40

150 0.21 0.21 0.20 0.19 0.949

CDA 110

Ti

150 0.23 0.23 0.22 0.24 0.990

CDA 110 C4130

150 0.2! NM NM NM 0.998

CDA 110

1625 150 0.21 NM NM NM 0.994

Not measured Number of observations was eight.
. AGENT STABILITY UNDER STORAGE

c

CP <

C

o

CDo

c

o

i_o

(/) <

C

D

O

D

0.3

0.2

0.1

0.3 r

0.2

0.1

All samples at 150 °C

J I I \ L J \ I L 0 4 8 12 16 20 24 28 32 36 40 44 48 52

o Blank

V N40

Ti Blank at 100 °C

A C4130

1 1 1 1 1 1 1

o

1 1

1625

1 1 1 0 4 8 12 16 20 24 28 32 36 40 44 48 52

Time (weeks)

Figure 26. Integrated agent peak areas for all samples of CF3I tested in the moist condition with

copper at 100 °C and 150 °C plotted as a function of time.
304 7. AGENT STABILITY UNDER STORAGE

7.3.5.2 Changes in the Impurity Peak Areas Appearing in the 5330 Pa Spectra. The peak

area data in the remaining tables represent a single value for each analysis obtained over the course of

the study. Tables 22-25 show the integrated peak areas for the absorbance band at 670 cm ' for COj

for all temperatures and test conditions of CF3I. Figures 27-30 are the respective graphical representations of the data. In both of the test conditions where ambient samples were tested (Figures 27 and

28), no changes in the peak areas were noted. The CO, concentration in the initial cylinders ranged

from 26 to 460 [xL/L (based on a 0. 1 % by volume standard of COj in Nj.) However, all samples

tested at elevated temperatures showed an increase in the CO2 peak area. Even samples that initially

contained CO2 as low as 26 pL/L (C4130 and 1625) experienced an increase in CO2 concentration to roughly 900 pL/L. The CO2 concentration in the cylinder containing 1625 metal in the dry condition

with and without copper and the moist condition without copper increased from 400 pL/L to nearly

1900 pL/L. The dramatic increase in the CO2 did not occur for the 1625 in the moist condition with

copper, but the initial CO2 content in this cylinder was only 26 pL/L. The COj content in the

cylinders with the other metals only increased about half this amount. The CO2 production continued

to increase after 52 weeks. Also, all samples tested at 100 °C showed an increase in the CO2 peak

area. The actual concentration was about half that in the 150 °C cylinders, except for C4130 and 1625

in the dry condition without copper, which had very little COj present at the outset of testing.

Tables 26-29 show the integrated peak areas for the absorbance band at 700 cm"' in the 5330 Pa

spectrum for the CF3H impurity. Figures 31-34 show the graphs for the respective data. There was

no increase in the CF3H peak area for any of the conditions at 23 °C or 100 °C. However, at 150 °C

the generation of CF3H was increasing with time and, like the CO2, was still increasing after 52 weeks.

Noticeably different for the CF3H increase was a longer time before the slope started to show an

increase. 1625 in both the dry and moist conditions without copper (Figures 31 and 33, respectively)

showed a more rapid increase than the other metals. In the dry condition with copper (Figure 32) it was the nitronic 40 metal that showed a more rapid increase.

Tables 30-33 show the integrated peak areas for the absorbance band at 950 cm"' of the fluorinated alkene. Figures 35-38 show the graphs for the respective data. In this case, none of the initial samples showed any absorbance in the 950 cm"' band and at 23 °C for as long as 48 weeks there was

no evidence of formation. After about 8 weeks at 100 °C, a small but easily resolved peak started to appear. In both the dry and moist condition without copper (Figures 35 and 37, respectively), the

concentration was still increasing after 8 weeks. In both the dry and moist condition with copper

(Figures 36 and 38, respectively), the concentration increased then equilibrated. However, after only 4

weeks at 150 °C, the concentration of the fluorinated alkene increased rapidly then began to equilibrate

after 20 weeks. In the dry condition without copper at 150 °C (Figure 35), the fluorinated alkene

content maximized around 20 weeks, then equilibrated. There may have been some differentiation of

the other metals with copper present at 150 °C (Figures 36 and 38). The blank containing only copper

strips, and the Ti- 15-3-3-3 with copper strips in the dry condition (Figure 36) maximized in about 8

weeks, then started to drop and after 52 weeks was approaching a very low level. The Nitronic 40

and 1625 metals with the copper caused more of the fluorinated alkene to appear and the decline was

even more gradual. The C4130 in combination with copper resulted in the most fluorinated alkene

appearing and there was no decline in the compound.

Figure 38 shows the data for the moist condition with copper. Note the similarity of this data to Figure 36 with respect to the alkene content and metal, but equilibration occurred after 20 weeks.

Also of note is that the blank, which contains only copper strips, did not produce detectable amounts

of the alkene at 150 °C. However, the blank at 100 °C showed the presence of detectable fluorinated

alkene; the peak area did appear to equilibrate with time. A possible explanation is that the copper is causing the double bond to break, and a higher temperature is needed to activate the reaction.
7. AGENT STABILITY UNDER STORAGE 305

Although the characteristic band at 2150 cm ' for CO appears in all of the elevated temperature

samples for CF3I, this absorbance peak was not integrated as a function of time as with the other

impurity peaks. It is reasonable to expect the presence of CO along with COj.

The initial and final spectra for all samples of CF3I are found in appendixes C, D, E, and F. These spectra are included for reference in order of dry condition, without copper (Appendix C); dry

condition, with copper (Appendix D); moist condition, without copper (Appendix E) and moist

condition, with copper (Appendix F).

7.3,6 Summary of the CF3I Area Data. The peak area data for an absorbance band of the CF3I

presented in the above graphs (Figures 23-26) suggest that the areas are decreasing slightly as a

function of time at elevated temperatures, especially with moisture present. Table 34 is a summary of

the data in Tables 18-21 showing only the initial and final peak areas along with the time of aging at

all temperatures and conditions. The last column shows whether the final area was significantly

decreased at the 99 % confidence interval. None of the samples at 23 or 100 °C showed a statistically

significant change. However, some samples at 150 °C aging started to show statistically significant

decreases, especially those that aged for 52 weeks. In the dry condition without copper at 150 °C

none of the final areas were significantly decreased (the initial area for this sample was higher than

expected and might be an outlier.) In the dry condition with copper, three of the areas decreased

significantly and the other two which did not age as long appear to be decreasing and in time will

become significantly lower. In the moist condition without copper, four of the five samples show a

significant decrease, the other sample appears to be going down, also. Finally, in the moist condition

with copper, three of the samples show a significant decrease. The other two samples are decreasing

and at longer time might become significant. Even though some of the areas at 150 °C are showing

"statistically significant" differences, the actual loss in agent is probably quite small and poses no

problem to the fire extinguishing capability of the CF3I.

The data presented in Table 35 show the spectral comparisons for all of the CF3I samples aged at 150 °C compared to their respective samples maintained at 23 °C. Keep in mind that the spectral

comparisons for the elevated temperature samples were for the longest aged sample compared to its

initial spectrum and represents a single value, whereas the spectral comparisons for the 23 °C samples

are an average of each aged spectrum compared to its respective initial spectrum. Although not all samples have been aged the same length of time, these data do suggest some significant changes in the

aged spectra with time. Also one should keep in mind that the impurity peaks are small in comparison

to the agent peaks, so the increase in the areas of all combined peaks resulted in significant changes.

Since the agent concentration in the cylinders is greater than 100 000 fiL/L, and the highest measured

concentration of CO2 impurity was only 1900 |jL/L, the amount of impurity from all sources is probably not enough to have an affect on the overall performance of the CF3L Some fluorinated

alkenes are highly toxic, and if the peak at 950 cm"' is a result of an alkene, the only question is whether enough is being generated to form a harmful mixture.

The data for CF3I suggested that the weak C-I bond in CF3I was indeed breaking at the 100 and

150 °C temperatures. Once the bond broke the radical was free to recombine with another CF3

radical, an H atom or other radicals in the matrix. Since the CF3H concentration in the cylinders

continued to increase with time and the CF3I agent peak continued to show slight decreases, the CF3

radical combined with a hydrogen atom to form a more stable CF3H molecule. The presence of

moisture in the sample probably provided an increased source of hydrogen atoms, therefore the

degradation process was accelerated.
306 7. AGENT STABILITY UNDER STORAGE

Table 22. Integrated peak areas for an absorbance band of COj in the 5330 Pa spectra of CF3I in

the dry condition without copper

Integrated area from 667 to 671 cm (±0.0048)

Week number

Metal

Temp.

CC) 0 4 8 16 20 24 28 32 36

Blank 23 0.09 0.09 0.09 0.09 0.09 0.09 NM" 0.09 NM

N40 23 0.08 0.08 0.08 0.08 0.08 0.07 NM 0.08 NM

Ti 23 b 0.08 0.09 0.08 0.09 0.08 NM 0.08 NM

C4130 23 0.01 0.01 0.01 0.01 0.01 0.01 NM 0.01 NM

1625 23 0.18 0.17 0.16 0.16 0.16 0.17 NM 0.16 0.17

Blank 100 0.14 0.15 0.14 0.15 0.16 0.16 0.17 0.17 0.16

N40 100 0.14 0.14 0.14 0.16 0.16 0.16 0.17 0.17 c

Ti 100 0.14 0.14 0.14 0.15 0.15 0.16 0.17 0.16 0.16

C4130 100 0.01 0.01 0.02 0.02 0.03 0.03 NM 0.04 0.05

1625 100 0.01 0.01 0.01 0.02 0.02 0.02 NM 0.03 0.04

Blank 150 0.18 0.24 0.25 0.27 0.30 0.30 0.31 0.35 0.36

N40 150 0.17 0.26 0.27 0.28 0.30 0.29 0.30 0.31 0.32

Ti 150 0.17 0.20 0.22 0.24 0.28 0.25 0.26 0.29 0.30

C4130 150 0.01 0.09 0.12 0.16 0.20 0.23 NM 0.27 0.32

1625 150 0.17 0.68 0.66 0.69 0.68 0.72 NM 0.69 0.72
7. AGENT STABILITY UNDER STORAGE 307

Table 22 (cont.). Integrated peak areas for an absorbance band of COj in the 5330 Pa spectra of

CF3I in the dry condition without copper

Integrated area from 667 to 671 cm"' (±0.0048)

Week number

Metal

Temp.

(°C) 40 44 48 52

Blank 23 ft OR ft ft9 N40 23 ft ft? ft\J,\JO 08 NM

Ti 23 0.08 NM 0.09 NM

C4130 23 0.01 NM NM NM

1625 23 ft 16 ft 17 NM NM

Blank 100 ft 1 8 ft 19 NM NM

N40 100 c c c c

Ti 100 0.18 0.19 NM NM

C4130 100 0.05 NM NM NM

1625 100 0.04 NM NM NM

Blank 150 0.38 0.41 0.39 0.42

N40 150 0.32 0.32 0.31 0.33

Ti 150 0.32 0.34 0.35 0.40

C4130 150 0.36 NM NM NM

1625 150 0.73 0.75 NM NM

Not measured

Lost spectrum file Cylinder emptied after 32 weeks from a leak.
308 7. AGENT STABILITY UNDER STORAGE

0 4 8 12 16 20 24 28 32 36 40 44 48 52

0.3

0.2

0.1

0.0

o Blank

V N40

Ti A C4130

o 1625

0 4

All samples at 100 °C

12 16 20 24 28 32 36 40 44 48 52

0.3

0.2

All samples ot 23 °C

-4—4^—4^- 0.0

0 4 8 12 16 20 24 28 32 36 40 44 48 52

Time (weeks)

Figure 27. Integrated peak areas for CO^ for all samples of CF3I tested in the dry condition

without copper at 23 °C, 100 °C, and 150 °C plotted as a function of time.
7. AGENT STABILITY UNDER STORAGE 309

Table 23. Integrated peak areas of an absorbance band of COj in the 5330 Pa spectra of CF3I

tested in the dry condition with copper

Integrated area from 667 to 671 cm'' (±0.0048)

Week number

Metal

Temp.

(°C) 0 4 8 16 20 24 28 32 36

CDA 110

Blank 23 0.09 0.08 0.08 0.08 0.09 0.08 0.08 NM

CDA 110

N40 23 0.08 0.08 0.08 0.08 0.08 0.08 NM 0.08 NM

CDA 110

Ti 23 0.08 0.08 0.08 0.08 0.08 0.07 NM 0.08 NM

CDA 110

C4130 23 0.01 0.01 0.01 0.01 0.01 0.01 NM 0.01 NM

CDA 110

1625 23 0.17 0.16 0.15 0.15 0.15 0.16 NM 0.15 0.16

CDA 110

Blank 100 0.14 0.14 0.14 0.15 0.15 0.16 0.17 0.18 __b

CDA 1 10

Blank 150 0.17 0.23 0.25 0.27 0.30 0.27 0.28 0.30 0.30

CDA 110

N40 150 0.16 0.23 0.25 0.26 0.29 0.27 0.27 0.29 0.29

CDA 110

Ti 150 0.16 0.22 0.24 0.25 0.28 0.26 0.26 0.27 0.28

CDA 110

C4130 150 0.01 0.12 0.15 0.19 0.20 0.22 NM 0.23 0.26

CDA 110

1625 150 0.17 0.33 0.40 0.48 0.50 0.54 NM 0.54 0.57
310 7. AGENT STABILITY UNDER STORAGE

Table 23 (cont.). Integrated peak areas of an absorbance band of COj in the 5330 Pa spectra of

CFjI tested in the dry condition with copper

Integrated area from 667 to 671 cm"' (±0.0048)

Week number

Metal

Temp.

(°C) 40 44 48 52

CDA 110

Blank 23 0.08 NM 0.08 NM

CDA 110

0 08 NM 0 08 NM

CDA 110

Ti 23 0.08 NM 0.08 NM

CDA 110

C4130 23 0.01 NM NM NM

CDA 110

1625 23 0.16 0.16 NM NM

CDA 110

Blank 100 __b __b __b __b

CDA 110

Blank 150 0.30 0.30 0.30 0.32

CDA 110

N40 150 0.30 0.30 0.30 0.32

CDA 110

Ti 150 0.27 0.25 0.27 0.29

CDA 110

C4130 150 0.28 NM NM NM

CDA 110

1625 150 0.58 0.59 NM NM

Not measured

Cylinder emptied after 32 weeks from a leak
AGENT STABILITY UNDER STORAGE 311

0.8 r

0.6

0.4

0.2

0.0

0.3

0.2 -

0.1 -

0.0

All samples at 150 °C

J I 0 4 8 12 16 20 24 28 32 36 40 44 48 52

o Blank

V N40

Ti A C4130

o 1625

Blank at 100 °C

J \ I I \ \ I I \ I I 0 4 8 12 16 20 24 28 32 36 40 44 48 52

0.3

0.2

0.0

All samples at 23 °C

0 4 8 12 16 20 24 28 32 36 40 44 48 52 .- .

Time (weeks)

Figure 28. Integrated peak areas for COj for all samples of CF3I tested in the dry condition with

copper at 23 °C, 100 °C, and 150 °C plotted as a function of time.
312 7. AGENT STABILITY UNDER STORAGE

Table 24. Integrated peak areas for an absorbance band of COj in the 5330 Pa spectra of CF3I

tested in the moist condition without copper

Integrated area from 667 to 671 cm"' (±0.0048)

Week number

Metal

Temp.

(°C) 0 4 8 16 20 24 28 32 36

Blank 100 0.13 0.15 0.16 0.17 0.18 0.18 0.18 0.19 0.18

Blank 150 0.14 0.23 0.26 0.28 0.34 0.37 0.40 0.44 0.43

N40 150 0.14 0.21 0.23 0.23 0.25 0.25 0.25 0.27 0.27

Ti 150 0.13 0.22 0.24 0.25 0.28 0.28 0.29 0.32 0.30

C4130 150 0.01 0.10 0.13 0.20 0.23 0.27 NM^ 0.29 0.33

1625 150 0.16 0.68 0.66 0.66 0.67 0.70 NM 0.68 0.71

Table 24 (cont.). Integrated peak areas for an absorbance band of CO2 in the 5330 Pa spectra of

CF3I tested in the moist condition without copper

Integrated area from 667 to 671 cm '

(±0.0048)

Week number

Metal

Temp.

(°C) 40 44 48 52

Blank 100 0.20 0.21 NM NM

Blank 150 0.45 0.46 0.45 0.49

N40 150 0.27 0.27 0.27 0.29

Ti 150 0.34 0.34 0.34 0.37

C4130 150 0.35 NM NM NM

1625 150 0.71 0.75 NM NM

Not measured.
AGENT STABILITY UNDER STORAGE

CM O

CJ

C

om

o

c

o

o

D

c

o

0.0^

0 4 8 12 16 20 24 28 32 36 40 44 48 52

0.3 r

D

0)

o

q; Q.

X)

a

CD

0 4 8 12 16 20 24 28 32 36 40 44 48 52

Time (weeks)

Figure 29. Integrated peak areas for COj for all samples of CF3I tested in the moist condition

without copper at 100 °C and 150 °C plotted as a function of time.
314 7. AGENT STABILITY UNDER STORAGE

Table 25. Integrated peak areas of an absorbance band of COj in the 5330 Pa spectra of CF3I

tested in the moist condition with copper

Integrated area from 667 to 671 cm ' (±0.0048)

Week number

Metal Temp.

{°C) 0 4 8 16 20 24 28 32 36 CDA 110 Blank

100 0.13 0.16 0.17 0.18 0.19 0.19 0.19 0.20 0.20

CDA 110 Blank150 0.13 0.23 0.26 0.28 0.31 0.30 0.32 0.34 0.34

CDA 110 N40150 0.13 0.22 0.25 0.29 0.33 0.33 0.35 0.37 0.37

CDA 110

Ti

150 0.13 0.20 0.22 0.26 0.29 0.29 0.33 0.36 0.38

CDA 110 C4130150 0.01 0.13 0.16 0.21 0.22 0.25 NM' 0.27 0.30

CDA 110

1625

150 0.01 0.14 0.18 0.25 0.28 0.30 NM 0.32 0.35

Table 25 (cont.). Integrated peak areas of an absorbance band of CO2 in the 5330 Pa spectra of

CF3I tested in the moist condition with copper

Integrated area from 667 to 671 cm ' (±0.0048)

Week number

Metal Temp.

(°C) 40 44 48 52 CDA 110 Blank

100 0.21 0.23 NM NM

CDA 110 Blank150 0.35 0.34 0.33 0.36

CDA 110 N40150 0.37 0.38 0.36 0.39

CDA 110

Ti 150 0.39 0.41 0.39 0.43

CDA 110 C4130

150 0.33 NM NM NM

CDA 110

1625 150 0.37 NM NM NM

Not measured
AGENT STABILITY UNDER STORAGE

0.8 r

o

c

o

CD

<DO

c

o

i_o

in <

c

o

D

O

CL

0.3

0.0

All samples at 150 °C

0 4 8 12 16 20 24 28 32 36 40 44 48 52

0 4 8 12 16 20 24 28 32 36 40 44 48 52

Time (weeks)

Figure 30. Integrated peak areas for COj for all samples of CF3I tested in the moist condition

with copper at 100 °C and 150 °C plotted as a function of time.
316 7. AGENT STABILITY UNDER STORAGE

Table 26. Integrated peak areas of an absorbance band for CFjH in the 5330 Pa spectra of CF3I

tested in the dry condition without copper

^ ^ Integrated area from 698 to 701 cm"' (±0.0011)

Week number

Metal

Temp.

(°C) 0 4 8 16 20 24 28 32 36

Blank 23 0.01 0.01 0.01 0.01 0.01 0.01 NM'' 0.01 NM

N40 23 0.01 0.01 0.01 0.01 0.01 0.01 NM 0.01 NM

Ti 23 b 0.01 0.01 0.01 0.01 0.01 NM 0.01 NM

C4130 23 ND ND ND ND ND NM ND NM

1625 23 0.04 0.04 0.04 0.04 0.04 0.04 NM 0.04 0.04

Blank 100 0.03 0.03 0.03 0.03 0.03 0.03 0.03 0.03 0.03

N40 100 0.03 0.03 0.03 0.03 0.03 0.03 0.03 0.03 d

Ti 100 0.03 0.03 0.03 0.03 0.03 0.03 0.03 0.03 0.03

C4130 100 ND ND ND ND ND ND NM ND ND

1625 100 ND ND ND ND ND ND NM ND ND

Blank 150 0.02 0.02 0.03 0.04 0.05 0.06 0.08 0.11 0.15

N40 150 0.02 0.03 0.04 0.06 0.07 0.09 0.10 0.11 0.12

Ti 150 0.02 0.02 0.03 0.03 0.03 0.03 0.04 0.04 0.05

C4130 150 ND <0.01 0.01 0.03 0.04 0.06 NM 0.12 0.15

1625 150 0.04 0.06 0.10 0.22 0.26 0.30 NM 0.35 0.37
7. AGENT STABILITY UNDER STORAGE 317

Table 26 (cont.). Integrated peak areas of an absorbance band for CF3H in the 5330 Pa spectra of

CF3I tested in the dry condition without copper

Integrated area from 698 to 701 cm"' (±0.0011)

Week number

Metal

Temp.

(°C) 40 44 48 52

Blank 23 0.01 NM 0.01 NM

N40 23 0.01 NM 0.01 NM

Ti 23 0.01 NM 0.01 NM

C4130 23 ND NM NM NM

1625 23 0.04 0.04 NM NM

Blank 100 0.03 0.04 NM NM

N40 100 __d __d __d

Ti 100 0.03 0.04 NM NM

C4130 100 ND NM NM NM

1625 100 ND NM NM NM

Blank 150 0.19 0.24 0.28 0.32

N40 150 0.12 0.13 0.14 0.15

Ti 150 0.06 0.07 0.10 0.12

C4130 150 0.19 NM NM NM

1625 150 0.39 0.41 NM NM

Not measured

Lost spectrum file Not detected

Cylinder emptied after 32 weeks from a leak.
318 7. AGENT STABILITY UNDER STORAGE

to

C

o

QQ

(DU

Co

i_o

(/) <

o

o

CL

XI

0)

0.04

«3-

0.02

< 0.00 <i o Ct

O.OAo

0.02

0.00^^

All samples at 150 °C

12 16 20 24 28 32 36 40 44 48 52

-Q O- -O O O O O OAll samples at 100 °C

-o—o Of- -o—o0 4 8 12 16 20 24 28 32 36 40 44 48 52

All samples at 23 °C

-o—©- -O" O- -o-

-A A-

_J L

-AI 0 4 8 12 16 20 24 28 32 36 40 44 48 52

Time (weeks)

Figure 31. Integrated peak areas for CFjH for all samples of CF3I tested in the dry condition

without copper at 23 °C, 100 °C, and 150 °C plotted as a function of time.
7. AGENT STABILITY UNDER STORAGE 319

Table 27. Integrated peak areas of an absorbance band for CF3H in the 5330 Pa spectra of CF3I

tested in the dry condition with copper

Integrated area from 698 to 701 cm"' (±0.0011)

Week number

Metal

Temp.

(°C) 0 4 8 16 20 24 28 32 36

CDA 110

DialIK /.J 0u.w011 0 01 0 01 0\J.\J011 0 01 0 01 NM 0 01

CDA 110

N40 23 0.01 0.01 0.01 0.01 0.01 0.01 NM^ 0.01 NM

pn A 1 1

n

Ti 23 0.01 0.01 0.01 0.01 0.01 0.01 NM 0.01 NM

CDA 110

C4130 23 ND ND ND ND XTT> ND ND NM ND NM

CDA 110

1625 23 0.04 0.04 0.04 0.04 0.04 0.04 NM 0.04 0.04

CDA 110

Blank 100 0.03 0.03 0.03 0.03 0.03 0.03 0.03 0.03

CDA 110

Blank 150 0.02 0.02 0.03 0.06 0.07 0.09 0.12 0.15 0.19

CDA 110

N40 150 0.02 0.03 0.05 0.16 0.22 0.27 0.32 0.37 0.39

CDA 110

Ti 150 0.02 0.02 0.03 0.05 0.07 0.09 0.10 0.12 0.14

CDA 110

C4130 150 ND 0.01 0.02 0.03 0.05 0.05 NM 0.07 0.09

CDA 110

1625 150 0.04 0.04 0.04 0.05 0.07 0.08 NM 0.13 0.16
320 7. AGENT STABILITY UNDER STORAGE

Table 27 (cont.)- Integrated peak areas of an absorbance band for CF3H in the 5330 Pa spectra of

CF3I tested in the dry condition with copper

Integrated area from 698 to 701 cm ' (±0.0011)

Week number

Metal

Tpmn

(°C) 40 44 48 52

CDA 110

Blank 23 0.01 NM 0.01 NM

CDA 110

N40 23 0.01 NM 0.01 NM

CDA 110

Ti 23 0.01 NM 0.01 NM

CDA 110

C4130 23 ND NM NM NM

CDA 110

1625 23 0.04 0.04 NM NM

CDA 110

Blank 100 c c c c

Blank 150 0.23 0.28 0.33 0.37

CDA 110

N40 150 0.41 0.45 0.47 0.49

CDA 110

Ti 150 0.16 0.18 0.20 0.22

CDA no

C4130 150 0.11 NM NM NM

CDA 110

1625 150 0.21 0.23 NM NM

Cylinder emptied after 32 weeks from a leak
7. AGENT STABILITY UNDER STORAGE 321

0.6

Li_ O

C

o

CD

0)u

coO

(/) JD<

cD

O

(D <

O

(U CLX)

(U

cn

0)

0.04 -

0.02

0.00

0.02

0.00^^

o Blank

V N40

Ti A C4130

o 1625

All samples at 150 °C

0 4 8 12 16 20 24 28 52 36 40 44 48 52

Blank at 100 °C

I I I I I I I I 0 4 8 12 16 20 24 28 52 36 40 44 48 52

0.04 <^—o—

All samples at 23 °C

-O—o~

-A A-

_1 L

-Q Q- -o- -o

-A-

-L.

-A ^A- -A

I 0 4 8 12 16 20 24 28 32 56 40 44 48 52

Time (weeks)

Figure 32. Integrated peak areas for CF3H for all samples of CF3I tested in the dry condition

with copper at 23 °C, 100 °C, and 150 °C plotted as a function of time.
322 7. AGENT STABILITY UNDER STORAGE

Table 28. Integrated peak areas of an absorbance band for CF3H in the 5330 Pa spectra of CF3I

tested in the moist condition without copper

Integrated area from 698 to 701 cm"' (±0.0011)

Week number

Metal

Temp.

0

_______ 4 8 16 20 24 28 32 36

Blank 100 0.03 0.03 0.03 0.03 0.03 0.03 0.03 0.03 0.03

Blank 150 0.01 0.02 0.03 0.07 0.11 0.18 0.24 0.29 0.34

N40 150 0.01 0.03 0.05 0.08 0.10 0.12 0.14 0.15 0.16

Ti 150 0.02 0.02 0.03 0.05 0.07 0.09 0.12 0.14 0.17

C4130 150 ND" 0.01 0.01 0.05 0.08 0.11 NM" 0.18 0.22

1625 150 0.04 0.12 0.25 0.36 0.38 0.41 NM 0.44 0.46

Table 28 (cont.). Integrated peak areas of an absorbance band for CF3H in the 5330 Pa spectra of

CF3I tested in the moist condition without copper

Integrated area from 698 to 701 cm ' (±0.0011)

Week number

Metal

Temp.

(°C) 40 44 48 52

Blank 100 0.03 0.03 NM NM

Blank 150 0.37 0.40 0.42 0.45

N40 150 0.17 0.19 0.20 0.21

Ti 150 0.19 0.22 0.24 0.27

C4130 150 0.27 NM NM NM

1625 150 0.48 0.51 NM NM

Not detected

Not measured.
7. AGENT STABILITY UNDER STORAGE 323

(J

c

D

CD

(DO

co

o

C/) <

co

0.6

o Blank

V N40

Ti A C4130

o 1625

All samples at 150 °C

J \ i \ \

I I 0 4 8 12 16 20 24 28 32 36 40 44 48 52

o

o

0) Q_

0)

0.04

0.02

0.00

Blank at 100 °C

0 4 8 12 16 20 24 28 32 36 40 44 48 52

Time (weeks)

Figure 33. Integrated peak areas for CF3H for all samples of CF3I tested in the moist condition

without copper at 100 °C and 150 °C plotted as a function of time.
324 7. AGENT STABILITY UNDER STORAGE

Table 29. Integrated peak areas of an absorbance band for CF3H in the 5330 Pa spectra of CF3I

tested in the moist condition with copper

Integrated area from 698 to 701 cm ' (±0.0011)

Week number

Metal Temp,

0 4 8 16 20 24 28 32 36 CDA 110 Blank

100 0.03 0.03 0.03 0.03 0.03 0.03 0.03 0.03 0.03

CDA 110 Blank150 0.01 0.02 0.02 0.06 0.09 0.14 0.19 0.24 0.29

CDA 110 N40150 0.01 0.02 0.03 0.07 0.10 0.14 0.19 0.26 0.35

CDA 110

Ti

150 0.01 0.02 0.05 0.09 0.11 0.13 0.15 0.16 0.18

CDA 110 C4130150 <0.01 0.01 0.02 0.04 0.05 0.05 NM" 0.06 0.07

CDA 110

1625

150 0.03 0.01 0.01 0.03 0.05 0.07 NM 0.13 0.16

Table 29 (cont.). Integrated Peak Areas of an Absorbance Band for CF3H in the 5330 Pa Spectra

of CF3I tested in the Moist Condition with Copper

Integrated Area from 698 to 701 cm ' (±0.0011)

Week Number

Metal Temp.

40 44 48 52 CDA 110 Blank

100 0.03 0.03 NM NM

CDA 110 Blank150 0.34 0.37 0.40 0.44

CDA 110 N40

150 0.41 0.50 0.55 0.63

CDA 110

Ti

150 0.19 0.20 0.22 0.24

CDA 110 C4130150 0.08 NM NM NM

CDA 110

1625 150 0.21 NM NM NM

Not measured
7. AGENT STABILITY UNDER STORAGE 325

All samples at 150 °C

u

M—O

XJc

D

0)o

co

i_o

CO <

c

D

0.0

0 4 12 16 20 24 28 32 36 40 44 48 52

D

OJ <

D

D

CL

0.04 -

0.02 -

0.00 - Blank at 100 °C

J L J I \ I I I 0 4 8 12 16 20 24 28 32 36 40 44 48 52

Time (weeks)

Figure 34. Integrated peak areas for CF3H for all samples of CF3I tested in the moist condition

with copper at 100 °C and 150 °C plotted as a function of time.
326 7. AGENT STABILITY UNDER STORAGE

Table 30. Integrated peak areas for an absorbance band of the fluorinated alkene in the 5330 Pa

spectra of CF3I tested in the dry condition without copper

Integrated area from 944 to 953 cm"' (±0.011)

Week number

Metal

Temp.

(°C) 0 4 8 16 20 24 28 32 36

Blank 23 ND" ND ND ND ND ND NM" ND NM

N40 23 ND ND ND ND ND ND NM ND NM

Ti 23 c ND ND ND ND ND NM ND NM

C4130 23 ND ND ND ND ND ND NM ND NM

1625 23 ND ND ND ND ND ND NM ND ND

Blank 100 ND ND ND 0.01 0.01 0.01 .02 .02 0.02

N40 100 ND <0.01 0.01 0.02 0.03 0.03 .03 .04

d

Ti 100 ND ND 0.01 0.02 0.02 0.03 .04 .04 0.05

C4130 100 ND <0.01 0.01 0.04 0.05 0.06 NM 0.06 0.07

1625 100 ND ND 0.01 0.02 0.03 0.03 NM 0.06 0.07

Blank 150 ND 0.06 0.09 0.10 0.10 0.11 0.10 0.11 0.11

N40 150 ND 0.01 0.10 0.10 0.11 0.11 0.11 0.11 0.11

Ti 150 ND 0.03 0.06 0.07 0.12 0.08 0.08 0.09 0.09

C4130 150 ND 0.06 0.09 0.11 0.11 0.12 NM 0.11 0.12

1625 150 ND 0.06 0.06 0.12 0.17 0.12 NM 0.15 0.13
7. AGENT STABILITY UNDER STORAGE 327

Table 30 (cont.). Integrated peak areas for an absorbance band of the fluorinated alkene in the

5330 Pa spectra of CF3I tested in the dry condition without copper

Integrated area from 944 to 953 cm"' (±0.011)

Week number

Metal

Temp.

( C) A A41) A A 44 A 0 4o

Blank 23 ND NM NU XTX NMK

N40 23 ND NM ND XTA NMH

Ti 23 ND NM ND XTA K NM

C4130 23 ND NM NM NM

1625 23 ND ND NM NM

Blank 100 0.02 0.03 XTA K NM XTA NMK

N40 100 d d d d

Ti 100 0.05 0.06 XTA NMK XTX NMK

C4130 100 0.07 NM XTX MMH XTX NMH

1625 100 0.08 NM NM NM

Blank 150 0.11 0.11 0.11 0.11

N40 150 0.12 0.12 0.12 0.12

Ti 150 0.10 0.10 0.10 0.10

C4130 150 0.14 NM NM NM

1625 150 0.12 0.12 NM NM

Not detected

Not measured

Lost spectrum file Cylinder emptied after 32 weeks from a leak.
328 7. AGENT STABILITY UNDER STORAGE

0.3

CDC

(D

X)

cu

+-'o

o

0)

All samples at 150 °C

\ \ \ I I 0 4 8 12 16 20 24 28 32 36 40 44 48 52

"DC

CD

CDa

co

o

(/) <

Co

o

cu <

D

0) CL

X)

CUo

en

cu

0.08 -

0.04

0.00 <^

0.04

0.02

0.00 a

All samples at 100 °C

_J I I \ I 0 4 8 12 16 20 24 28 32 36 40 44 48 52

All samples at 23 °C

-o-

-L _L_

-O-

_L_

-O

I 0 4 8 12 16 20 24 28 32 36 40 44 48 52

Time (weeks)

Figure 35. Integrated peak areas for the fluorinated alkene for all samples of CF3I tested in the

dry condition without copper at 23 °C, 100 °C, and 150 °C plotted as a function of

time.
7. AGENT STABILITY UNDER STORAGE 329

Table 31. Integrated peak areas for an absorbance band of the fluorinated alkene in the 5330 Pa

spectra of CF3I tested in the dry condition with copper

Integrated area from 944 to 953 cm"' (±0.011)

Week number

Metal

Temp.

(°C) 0 4 8 16 20 24 28 32 36

CDA 110

DianK Z J IN LJ IN LJ L\LJ iMiVl CDA 110

N40 23 ND ND ND ND ND ND NM ND NM

\^L/J\ IIU

Ti 23 ND ND ND ND ND ND NM ND NM

CDA 110

C4130 23 ND ND ND ND ND ND NM ND NM

CDA 110

1625 23 ND ND ND ND ND ND NM ND ND

CDA 110

Blank 100 ND ND ND 0.01 0.01 0.01 .02 .02

c

CDA 110

Blank 150 ND 0.11 0.14 0.08 0.08 0.06 0.06 0.06 0.05

CDA 110

N40 150 ND 0.13 0.14 0.13 0.14 0.12 0.12 0.12 0.10

CDA 110

Ti 150 ND 0.10 0.12 0.10 0.09 0.07 0.06 0.05 0.04

CDA 110

C4130 150 ND 0.11 0.14 0.18 0.19 0.19 NM 0.19 0.19

CDA 110

1625 150 ND 0.04 0.10 0.14 0.16 0.14 NM 0.14 0.12
330 7. AGENT STABILITY UNDER STORAGE

Table 31 (cont.)- Integrated peak areas for an absorbance band of the fluorinated alkene in the

5330 Pa spectra of CF3I tested in the dry condition with copper

Integrated area from 944 to 953 cm ' (±0.011)

Week number

Metal

Temp.

(°C) 40 44 48 52

CDA 110

Blank 23 ND NM ND NM

CDA 110

N40 23 ND NM ND NM

CDA 110

Ti 23 ND NM ND NM

CDA 110

C4130 23 ND NM NM NM

CDA 110

1625 23 ND ND NM NM

CDA 110

Blank 100 c c c c

CDA 110

Blank 150 0.05 0.04 0.04 0.03

CDA 110

N40 150 0.10 0.10 0.09 0.08

CDA 110

Ti 150 0.03 0.03 0.02 0.02

CDA 110

C4130 150 0.19 NM NM NM

CDA 110

1625 150 0.12 0.12 NM NM

Not detected

Not measured

Cylinder emptied after 32 weeks from a leak.
AGENT STABILITY UNDER STORAGE

0.3 r

0.2

0.1 -

.-^ 0.0 -

0.08

0.04

0.00 (>

All samples at 150 °C

4 8 12 16 20 24 28 32 36 40 44 48 52

o Blank

V N40

Ti A C4130

o 1625

Blank at 100 °C

0 4 8 12 16 20 24 28 32 36 40 44 48 52

0.04

0.02

0.00 -o—oJ L

All samples at 23 °C

-o—o—

o

J \ \ L I I I I 0 4 8 12 15 20 24 28 32 36 40 44 48 52

Time (weeks)

Figure 36. Integrated peak areas for the fluorinated alkene for all samples of CF3I tested in dry condition with copper at 23 °C, 100 °C, and 150 °C plotted as a function of

*'me.
332 7. AGENT STABILITY UNDER STORAGE

Table 32. Integrated peak areas of an absorbance band of the fluorinated aikene in the 5330 Pa

spectra of CF3I tested in the moist condition without copper

Integrated area from 944 to 953 cm"' (±0.011)

Week number

Metal

Temp.

( C) 0 4

08 16 20 24 0 28 32 36

Blank 100 ND^ ND 0.01 0.02 0.03 0.03 0.04 0.04 0.04

Blank 150 ND 0.08 0.10 0.16 0.20 0.18 0.17 0.18 0.18

N40 150 ND 0.09 0.10 0.10 0.11 0.12 0.11 0.12 0.12

Ti 150 ND 0.08 0.07 0.10 0.10 0.10 0.09 0.09 0.09

C4130 150 ND 0.11 0.13 0.23 0.24 0.19 NM" 0.19 0.18

1625 150 ND 0.12 0.14 0.12 0.13 0.12 NM 0.12 0.12

Table 32 (cont.). Integrated peak areas of an absorbance band of the fluorinated aikene in the 5330

Pa spectra of CF3I tested in the moist condition without copper

Integrated area from 944 to 953 cm"' (±0.011)

Week number

Metal

Temp.

(°C) 40 44 48 52

Blank 100 0.04 0.05 NM NM

Blank 150 0.17 0.18 0.18 0.19

N40 150 0.12 0.12 0.12 0.12

Ti 150 0.09 0.08 0.09 0.09

C4130 150 0.20 NM NM NM

1625 150 0.12 0.13 NM NM

Not detected

Not measured.
7. AGENT STABILITY UNDER STORAGE 333

C

CD <

-oD

C

'i_o

o

"D

£ZD

CD

<DO

C

o

i_O

CO <

c

o

0.3

0.2

0.1

0.0

All samples at 150 °C

\ \ \ \ \ L_ 0 4 8 12 16 20 24 28 32 36 40 44 48 52

D

CD <

D

CD CL

TD

0)

0.08

0.04 -

0.00 (>

0 4 8 12 16 20 24 28 32 36 40 44 48 52

Time (weeks)

Figure 37. Integrated peak areas for the fluorinated alkene for all samples of CF3I tested in the

moist condition without copper at 100 °C and 150 °C plotted as a function of time.
334 7. AGENT STABILITY UNDER STORAGE

Table 33. Integrated peak areas of an absorbance band of the fluorinated alkene in the 5330 Pa

spectra of CF3I tested in the moist condition with copper

Integrated area from 944 to 953 cm ' (±0.011) Week number

Metal Temp.

(°C) 0 4 8 16 20 24 28 32 36 CDA 110 Blank

100 ND' ND 0.01 0.01 0.02 0.02 0.02 0.02 0.03

CDA 110 Blank150 ND ND ND ND ND ND ND ND ND

CDA 110 N40150 ND 0,02 0.04 0.11 0.12 0.12 0.12 0.12 0.12

CDA 110

Ti

150 ND 0.04 0.07 0.04 0.03 0.03 0.03 0.03 0.03

CDA 110 C4130150 ND 0.12 0.18 0.22 0.23 0.17 NM" 0.18 0.17

CDA 110

1625

150 ND 0.11 0.04 0.06 0.04 0.03 NM 0.03 0.03

Table 33 (cont.). Integrated peak areas of an absorbance band of the fluorinated alkene in the 5330

Pa spectra of CF3I tested in the moist condition with copper

Integrated area from 944 to 953 cm"' (±0.011)

Week number

Metal Temp.

(°C) 40 44 48 52 CDA 110 Blank

100 0.03 0.03 NM NM

CDA 110 Blank150 ND ND ND ND

CDA 110 N40150 0.12 0.12 0.12 0.12

CDA 110

Ti

150 0.03 0.03 0.03 0.03

CDA 110 C4130

150 0.18 NM NM NM

CDA 110

1625

150 0.04 NM NM NM

Not detected Not measured.
7. AGENT STABILITY UNDER STORAGE 335

Table 34. Summary of the initial and final peak areas of CF3I at all temperatures and conditions

Temp.

(°C)

Area (± 0.011)

Significant @

Metal Number of weeks Init. Final 3 X uncertainty

Condition: Dry without Copper

blank 48 23 0.25 0.23 no N40 48 23 0.19^ 0 24

Ti 48 23 0.26 0.23 no C4130 40 23 0.24 0.25 no

1625 44 23 0.27 0.24 no

blank 44 100 0.25 0.25 no N40 32 100 0.24 0 24

Ti 44 100 0.25 0.24 no C4130 40 100 0.25 0.24 no

1625 40 100 0.25 0.24 no

blank 52 150 0.28 0.25 no N40 52 150 0.27 0.25 no

Ti 52 150 0.28 0.25 no C4130 40 150 0.29' 0.23

1625 44 150 0.25 0.22 no

Condition: Dry with Copper

blank 48 23 0.26 0.23 no N40 48 23 0.26 0.24 no

Ti 48 23 0.27 0.25 no C4130 40 23 0.23 0.24 no

1625 44 23 0.26 0.23 no

blank 32 100 0.25 0.26 no

blank 52 150 0.29 0.23 yes

N40 52 150 0.26 0.19 yes

Ti 52 150 0.28 0.23 yes

C4130 40 150 0.24 0.22 no

1625 44 150 0.25 0.23 no
336 7. AGENT STABILITY UNDER STORAGE

Table 34. (continued) Summary of the initial and final peak areas of CF3I at all temperatures and

conditions

Metal Number of weeks Temp. °C Area (± 0.011) Significant @

Init. Final

3 X uncertainty

Condition: Moist without copper

blank 44 100 0.25 0.24 no

blank 52 150 0.27 0.23 yes

N40 52 150 0.28 0.23 yes

Ti 52 150 0.28 0.23 yes

C4130 40 150 0.24 0.21 no

1625 44 150 0.25 0.21 yes

Condition: Moist with copper

blank 44 100 0.25 0.24 no

blank 52 150 0.27 0.22 yes

N40 52 150 0.27 0.19 yes

Ti 52 150 0.28 0.21 yes

C4130 40 150 0.23 0.21 no

1625 40 150 0.23 0.21 no

Suspected to be an outlier because of improper gas cell filling.

Table 35. Spectral comparisons of initial and aged CF3I samples

Condition Without copper (± 0.0018) With copper (± 0.0018)

Metal Weeks Dry at Dry at Moist at Dry at Dry at Moist at tested 23 °C 150 °C 150 °C 23 °C 150 °C 150 °C

Blank 52 0.996 0.991 0.984 0.997 0.980 0.981

N40 52 0.998 0.995 0.992 0.997 0.875 0.949

Ti 52 0.998 0.996 0.990 0.997 0.985 0.990

C4130 40 0.999 0.994 0.990 0.999 0.997 0.998

1625 44 0.998 0.988 0.983 0.999 0.994 0.994
AGENT STABILITY UNDER STORAGE

0)c<

O)ac

'i_o

13

(D

C

o

m

(Jco

Xlo

CO <

cD

O

D

a;

o

CL

cn

0)

0.0

0.08 -

0.04

0.00 c>

All samples at 150 °C

0 4 8 12 16 20 24 28 32 36 40 44 48 52

o Blank

V N40

Ti A C4130

o 1625

Blank at 100 °C

0 4 8 12 16 20 24 28 32 56 40 44 48 52

Time (weeks)

Figure 38. Integrated peak areas for the fluorinated alkene for all samples of CF3I tested in moist condition with copper at 100 °C and 150 °C plotted as a function of time.
338 7. AGENT STABILITY UNDER STORAGE

7.4 Conclusions

The long-term stability project showed the following:

• the fiuorocarbon agents FC-218, HFC- 125, and HFC-227ea were stable at temperatures as high

as 150 °C for as long as 48 weeks. No by-products were formed.

• CF3I degraded at 100 °C and was accelerated at 150 °C.

• CF3H, CO2, and CO were produced in low levels as degradation products of CF3I.

• the presence of moisture accelerates the degradation of CF3I.

• the presence of copper accelerates the degradation of CF3I.

• the presence of copper and moisture accelerate the degradation of CF3I.

• an absorbance band at 950 cm ' was generated in the CF3I samples that may be from a

fluorinated alkene; the presence of copper at 150 °C caused the double bond to break.

• storage at ambient conditions of any of the four agents is feasible, but storage at elevated

temperatures for CF3I needs more study.

7.5 Acknowledgments

The author wishes to acknowledge a number of additional NIST staff members who made contributions to this section. These include Mr. Richard Peacock, Dr. Marc Nyden, Mr. Thomas Cleary,

Dr. Takashi Kashiwagi, Dr. Richard Gann, and Mr. Darren Lowe for providing their technical

expertise. Also, Ms. Paula Garrett for organizing and editing the section.

In addition the author wishes to acknowledge Mr. Richard Sears of Walter Kidde Aerospace for

coordinating the treatment of the C4130 alloy steel and Mr. Larry Braden of Metal Samples Co., Inc. for assistance in providing the numerous metal coupons used. Dr. Douglas Dierdorf of Pacific

Scientific helped to try and identify and confirm the presence of a fluorinated alkene in aged CF3I

samples. Mr. Thomas Austin of Washington Valve and Fitting Co. recommended equipment that made the FTIR analyses procedure very precise.

7.6 References

Balzhiser, R.E., Samuels, M.R., and Eliassen, J.D., Chemical Engineering Thermodynamics, The

Study of Energy, Entropy, and Equilibrium, Prentice-Hall, Englewood Cliffs, NJ, 1972.

Dierdorf, D., personal communication, 1995.

Felder, P., "The influence of the molecular beam temperature on the photodissociation of CF3I at 308 nm," Chem. Phys. Lett. 197, 425 (1992).

Fourier Infrared Software Tools for Microsoft Windows, User's Reference Guide, Mattson

Instruments, Inc., 1992.

McGee, R R., Cleveland, F. R, Meister, A. G., Decker, C. E., and Miller, S. I., "Substituted

Methanes.X. Infrared Spectral Data, Assignments, Potential Constants, and Calculated Thermodynamic

Properties for CFjBr and CF3I," J. of Chem. Phys., 21, 2, 1948.
7. AGENT STABILITY UNDER STORAGE 339

Mendenhall, W. and Sincich, T., Statistics for Engineering and the Sciences, Dellen Publishing

Co., p. 213, 214, 1992.

Peacock, R.D., Cleary, T.G., and Harris, Jr., R.H., "Agent Stability under Storage and Discharge

Residue," Section 6 in Evaluation of Alternative In-Flight Fire Suppressants for Full-Scale Testing in Simulated Aircraft Engine Nacelles and Dry Bays, NIST Special Publication 861, Grosshandler, W. L., Gann, R. G., and Pitts, W. M., eds, U.S. Department of Commerce, Washington, DC 1994. 643-668.

Ricker, R.E., Stoudt, M.R., Dante, J.F., Fink, J.L., Beauchamp, C.R., and Moffat, T.P, "Corrosion

of Metals," Section 7 in Evaluation of Alternative In-Flight Fire Suppressants for Full-Scale Testing in Simulated Aircraft Engine Nacelles and Dry Bays, NIST Special Publication 861, Grosshandler, W. L., Gann, R. G., and Pitts, W. M., eds, U.S. Department of Commerce, Washington, DC 1994. 643-668.

Taylor, B.N., Kuyatt, C.E., Guidelines for Evaluating and Expressing the Uncertainty of NIST

Measurement Results, NIST Technical Note 1297, U.S. Department of Commerce, Washington DC.

1994.
340 7. AGENT STABILITY UNDER STORAGE

Appendix A. Initial and Final FTIR Spectra for HFC-125

The spectra in Appendix A are those of the HFC-125. The gas cell pressure for all was 5330 Pa. The

lower spectrum is for the initial analysis; the upper spectrum is for the final aged analysis. The blanks

for this agent appear in Figures 13 and 14.
7. AGENT STABILITY UNDER STORAGE 341

Figure A-1. Initial (lower) and 48 week (upper) spectra for Nitronic 40 in HFC- 125 at 23 °C.
342 7. AGENT STABILITY UNDER STORAGE

Figure A-2. Initial (lower) and 48 week (upper) spectra for Ti- 15-3-3-3 in HFC- 125 at 23 °C.
7. AGENT STABILITY UNDER STORAGE 343

Figure A-3. Initial (lower) and 32 week (upper) spectra for C4130 in HFC- 125 at 23 °C.
344 7. AGENT STABILITY UNDER STORAGE

Figure A-4. Initial (lower) and 48 week (upper) spectra for 1625 in HFC- 125 at 23 °C.
7. AGENT STABILITY UNDER STORAGE 345

Figure A-5. Initial (lower) and 48 week (upper) spectra for nitronic 40 in HFC- 125 at 150 °C.
7. AGENT STABILITY UNDER STORAGE

Figure A-6. Initial (lower) and 48 week (upper) spectra for Ti- 15-3-3-3 in HFC- 125 at 150 °C.
7. AGENT STABILITY UNDER STORAGE 347

Figure A-7. Initial (lower) and 40 week (upper) spectra for C4130 in HFC- 125 at 150 °C.
348 7. AGENT STABILITY UNDER STORAGE

Figure A-8. Initial (lower) and 48 week (upper) spectra for 1625 in HFC- 125 at 150 °C.
7. AGENT STABILITY UNDER STORAGE 349

Appendix B. Initial and Final FTIR Spectra of HFC-227ea

The spectra in Appendix B are those of the HFC-227ea. The gas cell pressure for all was 5330 Pa. The lower spectrum is for the initial analysis; the upper spectrum is for the final aged analysis. The

blanks for this agent appear in Figures 17-19.
350 7. AGENT STABILITY UNDER STORAGE

Figure B-1. Initial (lower) and 48 week (upper) spectra for nitronic 40 in HFC-227ea at 23 °C.
7. AGENT STABILITY UNDER STORAGE 351

3500 3000 2500 2000 1500

Wavenumbers (cm-1

)

Figure B-2. Initial (lower) and 40 week (upper) spectra for Ti- 15-3-3-3 in HFC-227ea at 23 °C.
352 7. AGENT STABILITY UNDER STORAGE

Figure B-3. Initial (lower) and 32 week (upper) spectra for C4130 in HFC-227ea at 23 °C.
7. AGENT STABILITY UNDER STORAGE 353

Figure B-4. Initial (lower) and 48 week (upper) spectra for 1625 in HFC-227ea at 23 °C.
354 7. AGENT STABILITY UNDER STORAGE

3500 3000 2500 2000 1500 1000

Wavenumbers (cm-1

)

Figure B-5. Initial (lower) and 40 week (upper) spectra for nitronic 40 in HFC-227ea at 125 °C.
7. AGENT STABILITY UNDER STORAGE 355

Figure B-6. Initial (lower) and 40 week (upper) spectra for Ti- 15-3-3-3 in HFC-227ea at 125 °C.
356 7. AGENT STABILITY UNDER STORAGE

Figure B-7. Initial (lower) and 40 week (upper) spectra for C4130 in HFC-227ea at 125 °C.
7. AGENT STABILITY UNDER STORAGE 357

Figure B-8. Initial (lower) and 40 week (upper) spectra for 1625 in HFC-227ea at 125 °C.
358 7. AGENT STABILITY UNDER STORAGE

Figure B-9. Initial (lower) and 48 week (upper) spectra for nitronic 40 in HFC-227ea at 150 °C.
7. AGENT STABILITY UNDER STORAGE 359

3500 3000 2500 2000 1500 1000

Wavenumbers (cm-1

)

Figure B-10. Initial (lower) and 48 week (upper) spectra for Ti- 15-3-3-3 in HFC-227ea at 150 °C.
360

7. AGENT STABILITY UNDER STORAGE

3500 3000 2500 2000 1500 1000

Wavenumbers (cm-1

)

Figure B-U. Initial (lower) and 40 week (upper) spectra for C4130 in HFC-227ea at 150 °C.
7. AGENT STABILITY UNDER STORAGE

Figure B-12. Initial (lower) and 48 week (upper) spectra for 1625 in HFC-227ea at 150 °C.
362 7. AGENT STABILITY UNDER STORAGE

Appendix C. Initial and Final FTIR Spectra of CF3I Tested in the Dry

Condition without Copper

The spectra in Appendix C are those of the CF3I. The gas cell pressure for all was 5330 Pa. The

lower spectrum is for the initial analysis; the upper spectrum is for the final aged analysis.
7. AGENT STABILITY UNDER STORAGE 363

Figure C-1. Initial (lower) and 48 week (upper) spectra for the blank in CF3I tested in the dry

condition without copper at 23 °C.
364 7. AGENT STABILITY UNDER STORAGE

Figure C-2. Initial (lower) and 48 week (upper) spectra for nitronic 40 in CF3I tested in the dry

condition without copper at 23 °C.
7. AGENT STABILITY UNDER STORAGE 365

Figure C-3. Initial (lower) and 48 week (upper) spectra for Ti- 15-3-3-3 in CF3I tested in the dry

condition without copper at 23 °C. . .
366 7. AGENT STABILITY UNDER STORAGE

3500 3000 2500 2000 1500 1000

Wavenumbers (cm-1

)

Figure C-4. Initial (lower) and 40 week (upper) spectra for C4130 in CF3I tested in the dry condition

without copper at 23 °C.
7. AGENT STABILITY UNDER STORAGE 367

Figure C-5. Initial (lower) and 44 week (upper) spectra for 1625 in CF3I tested in the dry condition

without copper at 23 °C. . _
368 7. AGENT STABILITY UNDER STORAGE

Figure C-6. Initial (lower) and 44 week (upper) spectra for the blank in CF3I tested in the dry

condition without copper at 100 °C.
7. AGENT STABILITY UNDER STORAGE 369

Figure C-7. Initial (lower) and 32 week (upper) spectra for nitronic 40 in CF3I tested in the dry

condition without copper at 100 °C.
370 7. AGENT STABILITY UNDER STORAGE

Figure C-8. Initial (lower) and 44 week (upper) spectra for Ti- 15-3-3-3 in CF3I tested in the dry

condition without copper at 100 °C.
7. AGENT STABILITY UNDER STORAGE 371

3500 3000 2500 2000 1500 1000

Wavenumbers (cm-1

)

{ Figure C-9. Initial (lower) and 40 week (upper) spectra for C4130 in CF3I tested in the dry condition

without copper at 100 °C. = ,
372 7. AGENT STABILITY UNDER STORAGE

Figure C-10. Initial (lower) and 40 week (upper) spectra for 1625 in CF3I tested in the dry

condition without copper at 100 °C.
7. AGENT STABILITY UNDER STORAGE 373

Figure C-11. Initial (lower) and 52 week (upper) spectra for the blank in CF3I tested in the dry

condition without copper at 150 °C.

I
374 7. AGENT STABILITY UNDER STORAGE

Figure C-12. Initial (lower) and 52 week (upper) spectra for nitronic 40 in CF3I tested in the dry

condition without copper at 150 °C.
7. AGENT STABILITY UNDER STORAGE 375

Figure C-13. Initial (lower) and 52 week (upper) spectra for Ti-15-3-3-3 in CF3I tested in the dry

condition without copper at 150 °C. : . .
376 7. AGENT STABILITY UNDER STORAGE

Figure C-14. Initial (lower) and 40 week (upper) spectra for C4130 in CF3I tested in the dry

condition without copper at 150 °C.
7. AGENT STABILITY UNDER STORAGE 377

Figure C-15. Initial (lower) and 44 week (upper) spectra for 1625 in CF3I tested in the dry

condition without copper at 150 °C.
378 7. AGENT STABILITY UNDER STORAGE

Appendix D. Initial and Final FTIR Spectra of CF3I Tested in the Dry

Condition with Copper

The spectra in Appendix D are those of the CF3I. The gas cell pressure for all was 5330 Pa. The

lower spectrum is for the initial analysis; the upper spectrum is for the final aged analysis.
7. AGENT STABILITY UNDER STORAGE 379

Figure D-1. Initial (lower) and 48 week (upper) spectra for the blank in CF3I tested in the dry

condition with copper at 23 °C.
380 7. AGENT STABILITY UNDER STORAGE

Figure D-2. Initial (lower) and 48 week (upper) spectra for nitronic 40 in CF3I tested in the dry

condition with copper at 23 °C.
7. AGENT STABILITY UNDER STORAGE 381

Figure D-3. Initial (lower) and 48 week (upper) spectra for Ti-15-3-3-3 in CF3I tested in the dry

condition with copper at 23 °C.
7. AGENT STABILITY UNDER STORAGE

Figure D-4. Initial (lower) and 40 week (upper) spectra for C4130 in CF3I tested in the dry

condition with copper at 23 °C.
7. AGENT STABILITY UNDER STORAGE 383

r

^

J;

1 a.

aI L/U

3500 3000 2500 2000 1500 1000

Wavenumbers (cm-1

)

Figure D-5. Initial (lower) and 44 week (upper) spectra for 1625 in CF3I tested in the dry

condition with copper at 23 °C.
384 7. AGENT STABILITY UNDER STORAGE

3500 3000 2500 2000 1500 1000

Wavenumbers (cm-1

)

Figure D-6. Initial (lower) and 32 week (upper) spectra for the blank in CF3I tested in the dry

condition with copper at 100 °C.
7. AGENT STABILITY UNDER STORAGE 385

Figure D-7, Initial (lower) and 52 week (upper) spectra for the blank in CF3I tested in the dry

condition with copper at 150 °C.
386 7. AGENT STABILITY UNDER STORAGE

Figure D-8. Initial (lower) and 52 week (upper) spectra for nitronic 40 in CF3I tested in the dry

condition with copper at 150 °C.
7. AGENT STABILITY UNDER STORAGE 387

3500 3000 2500 2000 1500 1000

Wavenumbers (cm-1

)

Figure D-9. Initial (lower) and 52 week (upper) spectra for Ti- 15-3-3-3 in CF3I tested in the dry

condition with copper at 150 °C.
388 7. AGENT STABILITY UNDER STORAGE

1

u

]

/I

4000 3500 3000 2500 2000 1500 1000

Wavenumbers (cm-l

)

Figure D-10. Initial (lower) and 40 week (upper) spectra for C4130 in CF3I tested in the dry

condition with copper at 150 °C.
AGENT STABILITY UNDER STORAGE 389

Figure D-11. Initial (lower) and 44 week (upper) spectra for 1625 in CF3I tested in the dry

condition with copper at 150 °C.
390 7. AGENT STABILITY UNDER STORAGE

Appendix E, Initial and Final FTIR Spectra of CF3I Tested in the Moist

Condition without Copper

The spectra in Appendix E are those of the CF3I. The gas cell pressure for all was 5330 Pa. The

lower spectrum is for the initial analysis; the upper spectrum is for the final aged analysis.
7. AGENT STABILITY UNDER STORAGE 391

Figure E-1. Initial (lower) and 44 week (upper) spectra for the blank in CF3I tested in the moist

condition without copper at 100 °C.
392 7. AGENT STABILITY UNDER STORAGE

Figure E-2. Initial (lower) and 52 week (upper) spectra for the blank in CF3I tested in the moist

condition without copper at 150 °C.
7. AGENT STABILITY UNDER STORAGE 393

Figure E-3. Initial (lower) and 52 week (upper) spectra for nitronic 40 in CF3I tested in the moist

condition without copper at 150 °C.
394 7. AGENT STABILITY UNDER STORAGE

Figure E-4. Initial (lower) and 52 week (upper) spectra for Ti- 15-3-3-3 in CF3I tested in the moist

condition without copper at 150 °C.
7. AGENT STABILITY UNDER STORAGE 395

Figure E-5. Initial (lower) and 40 week (upper) spectra for C4130 in CF3I tested in the moist condition

without copper at 150 °C.
396 7. AGENT STABILITY UNDER STORAGE

Figure E-6. Initial (lower) and 44 week (upper) spectra for 1625 in CF3I tested in the moist condition

without copper at 150 °C.
7. AGENT STABILITY UNDER STORAGE

Appendix F. Initial and final FTIR Spectra of CF3I Tested in the Moist

Condition with Copper

The spectra in Appendix F are those of the CF3I. The gas cell pressure for all was 5330 Pa. The

lower spectrum is for the initial analysis; the upper spectrum is for the final aged analysis.
398 7. AGENT STABILITY UNDER STORAGE

3500 3000 2500 2000 1500 1000

Wavenumbers (cm-1

)

Figure F-1. Initial (lower) and 44 week (upper) spectra for the blank in CF3I tested in the moist

condition with copper at 100 °C.
7. AGENT STABILITY UNDER STORAGE 399

3500 3000 2500 2000 1500 1000

Wavenumbers (cm-1

)

Figure F-2. Initial (lower) and 52 week (upper) spectra for the blank in CF3I tested in the moist

condition with copper at 150 °C.
400 7. AGENT STABILITY UNDER STORAGE

Figure F-3. Initial (lower) and 52 week (upper) spectra for nitronic 40 in CF3I tested in the moist

condition with copper at 150 °C.
7. AGENT STABILITY UNDER STORAGE 401

Figure F-4. Initial (lower) and 52 week (upper) spectra for Ti- 15-3-3-3 in CF3I tested in the moist

condition with copper at 150 °C.
402 7. AGENT STABILITY UNDER STORAGE

Figure F-5. Initial (lower) and 40 week (upper) spectra for C4130 in CF3I tested in the moist condition

with copper at 150 °C.
7. AGENT STABILITY UNDER STORAGE 403

Figure F-6. Initial (lower) and 40 week (upper) spectra for 1625 in CF3I tested in the moist condition

with copper at 150 °C. -

PART 2

ENGINEERING DESIGN CRITERIA
1
407

8. OPTIMIZATION OF SYSTEM DISCHARGE

Jiann C. Yang, Thomas G. Cleary, Isaura Vdzquez, Charles I. Boyer, Michelle D. King,

Brett D. Breuel, Carole Womeldorf, and William L. Grosshandler

Building and Fire Research Laboratory

Marcia L. Ruber and Lloyd Weber

Chemical Science and Technology Laboratory

Grzegorz Gmurczyk

Science Applications International Corporation

Contents

Page

8. OPTIMIZATION OF SYSTEM DISCHARGE 407

8.1 Overview 410

8.2 System Constraints 410

8.3 Thermophysical Properties of Selected Agents and Agent/Nitrogen Mixtures 413

8.3.1 Introduction 413

8.3.2 Experimental Apparatus and Procedures 414

8.3.2.1 Agent FiUing Calculations 414

8.3.2.2 Agent Preparation 416

8.3.2.3 Pure Agent Measurements 416

8.3.2.4 Nitrogen Solubility in Agents 421

8.3.2.5 Agent/Nitrogen Mixture Measurements 421

8.3.3 Modeling Thermodynamic Properties of Agent/Nitrogen Mixtures 423

8.3.3.1 Introduction 423

8.3.3.2 Extended Corresponding States Model 423

8.3.3.3 Calculation Procedure 427

8.3.4 Results and Discussion 427

8.3.4.1 Pure Agents 427

8.3.4.2 Solubility of Nitrogen in Agents 429

8.3.4.3 Agent/Nitrogen Mixtures 432

8.3.5 Summary and Conclusions 438

8.4 Discharge of Agent/Nitrogen Mixtures in a Simulated Dry Bay 438

8.4.1 Introduction 438

8.4.2 Experimental Methods 439

8.4.3 Results and Discussion 447

8.4.3.1 General Visual Observations 447
408 8. OPTIMIZATION OF SYSTEM DISCHARGE

8.4.3.2 Marotta Solenoid Valve 448

8.4.3.3 HTL Solenoid Valve 467

8.4.3.4 Kidde-Graviner Valve 468

8.4.3.5 HTL Squibs 471

8.4.3.6 Effect of an Orifice Plate 471

8.4.3.7 Effect of Fill Density 476

8.4.3.8 Effect of Initial Nitrogen Charge Pressure 476

8.4.3.9 Effect of Initial Vessel Temperature 494

8.4.3.10 Effect of Vessel Orientation 500

8.4.3.11 Effect of Partial Nitrogen Saturation 502

8.4.3.12 Effect of Vessel Geometry 509

8.4.3.13 Model Predictions 509

8.4.4 Summary and Conclusions 531

8.5 Computer Simulation of Agent Dispersion 531

8.5.1 Introduction 531

8.5.2 Modeling of Agent Dispersion Process 532

8.5.3 Solution Method 533

8.5.4 Numerical Results and Discussion 533

8.5.5 Conclusions 562

8.6 Pipe Flow Characteristics of Alternative Agents 563

8.6.1 Introduction 563

8.6.2 Agent Discharge in Piping 563

8.6.3 Single-Phase Flow 564

8.6.4 Two-Phase Flow 565

8.6.5 Experimental Apparatus 566

8.6.5.1 Discharge Vessel and Piping Design 566

8.6.5.2 Experimental Measurements 568

8.6.5.3 Test Matrix 572

8.6.6 Results and Discussion 574

8.6.6.1 Constant-Head Discharge 576

8.6.6.2 Transient Discharge 579

8.6.6.3 Straight Pipe Configuration 583

8.6.6.4 Piping with a 90° Angle Bend 584

8.6.6.5 Straight Pipe with a Sudden Expansion 584

8.6.6.6 Straight Pipe with a Sudden Contraction 585

8.6.6.7 Tee Configuration 588

8.6.6.8 Cooled Agent Discharge 588

8.6.6.9 Heated Agent Discharge 588

8.6.6.10 Visual Observations 596

8.6.7 Transient Discharge Model 596

8.6.7.1 Model Equations 598

8.6.7.2 Fill Conditions and Thermodynamic Inputs 604

8.6.7.3 Friction Factor 604

8.6.8 Model Calculations 613

8.6.8.1 Constant-Head Conditions 613

8.6.8.2 Transient Conditions 613

8.6.9 Hardware Design Scheme 627

8.6.10 Conclusions and Recommendations 639
8. OPTIMIZATION OF SYSTEM DISCHARGE 409

8.7 Concluding Remarks 643

8.8 Acknowledgments 645

8.9 References 645

Appendix A. A User's Guide to PROFISSY 649

Appendix B. Tabulated fill Conditions and Pressure Traces for All Pipe Flow Discharge

Tests 668

Appendix C. Thermodynamic and Transport Properties of Agent/Nitrogen Mixtures 773
410 8. OPTIMIZATION OF SYSTEM DISCHARGE

8.1 Overview

Current aircraft fire suppression bottles for dry bay and engine nacelle applications, which are designed to meet Military Specification MIL-C-22284A (proof pressure of 9.62 MPa and minimum

burst pressure of 12.37 MPa), are normally filled with liquid halon 1301 (CFjBr) to about half of the

bottle volume, and the bottle is then pressurized with nitrogen to a specified equilibrium pressure

(typically 4.1 MPa) at room temperature. The purpose of using the pressurization gas is to expedite

the discharge of the agent and to facilitate the dispersion of the agent. Without nitrogen pressurization, the bottle pressure, which is simply the vapor pressure of the agent, can be so low (even subatmospheric) at extremely cold ambience that there is virtually no driving force to expel the agent

from the bottle in case of a fire, thus hindering a rapid release of the mixture.

From the above description, three important issues have emerged and need be considered when

using a halon alternative as an in-flight fire suppressant: (1) the system hardware, (2) the thermophysical properties of the agent/nitrogen mixture, and (3) the agent/nitrogen mixture behavior during a

discharge.

In this section, the last two issues which are pertinent to agent storage and its subsequent

discharge will be addressed in detail. With respect to the first issue, a survey of the system constraints on existing in-flight halon 1301 fire protection systems will be given. This survey is also used to provide guidelines for identifying important experimental parameters for subsequent agent/nitrogen

storage and discharge studies. The information on system constraints consists of bottle size, percent

liquid-fill, nitrogen pressurization, range of operating temperatures, and proof and burst pressures of

the vessel. In addition, the information on piping systems used to distribute halon 1301 for engine

nacelle fire protection applications will be discussed. The thermophysical properties of selected agents

and agent/nitrogen mixtures are presented under conditions commensurate with those identified in the

system constraint studies. This information is also relevant to the establishment of the initial conditions of the vessel before the discharge of the mixture. Finally, parametric studies on the

discharges of agent/nitrogen mixtures without (to simulate dry bay applications) and with (to simulate

engine nacelle applications) a piping system will be discussed. Numerical calculations using

CONCHAS-SPRAY and KIVA II were also performed in order to explore the feasibility of applying

such complex codes to the study of a flashing discharge into an enclosure. This section focuses on the

four selected agents (HFC-125, FC-218, CF3I, and HFC-227ea) for dry bay (HFC-125, FC-218, and

CF3I) and engine nacelle (HFC-125, CF3I, and HFC-227ea) applications.

The results obtained from this study provide important technical information on bottle design and

agent discharge for new generation aircraft that may use the halon alternatives, for the existing aircraft that may undergo retrofitting, or simply for possible "drop-in" replacements.

8.2 System Constraints

This section describes a survey summarizing the current in-flight halon 1301 systems. This work was

conducted with the cooperation of several aircraft fire extinguisher manufacturers, aircraft and airframe

manufacturers, defense contractors, the three branches of the U.S. Military Service, and the Federal

Aviation Administration (FAA). This survey is the result of examining the halon 1301 system

specifications for over fifty types of aircraft (both military and commercial). The halon 1301 systems

included in this survey are used for fire protection in engine nacelles, dry bays, and auxiliary power

units (APU). Table 1 lists the participants in the survey and the acquired information. Since some of

the information obtained was considered proprietary, it was not possible to tabulate the system
8. OPTIMIZATION OF SYSTEM DISCHARGE 411

Table 1. Names of the participants in the survey and the acquired information on current halon

1301 system constraints

Participant Acquired information

Boeing Piping

Duui-AUcn ol ndniiiiun, inc. Bottle sizes, agent fill volumes, operating temperaures,

vessel growth potential, nitrogen pressurization

rcucrai AvidLiun /Auriiini^uaiiun Bottle sizes and agent fill volumes

Lockheed Martin Piping, and bottle sizes

McDonnell Douglas Vessel growth potential

Northrop Grumman Vessel growth potential

Pacific Scientific Bottle sizes, agent fill volumes, operating temperatures, and

nitrogen pressurization

Systron-Donner Bottle sizes, agent fill volumes, operating temperatures, and

nitrogen pressurization

Walter Kidde Bottle sizes, agent fill volumes, operating temperatures, and

nitrogen pressurization

U.S. Air Force Bottle sizes, agent fill volumes, operating temperatures, and

nitrogen pressurization

U.S. Army Bottle sizes, agent fill volumes, operating temperatures, and

nitrogen pressurization

U.S. Navy Bottle sizes, agent fill volumes, operating temperatures, and

nitrogen pressurization

specifications for each individual aircraft in detail. Only the minimum and maximum operating ranges

of the parameters were compiled in this study. Note that no differentiation among the areas of

application were made in identifying the parameters. The survey is summarized in Table 2. The

percent of liquid fill was calculated based on the total amount and liquid density of pure agent at 25 °C. In most applications, the bottle specifications closely conform to Military Specification MIL-C22284A. The vessel growth potential is defined as the allowable percent increase in the original vessel

size. However, it is not clear from the literature whether the growth potential refers to the increase in vessel size at the same mounting location of the vessel or different, new location when rerouting of

agent distribution is feasible. Based on the volume factor, it is not possible, in most cases, to use the existing containers for the

selected halon alternatives as "drop-in" replacements because the containers may not be large enough

to hold the amount of agent required for fire extinguishment or may encounter the so-called liquid-full

condition (Grosshandler et ai, 1994) sooner even at moderately high ambient temperatures due to an

initial small ullage as a result of an increased initial fill volume of the liquid halon alternative. The

volume factor is defined as a measure of the liquid agent volume in the storage vessel normalized by

the liquid volume of halon 1301 required to suppress a flame under identical conditions (Grosshandler

et al., 1994). The following calculations, using HFC-125 as an example, illustrate how the system

constraints on the bottle size can be assessed based only on thermodynamic considerations. Note that
412 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 2. Summary of current halon 1301 system specifications for engine nacelle, dry bay, and

auxiliary power unit applications

Property Range

Bottle size (cm"^) 650 to 23000

Percent liquid fill (%) 25 to 70 .

•

Operating temperature (°C) - 54 to 160

Nitrogen pressurization (MPa) 1.4 to 5.7

Proof pressure (MPa) 6.4 to 24.7

Burst pressure (MPa) 11.3 to 37.1

Vessel growth potential (%) 0 to 400

Other constraints (e.g., mechanical properties of the vessel, discharge time) have to be taken into

account in the overall system constraint evaluation. The intent here is not to provide a comprehensive

methodology for system constraint evaluation, but rather to provide system designers a set of

engineering tools to perform design calculations (in this case, on bottle size).

Assuming the original halon 1301 bottle is a military standard CF-6 type vessel with an internal

volume of 3.67 x 10"^ m'^ and contains 2.96 kg of halon 1301 pressurized to 4.12 MPa with nitrogen

at 21 °C, the percent of liquid fill at this condition is calculated to be 49.7 % using the computer code

PROFISSY (to be discussed in detail in Section 8.3). The volume of liquid HFC- 125 required is 5.47 X 10"^ m^, based on an average volume factor of 3 as reported in Grosshandler et al. (1994). If there is no vessel growth potential, then the original halon 1301 vessel can not be used for HFC- 125

because the liquid volume of the agent is greater than the vessel volume. Table 3 summarizes the

system constraint evaluation based on other vessel growth potentials. It is clear that even with a growth potential of 500 %, which results in 24.9 % liquid fill, the pressure at 150 °C is still very close

to the burst pressure of a standard CF-6 vessel. In the calculations, a temperature of 150 °C was

assumed to be the extreme operating temperature of the vessel. Similar calculations were also

performed for the other three agents, and the results are given in Tables 4, 5, and 6. To the best of our knowledge, no military specifications exist for storage and delivery systems for

dry bay discharge applications. Since the events typically occurring in a dry bay are on the order of

milliseconds, it is expected that the discharge will be directly from the vessel without going through

any piping systems. For engine nacelle applications, piping systems are required to transport halon

from the bottles (normally located outside the engine nacelle) to the point of application in the nacelle. The design of the piping system centers around Military Specification MIL-E-22285 (1959) and its amendment (1960). According to these documents, the discharge time, defined as the time measured

from the instant the agent (halon 1301) starts to leave the tubing ends to the time the required amount

of agent has been discharged, should be 1 second or less. The initial amount of agent in the

container(s) should be able to provide an agent concentration of at least 6 % by volume in all parts of

the affected zone for at least 0.5 second at normal cruising condition. The pipe lines should be less than 3.1 m (10 ft) long, where practicable, and be straight if possible. The fittings and turns in the

piping system should be kept to a minimum. Discharge piping should terminate in open ends without
8. OPTIMIZATION OF SYSTEM DISCHARGE 413

Table 3. Sample calculations showing vessel size constraint for HFC- 125 based on an average

volume factor of 3 and various vessel growth potentials

Vessel growth potential (%)

New vessel volume Percent liquid fill (%)

Vessel pressure @ 150 °C (MPa)

0 3.67 X 10"^ Not possible Not calculated

50 5.51 X IQ-^ 99.4 > 41.23

100 7.34 X IQ-^ 74.6 26.65

200 11.01 X 10"^ 49.7 16.77

300 14.68 X IQ-^ 37.3 14.01

400 18.35 X IQ-^ 29.8 12.61

500 22.02 X 10"^ 24.9 11.73

nozzles, and tubing and fittings should have a minimum burst pressure of 10.3 MPa. Tubing sizes are determined based on the desired discharge rates and on system-volume considerations {i.e., minimizing

the system volumes). It is likely that any system employing a halon replacement agent would, at least,

have to meet similar performance criteria and system constraints.

8.3 Thermophysical Properties of Selected Agents and Agent/Nitrogen

Mixtures

8.3.1 Introduction. The total pressure in the bottle is a complex function of ambient temperature

because of the temperature dependence of the agent vapor pressure, the partial pressure of nitrogen in the gas ullage, and the solubility of nitrogen in the liquid agent. For halon 1301, the bottle pressure- temperature relationship and the solubility of nitrogen in halon 1301 have been well characterized. By

contrast, such a relationship and solubility data are scarce or do not exist for some of the halon

alternatives.

The major objectives of this study are very similar to those reported in the previous work

(Grosshandler et al., 1994), the only difference being the initial liquid agent fill conditions at room

temperature and the number of agents used. The present study expands the temperature-pressure

database for the four selected agents (HFC- 125, FC-218, CF3I, and HFC-227ea) using two additional

initial fill conditions and an experimental apparatus that can accommodate much higher final pressures

at elevated temperatures than the earlier study.

The first subtask of this study was to determine the solubility of nitrogen in the four selected

agents at room temperature. Halon 1301 was also included in this study for the purposes of reference

and comparison with the alternative agents. Specifically, the objective was to measure the amount of

nitrogen that was needed to pressurize the bottle to a specified equilibrium pressure.

In the second subtask, the thermodynamic properties of the four selected agents together with

halon 1301 were measured. The objective was to determine the final pressure of the bottle when the

bottle, filled with either a pre-determined amount of pure agent or with agent pressurized with

nitrogen, was exposed to two extreme temperatures (- 60 °C and 150 °C). Experiments on pure
414 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 4. Sample calculations showing vessel size constraint for FC-218 based on an average

volume factor of 2.8 and various vessel growth potentials

Vessel growth potential (%)

New vessel volume

(m^)

Percent liquid fill (%)

Vessel pressure @ 150 °C (MPa)

0 3.67 X

10"^ Not possible Not calculated

50 5.51 X IQ-^ 92.8 30.81

100 7.34 X 10"^ 69.6 18.43

200 11.01 X 10"^ 46.4 12.88

300 14.68 X 10"^ 34.8 11.22

400 18.35 X 10"^ 27.9 10.38

500 22.02 X

10"^ 23.2 9.83

agents were performed in order to obtain pressure-temperature relationships for some of the agents

(especially CF3I) where documented data were not readily available.

Furthermore, in order to assimilate the results obtained from the two aforementioned subtasks into

useful engineering correlations that could be easily used by the bottle designers, the third subtask was

to develop a computer code that could facilitate such calculations.

For the purpose of reference, selected thermophysical properties of the four selected agents are tabulated in Table 7 where MW is the molecular weight, is the normal boiling point (at

0.101 MPa), T^. is the critical temperature, P^^^ is the saturation vapor pressure at 25 °C, is the

critical pressure, is the critical density, py is the saturated liquid density at 25 °C, C^j is the isobaric

liquid heat capacity at 7^, is the isobaric liquid heat capacity at 25 °C, and /y^ is the latent heat of

vaporization at T^. The references from which these data were taken are also given in the table.

8.3.2 Experimental Apparatus and Procedures. This section describes the experimental techniques

used to measure the amount of nitrogen required to pressurize the vessel initially filled with a fixed

amount of agent to a specified equilibrium pressure at room temperature and the pressure-temperature

relationship for pure agents and agent/nitrogen mixtures. The initial amounts of liquid agent dispensed

in the vessel before nitrogen pressurization corresponded to approximately half and two-thirds of the

vessel volume at room temperature. For agent/nitrogen mixtures, two initial equilibrium nitrogen

charge pressures (2.75 MPa and 4.12 MPa) at room temperature were used in this study. The amount

of nitrogen required to pressurize the vessel with agent to 2.75 MPa or 4.12 MPa, were measured only

at room temperature. To determine the final vessel pressure as a function of temperature, experiments were performed at - 60 °C, 22 °C, and 150 °C.

8.3.2.1 Agent Filling Calculations. This sub-section describes the method used to calculate the

total amount of agent needed such that the liquid volume fraction corresponds to the amount that is one-half or two-thirds of the vessel volume at room temperature. The calculation procedure is as follows.

Since two phases (liquid-vapor) exist in equilibrium for all the four selected agents at room

temperature, the total mass (m^ j) of agent in the vessel with a total volume is equal to the sum of

the masses of agent in both vapor and liquid phases.
8. OPTIMIZATION OF SYSTEM DISCHARGE 415

Table 5. Sample calculations showing vessel size constraint for CF3I based on an average volume

factor of 0.8 and various vessel growth potentials

Vessel growth potential (%)

New vessel volume Percent liquid fill (%)

Vessel pressure @ 150 °C (MPa)

0 3.67 X IQ-^ 39.8 11.53

50 5.51 X 10"^ 26.5 10.46

100 7.34 X IQ-^ 19.9 9.85

200 11.01 X

10"^ 13.3 9.04

300 14.68 X

10"^ 10.0 8.52

400 18.35 X

10"^ 8.0 8.18

500 22.02 X

10"^ 6.6 7.93

m = m + m , = 0 V + o ,V,

where m is the mass, p is the saturation density of the agent, V is the volume, subscripts t, a, v, and / represents total, agent, vapor phase, and liquid phase respectively. For a given agent at room

temperature, ^ and p^ ^ are known or can be estimated. If the vessel is half filled with liquid agent.

Similarly, for a two-thirds liquid full vessel,

27,

' 3

" 3

(3)

(4)

With V; and calculated by the method shown above, m^, can then be determined from Equation (1). In the calculations for HFC-227ea, FC-218, HFC- 125, and'halon 1301, p^^ and p^ / were obtained

from the database developed by Gallagher et al. (1993). For CF3I, p^^ was estimated by the modified

Rackett method (Reid et al, 1987), and p^ ^ was calculated by using the vapor pressure at room

temperature and assuming an ideal vapor phase.
416 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 6. Sample calculations showing vessel size constraint for HFC-227ea based on an average

volume factor of 2.5 and various vessel growth potentials

Vessel growth potential (%)

New vessel volume

(m^)

Percent liquid fill (%)

Vessel pressure @ 150 °C (MPa)

0 3.67 X

10"^ Not possible Not calculated

50 5.51 X

10"^ 82.9 26.65

100 7.34 X 10"^ 62.2 15.74

200 11.01 X 10"^ 41.4 11.80

300 14.68 X 10"^ 31.1 10.71

400 18.35 X 10"^ 24.9 10.07

500 22.02 X 10"^ 20.7 9.62

8.3.2.2 Agent Preparation. Before the agent was dispensed into the test vessel, it was degassed

to remove any dissolved noncondensible gases which might be present in the agent. Figure 1 is a schematic of the degassing apparatus. The entire apparatus was first evacuated to 1.33 Pa for at least half an hour before agent vapor was condensed in the sublimizer which was chilled in a dry ice bath. The sublimizer was a stainless steel vessel with a re-entrant well at the center. After sufficient liquid

had been condensed, liquid nitrogen was poured into the well. The liquid agent in the sublimizer was

then made to boil by applying vacuum. As the liquid boiled, it would give off its own vapor and the

noncondensible gases. Because the wall of the well was very cold due to the presence of liquid

nitrogen, the agent vapor was condensed on the wall, and the condensate then dripped back to the

boiling liquid agent. In this way, only the noncondensible gases and a very small amount of agent

vapor were removed from the sublimizer by the vacuum pump. Depending on the initial amount of

liquid agent in the sublimizer, this degassing process normally went on for at least half an hour.

Because the initial batch of CF3I was found to contain trace amounts of HI, CO2, H2O, and I2, a

purification system was constructed in order to remove impurities which might be present in other

batches. The purification system is shown schematically in Figure 2. After the entire system was

evacuated, the agent vapor passed through a column packed with potassium hydroxide whose main

function was to remove HI, then through a molecular sieve filter to remove water vapor and carbon

dioxide, and finally through a stainless steel coil which was immersed in an ice bath to remove I2. The purified vapor was then retrieved by condensation in a stainless steel container. The collected

agent was again subjected to boiling using a vacuum at dry-ice temperature for about half an hour to further remove noncondensible gases.

8.3.2.3 Pure Agent Measurements. For low temperature (- 60 °C) measurements of the pure

agents, the experimental set-up, which is similar to the one used in the previous agent screening

project (Grosshandler et ai, 1994), is shown in Figure 3. It consists of a stainless steel (SS 304)
8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 7. Selected thermophysical properties of agents

417

Agent MW

(kg/mol)

T, Tc Pc Pc- P/ <=P2 (kJ/kg) (°C) (MPa) (kg/m^) (kJ/kg K)

HFC-227ea 0.170 -16.4" 101.9' 2.95' 0.46' 621*' 1389' 1.074' 1.177' 126'

CF3I 0.196 -22.0' 122.0'* 4.04"* 0.49* 871'' 2106^ 0.542S 0.592S 106*'

FC-218 0.188 -36.8' 72.0' 2.68' 0.87' 629' 1251' 0.977' 1.190' 101'

HFC- 125 0.120 -48.6' 66.2' 3.63' 1.38' 57 IJ 1190' 1.126' 1.426' 164'

CFjBr 0.149 -57.8' 67.0' 4.02' 1.61' 745'' 1551' 0.670' 0.881' 111'

'From Gallagher et al. (1993)

''From Great Lakes Chemical Corporation (1993)

''From Kudchadker et al. (1979) "From Sladkov and Bogacheva (1992) 'Estimated by the method of Lee and Kesier (Reid et ai, 1987)

^Estimated by the modified Rackett method (Reid et al., 1987) ^Estimated by the method of Rowlinson (Reid et al., 1987)

''Estimated by Pitzer acentric factor correlation (Reid et al., 1987) 'From Braker and Mossman (1980) JFrom Allied Signal Inc. (1990)

''From ASHRAE Inc. (1969)

vessel (tested hydrostatically to 20.6 MPa), a needle valve (Whitey SS-0RS2) for dispensing the

agent and nitrogen, a K-type thermocouple (Omega TJ36-CASS-116U-12), and a pressure transducer

(Druck Model PDCR 330) with an uncertainty of 6.9 kPa. The volume of the whole system was

determined by filling the system with nitrogen to various specified pressures at room temperature and

weighing the pressurized system on an electronic scale with an uncertainty of 0.1 g. With the mass of

nitrogen, temperature, and pressure known, the volume was calculated by using a generalized

correlation for the compressibility factor (Smith and Van Ness, 1975). The total volume of the

apparatus was determined to have a mean of 52.2 cm with a standard deviation of 0.3 cm . The experimental procedure was as follows. The apparatus was evacuated to 1.33 Pa for at least

ten minutes. The set-up was then connected to the agent supply bottle (not shown in Figure 3). The

vessel was placed in a Dewar flask and chilled by using dry ice before gaseous agent was dispensed

through the needle valve into the vessel. Liquid agent was obtained by condensing the agent vapor, and the total amount was metered by placing the vessel and the Dewar flask on an electronic balance.

When the amount of agent reached the target mass, the needle valve was closed. The vessel was then

removed from the Dewar flask, warmed back to room temperature, and weighed on an electronic

balance with an uncertainty of 0.1 g to determine the actual mass in the system.

The filled vessel was then immersed in a heat transfer fluid (Dow Coming Syltherm XLT) which was chilled by two immersion coolers (NESLAB Models IBC-4A and CC 100-11) to - 60 °C. The

Certain commercial products are identified in this report in order to specify adequately the equipment

used. Such identification does not imply recommendation by the National Institute of Standards and

Technology, nor does it imply that this equipment is the best available for the purpose.
418 8. OPTIMIZATION OF SYSTEM DISCHARGE

Cold trap

Figure 1 . Schematic of the degassing apparatus.
8. OPTIMIZATION OF SYSTEM DISCHARGE 419

Vacuum pump

Vacuum

gauge

Relief (HTJ

valve '

Valve

Dry ice bath Molecular sieve

column

Agent

Figure 2. Schematic of CF3I purification system.
420 8. OPTIMIZATION OF SYSTEM DISCHARGE

Pressure

transducer

Thermocouple

Valve

^0UjJ

oo

Pressure

vessel

Figure 3. Experimental set-up for cold temperature measurements.
8. OPTIMIZATION OF SYSTEM DISCHARGE 421

final vessel pressure was recorded from the pressure transducer read-out at least an hour later, after the

internal temperature had reached thermal equilibrium with the bath temperature.

For high temperature (150 °C) measurements, a different set-up was constructed because the

vessel used in the cold temperature experiments could not withstand the pressure that resulted from the

high temperature condition. The apparatus is shown schematically in Figure 4. It consisted of a

stainless steel (SS 316) high pressure vessel with a woridng pressure rated for 68.7 MPa at 204 °C

(Jerguson Series 51 liquid level sight gage), a magnetic pump/mixer, a pressure transducer (Druck

Model PDCR 330), and taper seal high pressure valves (HiP Company). Heating was achieved by

placing the test apparatus in a temperature controlled oven. The internal volume of the entire system

with the valves to the vacuum pump, vent, nitrogen bottle, and sublimizer closed was determined by

evacuating the system followed by sucking distilled water into the system from a beaker. The amount

of water required to fill the system was then determined from the loss of weight of the beaker.

Knowing the mass of water and its density at room temperature, the internal system volume was

calculated to have an average of 42.6 cm^ with a standard deviation of 0.2 cm^. Each window on the pressure vessel was also calibrated in terms of the volume percent of liquid filled to facilitate the

filling process. The calibration was conducted by measuring the amount of water required to fill up to each window. In order to minimize photolysis of CF3I, agent filling was performed under a darkroom

safelight. In addition, when CF3I was used in a test, extreme caution was taken to prevent any stray room light from getting into the oven. The experimental procedure was as follows. The entire system was evacuated for at least half an

hour before liquid agent was dispensed from the sublimizer. The exact amount of agent in the system was determined by weighing the sublimizer. The oven was then turned on, and the test apparatus was

heated to the desired temperature. Because of space limitation, it was not possible to insert a thermocouple mto the vessel to monitor the internal temperature. To ensure that thermal equilibrium

had been reached, two thermocouples were placed at two different locations on the vessel wall, and the

final pressure was recorded only when the two thermocouple readings were within 1.0 °C of the set temperature of the oven and the final pressure maintained a steady reading for at least half an hour.

8.3.2.4 Nitrogen Solubility in Agents. To measure the total amount of nitrogen required to pressurize the vessel to a specified equilibrium pressure at room temperature, the set-up for the cold

temperature experiments was used. The agent filling procedure was the same as in the cold temperature studies. After a pure liquid agent was dispensed in the vessel, the nitrogen gas supply line (not

shown in Figure 3) was connected to the vessel. The supply line was evacuated for at least five minutes before initiating a slow flow of nitrogen. The needle valve was then opened to allow nitrogen

to bubble slowly through the liquid agent until an equilibrium pressure of approximately 2.8 MPa was

reached. Shaking the vessel intermittently and vigorously and repetitive bubbling of nitrogen was

required before the final equilibrium pressure was attained. The amount of nitrogen needed to pressurize the vessel was obtained by weighing the apparatus. This amount corresponded to the sum

of the mass of nitrogen in the gas ullage of the vessel and that dissolved in the liquid agent. The

charged vessel was then ready for the cold temperature experiments, as described in Section 8.3.2.5. When the cold experiments were completed, the charged vessel was warmed back to room temperature

and was pressurized again with nitrogen to approximately 4.1 MPa. The amount of nitrogen was then

determined, and the vessel was ready for another cold temperature experiment.

8.3.2.5 Agent/Nitrogen Mixture Measurements. The same two apparatus (see Figures 3 and 4)

used for pure agent studies were also utilized to perform agent/nitrogen mixture experiments. In fact,

since the experimental procedure for the solubility studies was equivalent to the preparation of the

vessel at room temperature for subsequent pressure-temperature measurements, cold temperature
422 8 OPTIMIZATION OF SYSTEM DISCHARGE

Vacuum pump

Pressure

display

Pressure

transducer

Pressure

vessel

0

o

o

o

o

o

o

Magnetic

mixer

Sight windows

To vent

Figure 4. Apparatus for high temperature experiments.
8. OPTIMIZATION OF SYSTEM DISCHARGE 423

experiments were therefore performed immediately after the solubility measurements were made by

simply immersing the same vessel in the chilled heat transfer fluid. For high temperature measurements, the experimental procedure was as follows. The same agent

filling procedure in the pure agent studies was used. After the agent was dispensed, the pressurization

of the vessel with nitrogen to a specified equilibrium pressure (approximately 2.8 MPa) at room

temperature was initiated with the magnetic pump running to facilitate the mixing of nitrogen with the

agent. After pressurization, the apparatus was heated to the final temperature (150 °C), and the final pressure of the vessel was recorded the same way as in the pure agent experiments. The apparatus was then cooled back to room temperature, pressurized with nitrogen to approximately 4.1 MPa, and

heated to the final temperature (150 °C), where the equilibrium pressure was recorded.

8.3.3 Modeling Thermodynamic Properties of Agent/Nitrogen Mixtures

8.3.3.1 Introduction. The thermodynamic properties of agent/nitrogen mixtures are modeled

using a computer program PROFISSY (acronym for Properties Of Fi^re Suppression Systems). The

program incorporates a model known as "extended corresponding states" (ECS). This model has been

quite successful in modeling the thermodynamic and transport properties of fluids. An early model of

this type, known as the TRAPP model (Ely and Hanley, 1981a, 1981b, and 1983), has been successful

in the prediction of hydrocarbon transport properties, including LNG (McCarty, 1982). Much of the

development of TRAPP occurred in the early 1980's at the National Institute of Standards and

Technology. Since that time, the model has been revised, improved, and used successfully to predict

the properties of hydrocarbon mixtures (Ely, 1982; Huber and Ely, 1990; Friend, 1992), air (Ely, 1990), and C02-rich hydrocarbon mixtures (Sherman et ai, 1989), and forms the basis of two

current Standard Reference Data Products (Huber and Ely, 1990; Friend, 1992). ECS has been used

successfully to predict equilibrium properties such as density, heat capacity, speed of sound, entropy,

enthalpy, Joule-Thomson coefficient, viscosity, and thermal conductivity. It is a powerful method,

applicable to the entire range of fluid states, from dense liquid to dilute gas, as well as to the

supercritical fluid regime. It may be used with only minimal information on a fluid: the critical point,

the normal boiling point, and the molecular weight. Additional information on a fluid, such as vapor

pressures, saturated liquid densities, and liquid viscosities can be used to refine the model predictions.

Recently ECS models have been used to model thermophysical properties of alternative refriger- ants and their mixtures (Huber and Ely, 1992; Huber et ai, 1992; Gallagher et al., 1995). In this work, we extended the model to calculate the thermodynamic properties of agent/nitrogen mixtures.

8.3.3.2 Extended Corresponding States Model. The central idea of extended corresponding

states is that all points on the PVT surface of any fluid may be represented by scaling the PVT surface

of a reference substance. These "scale factors" involve the critical properties of the fluid of interest and the reference fluid and may also be functions of temperature and density.

The basic equation in the extended corresponding states theory relates the dimensionless residual

Helmholtz free energy of a pure fluid, denoted by subscript j, to that of a reference fluid, denoted by

subscript o, whose thermodynamic properties are known, in principle, with great accuracy:

(5)
424 8. OPTIMIZATION OF SYSTEM DISCHARGE

where the residual Helmholtz free energy is defined as

^ Aip,T) - A'ip,T)

RT

(6)

In the above two equations, A is the Helmholtz free energy, the asterisk indicates an ideal gas

state value, p is the molar density, T is the absolute temperature, and R is the universal gas constant. The temperature and molar density of the fluid ; are scaled with the corresponding values for the

reference fluid o according to the following two equations:

r, = 1 (7)

Po = Pjhj (8)

where the scale factors fj and hj are called the equivalent substance reducing ratios. In extended

corresponding states between two pure fluids ; and o, they are related to the critical parameters of the

two fluids, namely

fj = ^

Qj(Pj,Tj) (9)

C, 0

'

ft^ = *^(p^.,7- ) (10)

' '

where the functions 9^ and are called shape factors, and subscript c denotes a critical property. In addition, the condition that the compressibility factors of the fluids be equal, Z ''y = Z is imposed.

Other dimensionless residual thermodynamic properties of a fluid can be derived from differentiation

of Equation (5) and can be found in Huber and Ely (1994).

To calculate properties of mixtures using extended corresponding states, we use what is known as a "one-fluid" mixture model. The properties of a mixture are found by first assuming the mixture to be a single hypothetical pure fluid (subscript x) through the use of some mixing and combining rules. In this way, the properties of the hypothetical pure fluid are then found by scaling with a single

reference fluid o.

(11)

In order to obtain the equivalent substance reducing ratios, and h^, for the hypothetical pure

fluid, we use the van der Waals one-fluid mixing rules,
8. OPTIMIZATION OF SYSTEM DISCHARGE 425

1=1 j=\

fh=yy x.x.f. h.. (13)

and the combining rules,

(14)

8

where x is the mole fraction, n is the number of components, subscripts / and j refer to component / and j respectively, k-j and /,y are the binary interaction parameters that may be non-zero when / ^ j and

k- and / are defined to be zero. The parameters are generally found empirically by fitting experimental data. However, in order to test the robustness of the ECS model without resorting to the experimental data which have to be obtained a priori, the binary interaction parameters are set equal to zero

in the model.

Several factors contribute to the ability of the ECS model to accurately predict thermophysical

properties. Two important factors are the choice of the reference fluid and the determination of the

shape factors. It is important to have a very well characterized empirical reference fluid equation of

state. The reference fluid used in this work is HFC- 134a (1,1,1,2-tetrafluoroethane) which is represented by the Jacobsen-Stewart (1973) modified 32 term Benedict-Webb-Rubin equation of state (MBWR-32) with coefficients as determined by Huber and McLinden (1992). This equation has been

used extensively to represent the properties of hydrocarbons, common inorganics, and most recently

refrigerants. Its functional form is essentially a polynomial in density and temperature and is given by

>2 15 j:a„(Dp'' ..-^^'^^^ a„(Dp^"-'^ (16)

n = l n = 10

The temperature dependence of the aJ,T} is summarized in Table 8. The 32 coefficients {bj to b^2)

for the reference fluid HFC- 134a are tabulated in Huber and Ely (1994). Equation (16), when applied

to HFC- 134a, typically reproduces the pressures within 0.4 % and the density within 0.2 %.

Exceptions are large density deviations in the near critical region, and large pressure errors in the lowtemperature compressed liquid region near the saturation boundary.
426 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 8. Temperature dependence of the MBWR-32 coefficients

a 1/ RT "9 - b,„ / r

^

«2 b,T + b2T^'^ + + bj/T+ b^/T^ °I0 = b,o/T'- + b„/T'

«.? bj + bj + bf^ / T + b^/T^ = b,,/T' + b2s/T'^

^4 biJ + b,i + bi2/T a/2 = b,,/T^

«5 «/j ^b,,/T'

«6 b,4/T+b,^/T^ a,

4

= b,,/T' + b2g/T^

«7 b„/T "15 = b,o/T'

«S b,j/T+b„/T^

The accuracy one obtains from ECS models also depends on how well the shape factors are determined. The shape factors are functions of both density and temperature. In principle, if one has

accurate, analytical equations of state for the two fluids, the exact shape factors relating them can be

found, as discussed by Ely (1990). However, often only one fluid (the reference) has an accurate,

analytical equation of state available. In this situation, one usually assumes an empirical form for the

shape factors. Leach et al. (1968) expressed the shape factors as weak functions of temperature and

density using the acentric factor as a third correlating parameter. Since the shape factors are much

stronger functions of temperature than density and most of the data which are available for refrigerants are for the saturation boundary, we have chosen a different approach utilizing information along the

saturation boundary of the fluids (Huber and Ely, 1994) which results in density independent shape

factors. It can be shown (Huber and Ely, 1994) that the vapor pressures and the saturated liquid

densities of a fluid can be related to the reference fluid properties through the following equations

p (r.) = p

(li) L (17) ^saty ^sat, ^ f

p (7.) = p (^) 1 (18)

"sat, J'' f^sat, " f fi

•'J J

where subscript sat refers to a property of a fluid along the saturation boundary. Equations (17) and

(IS) are simultaneously solved to obtain and hj along the saturation boundary. Vapor pressures and

saturated liquid densities are provided from empirical correlations of data, if available, or from

estimations (Reid et ai, 1987). Additional information on the procedure for obtaining shape factors can be found in Huber and Ely (1994).

We do not yet have extensive results on the performance of the ECS model for predicting the

density of agent/nitrogen mixtures; however, our experience with hydrocarbon mixtures and with

refrigerant mixtures gives us confidence that generally the densities of the liquid phase are accurate to within 3 %. This number can be improved with the use of the binary interaction parameters.
8. OPTIMIZATION OF SYSTEM DISCHARGE 427

8.3.3.3 Calculation Procedure. The computer code PROFISSY was developed with the primary

purpose of helping fire suppression bottle designers or users to obtain temperature-pressure characteristics of the bottle. In other words, given a vessel (V,) charged with agent and nitrogen at room

temperature, one would like to know what the final vessel pressure will be when the vessel is exposed

to a different temperature. In this aspect, only the information which is pertinent to the problem is provided in the output although the computer program can perform other thermodynamic property

calculations.

The first step in the calculation is to determine the initial conditions of the vessel, that is to determine the compositions of the liquid and vapor phases, the fraction of the mixture (agent/nitrogen)

that will be in the vapor phase, and the amount of nitrogen required to pressurize the vessel or the

initial total pressure of the vessel given an initial amount of nitrogen. To perform such calculations, it

is required that the amount of agent, the fill temperature, the vessel size, and the total pressure of the

vessel or the amount of nitrogen in the vessel be known. The calculation is essentially a flash calculation at fixed T and P. The flash calculation can be generally stated as follows:

Given a mixture, in this case agent/nitrogen, whose bulk compositions {Zj) are known at T and P,

determine yj, and fraction vaporized (a), where x is the liquid mole fraction and y is the vapor mole

fraction. The solution of the problem requires mass balance and vapor-liquid equilibrium calculations.

A detailed computational procedure can be found in Prausnitz et al. (1980) and Walas (1985). In our

case, the calculation is slightly different depending on whether ( 1 ) the amount of nitrogen (n^ ,) or (2)

the total initial charge pressure (P,) is given.

In Case (1), with V,, T^, n^^, and n^, given, we calculate P-, x^, y^, and a, where n is the number

of moles and subscripts, t, g, and a represent total, nitrogen, and agent respectively. Note that for a

binary system x^ and y^ are known once x^ and y^ are determined. In Case (2), knowing V,, T^, ^, and P:, then n , x , y , and a are calculated. Irrespective of Case (1) or (2), bubble point and dew

point calculations (see e.g., Prausnitz et al., 1980) need be carried out to ensure that the flash calculations are performed in the two-phase region. Flash calculations are feasible only when the

vessel conditions are above the bubble point and below the dew point. Below the bubble point, the

mixture is a subcooled liquid; in this case, x- = z,. Above the dew point, the mixture is a superheated

vapor; in this case y^ = z-. Once the initial conditions of the vessel are established, calculations of temperature-pressure

characteristics can proceed. In this case, given V,, and we want to calculate P^, x^, y^, and a

at 7^ if the resulting condition is in a two-phase region. The computational procedure is similar to Case (1) above. If the resulting condition in the vessel is in a single-phase region, then straightforward PVT calculations will provide given V^, Tp and z,- (= x^ or y^.

Only four pieces of information are required to run the program: (1) agent mass, (2) vessel

volume, (3) fill temperature, and (4) either nitrogen mass needed to pressurize the vessel, or the fill pressure of the vessel. Appendix A provides a typical session showing how to run the computer

program. The current version of the program only supports the agents used in this study although

other agents can be incorporated into the program by modifying the property data base.

8.3.4 Results and Discussion

8.3.4.1 Pure Agents. The results of the temperature-pressure measurements for pure agents are summarized in Table 9. In all cases, the combined standard uncertainty of the measurement was

estimated to be less than 0.2 MPa based on the law of propagation of uncertainty with the sensitivity

coefficients calculated using the ideal gas law. The total agent mass in the second column corresponds

to the initial liquid agents whose volumes in the two cases measured approximate one-half or two-
428

Table 9.

8. OPTIMIZATION OF SYSTEM DISCHARGE

Comparison of measured pressures of pure agents at 150 °C ± 2 °C with predicted values

Agent Agent, , (g)

(measured)

(MPa)

Py (predicted)

(MPa)

error^ (%)

HFC-227ea 28.3 6.83 6.47 5.3

30.3 7.10 6.74 5.1

40.7 10.82 10.33 4.5

38.2 10.09 9.32 7.6

CF3I 43.0 6.51 6.62 1.7

43.0 6.77 6.77 0.0

57.3 8.84 9.23 4.4

58.6 9.52 9.85 3.5

FC-218 28.9 8.33 8.01 3.8

29.1 8.32 8.01 3.7

37.4 11.90 11.72 1.5

37.3 11.85 11.74 0.9

HFC- 125 26.1 11.57 11.29 2.4

34.6 16.6! 16.01 3.6

32.3 15.21 14.83 2.5

CF3Br 34.9 11.98 11.50 4.0

33.8 11.68 11.21 4.0

44.4 16.28 15.55 4.5

45.6 16.97 16.26 4.2

"error (%) = | (measured) - (predicted) | x 100 / (measured)

thirds of the system volume at room temperature. The measurements were obtained at ~ 150 °C.

Only the high temperature results are reported here because at low temperature (- 60 °C) the pressures

for all the agents were found to be below 0. 1 MPa, which was the lowest value that could be

measured from the pressure transducer.

For all the tests with liquid agent initially half filling the vessel, the final pressures at 150 °C

were found to be less than 12.5 MPa. For an initially two-thirds filled vessel, the final pressures at 150 °C were measured to be less than 17.5 MPa.

The predicted final pressures obtained using PROFISSY are also tabulated in Table 9. Note that PROFISSY was developed specifically for agent/nitrogen mixtures with no intention to duplicate the

function of the REFPROP computer program (Gallagher et ai, 1993) (which is designed to estimate

thennophysical properties of pure and mixed refrigerants). An attempt to use REFPROP for the pure

agents proved to be unfruitful because the final conditions attained in the experiments exceeded the

calculational range of REFPROP. To use PROFISSY for the estimation, 0.00001 g of nitrogen was
8. OPTIMIZATION OF SYSTEM DISCHARGE 429

artificially input into the program to simulate the calculations for a pure agent. This procedure was

found to be justifiable because the calculated initial charge pressures at the initial fill temperatures were very close to the measured agent vapor pressures at the same temperatures. In most cases except

CF3I, PROFISSY tended to underestimate the final pressures. Overall, the predictions were found to be within 7.6 % or less of the measured values. Since very few previous studies related to CF3I vapor pressure measurements have been reported

in the literature (Smith and Srivastava, 1986), temperature-pressure relationships for CF3I were also

measured over a temperature range from - 20 °C to 150 °C. From these measurements, the vapor

pressure as a function of temperature was determined. Figure 5 shows three sets of PT measurements

using initial liquid agent volumes corresponding to one-third, half, and two-thirds of the vessel volume

at room temperature, respectively. The trend in Figure 5 can be explained based on a comparison

between the fill density, defined as the ratio of the total agent mass to vessel volume, and the critical density of CF3I (Grosshandler et ai, 1994). Irrespective of the fill conditions at room temperature, if the vessel is chilled to a lower temperature, the measured pressure will correspond to the vapor

pressure at that temperature. On the contrary, if the vessel is heated to a higher temperature, the pressure that one measures depends on whether the liquid and vapor phases still coexist at that temperature. To determine whether the two phases are still present at the temperature of interest, one

has to find out whether the fill density lies between the saturated vapor and liquid densities at that temperature (Smith and Van Ness, 1975). Using the same reasoning, vapor pressures that were

extracted from Figure 5 in the temperature range between - 20 °C and 70 °C are shown in Figure 6. The data were fitted in the form of log P^^, = A - {BIT) where A and B have values of 3.186 and 1037

respectively, P^^, is in MPa and T is in K. Note that the above equation should only be used within

the stated temperature range.

8.3.4.2 Solubility of Nitrogen in Agents. Table 10 summarizes the experimental results obtained under an initial total pressure (P,) of approximately 2.8 MPa or 4.1 MPa at room temperature

(r, = 23 °C ± 1 °C) and with an initial amount of liquid agent corresponding to one-half of the vessel

volume. For all the agents evaluated, the amounts of nitrogen (m^ ,) required to pressurize the vessel

to - 2.8 MPa and - 4.1 MPa were found experimentally to be less than 4 % and 6 % of the total mass

of the agent (m^ ,), respectively. The predicted nitrogen mass in the fifth column and the predicted

solubilities of nitrogen {x^i) in the agent under 7, and P- in the last column of Table 9 were obtained

using PROFISSY. The calculations were performed with V, (= 52.2 cm^), and 7- as input and

the two binary interaction parameters {k- and /,y) set equal to zero. The solubility is defined as the

mole fraction of dissolved nitrogen in the liquid agent/nitrogen mixture at pressure P-. The predicted

nitrogen mass compares favorably with the measured value even without taking and l^j into

consideration. The solubility of nitrogen in FC-218 was found to be the highest compared to the other

four agents, whereas nitrogen exhibited the lowest solubility in CF3Br when P- was ~ 2.8 MPa and in CF3I when P- was - 4.1 MPa.

Table 1 1 tabulates the experimental results as well as the predictions obtained under similar

experimental conditions as in Table 10, except that the initial amount of liquid agent corresponded to two-thirds of the vessel volume. The predicted nitrogen mass also compares favorably with the

measured value. For all the agents evaluated, the amount of nitrogen required to pressurize the vessel was found experimentally to be less than 3 % to 4 % of the total mass of the agent. Note that the

calculated solubilities in Table 11 are very close to those tabulated in Table 10. This is not surprising

because the solubility depends only on and 7, and does not depend on the amount of liquid agent.

Again, FC-218 was found to have the highest nitrogen solubility compared to the other four agents,

irrespective of P,. At low P, (~ 2.8 MPa), the solubility of nitrogen in CFgBr was calculated to be the
430 8. OPTIMIZATION OF SYSTEM DISCHARGE

Oh

12

11

10

9

8

7

6

5

4

3

2

1

1 r

1 r1 r

CF3I (pure)

o 1/3 fill • 1/2 fill V 2/3 fill

•

O

o

o

0

-40 -20 0 20 40 60 80 100 120 140 160 180 200

Temperature (^C)

Figure 5. Pressure-temperature measurements of pure CF3I.
8. OPTIMIZATION OF SYSTEM DISCHARGE 431

0.4

0.2

0.0

-0.2

IS -0.4

Q -0.6

-0.8

•1.0

1.2

2.5

ft

O 1/3 fill • 1/2 fill V 2/3 fill

fit: log P = 3.186 - (1037/T)

S3.t

3.0 3.5

1/T (K'S X 1000

4.0 4.5

Figure 6. Vapor pressure of CF3I as a function of temperature.
432 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 10. Amount of nitrogen (w^,) required to pressurize mixture to P- at T,- = 23 °C ± 1 °C with an initial amount of liquid agent corresponding to half of the vessel volume

Agent Agent, , (g)

Pi (MPa)

Nitrogen, , (g)Predicted

, (g)

Solubility,

, (predicted)

HFC-227ea 36.5 2.90 1.2 1.24 0.083

36.6 2.93 1.2 1.26 0.084

36.5 4.29 1.9 1.95 0.129

36.6 4.28 1.9 1.94 0.128

CF3I 54.9 2.87 1.0 1.15 0.056

54.8 2.87 1.1 1.15 0.056

54.9 4.21 1.6 1.79 0.087

54.8 4.21 1.7 1.78 0.087

FC-218 33.2 2.89 1.0 1.07 0.090

33.1 2.89 1.1 1.07 0,090

33.2 4.27 1.8 1.77 0.152

33.1 4.31 1.8 1.80 0.154

HFC- 125 31.3 2.78 0.9 0.81 0.053

31.1 2.79 0.8 0.81 0.053

31.3 4.24 1.6 1.62 0.106

31.1 4.23 1.6 1.61 0.106

CFjBr 40.9 2.89 0.7 0.75 0.046

40.7 2.87 0.6 0.75 0.046

40.9 4.29 1.4 1.53 0.095

40.7 4.29 1.4 1.53 0.095

lowest, whereas the solubility of nitrogen in CF3I was found to be the lowest at high P- (~ 4.1 MPa).

This is quite explainable by the great difference in the vapor pressures of these two agents, as noted in Table 7.

8.3.4.3 Agent/Nitrogen Mixtures. The experimental results for agent/nitrogen mixtures at - 60 °C with an initial volume of liquid agent corresponding to half of the vessel volume are shown in Table 12. In all cases, the combined standard uncertainty of the measurement was found to be less than 0.1 MPa. The sensitivity coefficients in the uncertainty estimation were calculated based on PROFISSY. Comparing to the vapor pressure of a pure agent at - 60 °C, the measurements showed

that there were still significant residual pressures in the vessel. These residual pressures are deemed

necessary to expel the cold liquid agents if the vessels are discharged under such extremely low

temperature. The fifth column in the table tabulates the predicted final pressures using PROFISSY.
8. OPTIMIZATION OF SYSTEM DISCHARGE 433

Table 1 1 . Amount of nitrogen (m^ ^) required to pressurize mixture to P- at T- (= 23 °C ± 1 °C) with an initial amount of liquid agent corresponding to two-thirds of the vessel volume

Agent Agent, m^, Pi Nitrogen, , Predicted

, Solubility,

i (g) (MPa) (g) (g) (predicted)

HFC-227ea 48.7 2.87 1.1 1.16 0.082

48.7 2.98 1.2 1.21 0.086

48.7 4.25 1.8 1.81 0.127

48.7 4,25 1.8 1.82 0.127

CF3I 72.8 2.77 0.9 1.02 0.054

72.7 2.79 0.9 1.02 0.054

72.8 4.14 1.4 1.62 0.086

72.7 4.16 1.4 1.63 0.086

FC-218 44.1 2.94 1.0 1.07 0.096

44.2 2.93 1.0 1.06 0.096

44.1 4.36 1.6 1.81 0.161

44.2 4.29 1.6 1.75 0.157

HFC- 125 41.9 2.70 0.8 0.79 0.051

41.8 2.85 0.8 0.83 0.055

41.9 4.18 1.5 1.56 0.104

41.8 4.18 1.5 1.51 0.102

CF3Br 54.9 2.92 0.7 0.76 0.048

54.8 2.87 0.7 0.74 0.046

54.9 4.25 1.4 1.48 0.094

54.8 4.25 1.4 1.49 0.094

The calculations were performed with k-j and l^j set equal to zero, V, = 52.2 cm , and the initial conditions given in the table. In general, the predicted values were all found to be within 6 % of the

measured values.

Table 13 summarizes the results obtained from experiments at - 60 °C using an initial liquid

agent volume equivalent to two-thirds of the vessel volume. Considerable final pressures in the vessel were also noted at - 60 °C. Predictions using PROFISSY with and 1- set equal to zero and = 52.2 cm^ were also carried out, and the comparisons between predicted and measured final pressures indicate that overall the errors were within 8 % for all the agents.
434 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 12. Initial conditions at 7. = 23 °C ± 1 °C and final pressures {P^^) at - 60 °C ± 1 °C for

agent/nitrogen mixtures (half filled volume)

Agent

(g) (MPa)

t^f ^ (MPa;

(measured)

^ (MPa)

(predicted)

error^ (%)

HFC-227ea 36.5 2.90 1.69 1.70 0.6

36.6 2.93 1.72 1.71 0.6

36.5 4.29 2.59 2.60 0.4

36.6 4.28 2.58 2.61 1.2

CF3I 54.9 2.87 1.71 1.61 5.9

54.8 2.87 1.71 1.62 5.3

54.9 4.21 2.62 2.49 5.0

54.8 4.21 2.61 2.47 5.4

FC-218 33.2 2.89 1.44 1.39 3.5

33.1 2.89 1.43 1.38 3.5

33.2 4.27 2.34 2.23 4.7

33.1 4.31 2.38 2.27 4.6

HFC- 1 25 31.3 2.78 1.11 1.07 3.6

31.1 2.79 1.11 1.08 2.7

31.3 4.24 2.10 2.04 2.9

31.1 4.23 2.09 2.03 2.9

CFjBr 40.9 2.89 1.04 1.02 1.9

40,7 2.87 1.00 1.02 2.0

40.9 4.29 1.95 1.94 0.5

40.7 4.29 1.99 1.97 1.0

"error (%) = | Py-^( measured) - Py;^(predicted) | x 100 / Py-^(measured)

The results for high temperature (150 °C) measurements are shown in Tables 14 and 15 for initial conditions corresponding to half and two-thirds filled volumes, respectively. The combined standard

uncertainty, in all cases, was estimated to be less than 0.2 MPa, and the sensitivity coefficients were

calculated based on an ideal gas mixture model. The predictions were obtained by using PROFISSY.

In the calculations, k^j and l^j were again assumed to be zero and V, = 42.6 cm was used. For these

two experimental conditions, the errors between the predicted and measured final pressures were found

to be within 10 %.
8. OPTIMIZATION OF SYSTEM DISCHARGE 435

Table 13. Initial conditions at = 23 °C ± 1 °C and final pressures (P^-^) at - 60 °C ± 1 °C for

agent/nitrogen mixtures (two-thirds filled volume) ,:,

Agent

(g) (MPa)

Pf,, (MPa)

(measured)

Pf,c (MPa)

(predicted)

error" (%)

HFC-227ea 48.7 2.87 1.66 1.67 0.6

48.7 2.98 1.70 1.73 1.8

48.7 4.25 2.55 2.57 0.8

48.7 4.25 2.51 2.59 3.2

CF3I 72.8 1.11 1.60 1.52 5.0

72.7 2.79 1.64 1.53 6.7

72.8 4.14 2.49 2.43 2.4

72.7 4.16 2.62 2.41 8.0

FC-218 44.1 2.94 1.41 1.41 0.0

44.2 2.93 1.39 1.38 0.7

44.1 4.36 2.32 2.32 0.0

44.2 4.29 2.30 2.27 1.3

HFC- 125 41.9 2.70 1.08 1.06 1.9

41.8 2.85 1.10 1.11 0.9

41.9 4.18 2.06 2.05 0.5

41.8 4.18 2.02 1.99 1.5

CFjBr 54.9 2.92 1.05 1.04 1.0

54.8 2.87 1.02 1.02 0.0

54.9 4.25 1.94 1.94 0.0

54.8 4.25 1.90 1.96 3.2

^error (%) = | Py-^Cmeasured) - Py-^Cpredicted) | x 100 / Py-^(measured)
436 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 14. Initial conditions at 7- = 23 °C ± 1 °C and final pressures (P^;,) at 150 °C ± 2 °C for

agent/nitrogen mixtures (half filled volume)

Agent '"a.t (g)

p. (MPa)

f,h v^" "/ (measured)

f,h V^*»^i/ (predicted)

ciTor'^ (%)

HFC-227ea 28.3 2.76 10.70 10.39 2.9

30.3 2.81 11.25 10.95 2.7

28.3 4.29 13.58 13.25 2.4

30.3 4.19 14.43 14.00 3.0

CF3I 43.2 2.75 10.38 10.55 1.6

43.2 2.74 10.22 10.35 1.3

43.2 4.13 12.71 13.03 2.5

43.2 4.16 12.92 13.05 1.0

FC-218 28.9 2.85 12.13 11.90 1.9

29.1 2.76 11.84 11.76 0.7

28.9 4.16 14.64 14.44 1.4

29.1 4.25 14.73 14.59 1.0

HFC- 125 26.4 2.79 14.45 14.43 0.1

26.1 2.78 14.35 14.41 0.4

26.4 4.18 17.32 17.56 1.4

26.1 4.21 17.20 17.36 0.9

CFjBr 34.9 2.80 14.32 14.04 2.0

34.9 4.22 17.11 16.90 1.2

"error (%) = | Pr^(measured) - P^^(predicted) | x 100/ fr;j(measured)
8. OPTIMIZATION OF SYSTEM DISCHARGE 437

Table 15. Initial conditions at T,- = 23 °C ± 1 °C and final pressures (Pfj) at 150 °C ± 1 °C for

agent/nitrogen mixtures (two-thirds filled volume)

Agent

(g)

Pi (MPa)

Pf,h (MPa)

(measured)

Pf.h (MPa)

(predicted)

eiTor'' (%)

HFC-227ea 38.4 2.80 15.56 14.57 6.4

39.7 2.78 16.77 15.34 8.5

38.4 4.14 19.19 17.71 7.7

39.7 4.29 20.54 18.87 8.1

CF3I 57.3 2.76 13.79 14.34 4.0

58.6 2.79 14.42 15.08 4.6

57.3 4.23 17.62 18.38 4.3

58.6 4.10 18.25 18.41 0.9

FC-218 37.4 2.88 16.39 16.24 0.9

37.3 2.78 16.51 16.68 1.0

37.4 4.16 20.03 20.01 0.1

37.3 4.23 19.99 19.94 0.3

HFC- 125 34.6 2.76 20.13 20.37 1.2

32.3 2.86 18.87 18.94 0.4

34.6 4.16 24.25 24.71 1.9

32.3 4.18 22.39 22.87 2.1

CFjBr 44.4 2.76 18.92 18.68 1.3

45.6 2.76 20.19 19.69 2.5

44.4 4,18 22.91 22.82 0.4

45.6 4.21 23.22 23.31 0.4

^error (%) = | /'^;,(measured) - Py-^(predicted) | x 100 / Py-^(measured)
438 8. OPTIMIZATION OF SYSTEM DISCHARGE

8.3.5 Summary and Conclusions. For all the pure agents studied under two fill conditions (1/2 and

2/3 filled), the final pressure that the vessel experienced when it was heated to 150 °C was found to be

less than 17.5 MPa. Irrespective of the initial fill conditions, the final pressures were all below

0.1 MPa when the vessel was chilled to - 60 °C and pure agents were used.

For an initial fill condition corresponding to approximately half of the vessel volume, the amount

of nitrogen required to pressurize the vessel up to 4.4 MPa was measured to be less than 6 % of the

total agent mass in the vessel in all cases. Under an initial condition of two-thirds liquid fill, the

required nitrogen mass was found to be less than 4 % of the total agent mass. Independent of the

initial amount of liquid agent in the vessel, the mole fractions of nitrogen in the liquid agents

(solubilities) at room temperature were calculated to be less than 0.10 and 0.17, respectively, at 2.8 MPa and 4.1 MPa.

For agent/nitrogen mixtures with an initial liquid agent fill condition of one half or two-thirds and

initial total pressures of approximately 2.8 MPa and 4.1 MPa at room temperature, the final pressures

measured at - 60 °C were found to be higher than 0.9 MPa and 1.8 MPa respectively in all cases. Compared to the vapor pressures of the pure agents, significant residual pressures were still present in the vessel at - 60 °C. These residual pressures, as the result of the initial nitrogen pressurization, are

vital to agent discharges at cold ambience.

For agent/mixtures with an initially half full condition and initial charge pressures of approximately 2.8 MPa and 4.1 MPa, the resulting pressures at 150 °C were measured to be less than 15 MPa

and 18 MPa respectively. For the two-thirds condition under similar initial charge pressures, the final pressures were found to be less than 21 MPa and 25 MPa respectively. Comparisons of the measure- ments among the four selected agents showed that HFC- 125 exhibited the highest final pressures at 150 °C under all the experimental conditions used in this study.

The thermodynamic properties of agent/nitrogen mixtures were calculated using the computer

code PROFISSY which is based on an extended corresponding states model. The code was developed

to help fire suppression bottle designers to facilitate the calculations of bottle temperature-pressure

relationships. Comparing to the experimental measurements, the predictions from PROFISSY were

generally found to be within 10 % or less of the measured values. Based on the experimental results from this study, the following conclusions are made:

1 . Under the current pressure vessel design criteria in military specification MIL-C-22284A

(proof pressure of 9.62 MPa and minimum burst pressure of 12.37 MPa), the structural

integrity of the vessel will be severely compromised by the resulting high pressure if the

vessel, when used with the four selected agents, is exposed to ambient temperatures up to 150 °C.

2. When using the computer code PROFISSY, extreme care should be taken because PROFISSY

tends to underestimate the final pressures at elevated temperatures by about 10 % in some

cases.

8.4 Discharge of Agent/Nitrogen Mixtures in a Simulated Dry Bay

8.4.1 Introduction. Currently, CF3Br (halon 1301) is used for in-flight dry bay fire protection

applications. It is normally stored under pressure above its vapor pressure at room temperature by

using nitrogen. Due to nitrogen pressurization, nitrogen is dissolved in the liquid agent, and the

presence of dissolved nitrogen complicates the discharge process in that the fluid leaving the vessel is
8. OPTIMIZATION OF SYSTEM DISCHARGE 439

not pure halon but a mixture of halon and nitrogen and that degassing of the dissolved nitrogen {i.e., the dissolved nitrogen coming out of the liquid in the form of bubbles) occurs inside the vessel during

discharge (depressurization).

The search for halon alternatives has resulted in three potential agents, HFC- 125, FC-218, and

CF3I, for dry bay fire protection applications from a list of more than ten candidates

(Grosshandler et ai, 1994). This task is an extension of the study of agent discharge fluid dynamics

previously reported as a part of the agent screening work in Grosshandler et al. (1994). In the

previous study, the agent was only pressurized with nitrogen a few minutes before the initiation of a

test. Given such a short contact time of nitrogen with agent, it was not possible to have an appreciable amount of nitrogen dissolved in the agent. The experimental protocol was dictated by an idealized

discharge mechanism (a rupture disc). Rupture discs, which offered simple flow geometry and quick

action and also facilitated the agent screening process, were chosen as the result of a compromise

between a simulated and a real discharge event. Furthermore, it was believed that the effect of

dissolved nitrogen on the discharge process was similar among all the agents evaluated because the

nitrogen solubilities in all the agents were measured to be very similar (Grosshandler et al, 1994) and

that any observable differences in the discharge dynamics among the agents studied were attributable

to the agents themselves rather than the dissolved nitrogen.

In this study, the role of dissolved nitrogen in the agent discharge dynamics was examined in a

systematic way. Actual discharge mechanisms (solenoid valves and squibs) were used to simulate

discharge processes. The main objective of this task was to study the discharge characteristics of the

three selected agents (CFjBr was also included as a reference) in an unconfined space to simulate a

dry bay (without airflow and obstacles) under various experimental conditions in greater detail than the

previous screening project. In particular, the parameters of interest include size of valve opening, fill density, nitrogen charge pressure, initial agent temperature, discharge orientation, and the degree of

nitrogen saturation in liquid agent.

8.4.2 Experimental Methods. This section describes the experimental hardware and procedures used

to performed the dry bay discharge tests. The simulated dry bay was essentially a room with

dimensions of 3 m (width) x 3 m (length) x 3.4 m (height). No clutter or obstacles, except a pressure

transducer located about 1 .4 m downstream from the vessel outlet, were placed in the flow path of the

flashing spray. Depending on the release mechanism used and discharge orientation, the vessel was

centrally located either on one side of the room, near the ceiling, or near the floor. There were two

major components in the experimental set-up: (1) the pressure vessel and (2) the release mechanism.

Three pressure vessels, a cylindrical, a spherical, and a cylindrical with sight gages, were used in

this study. The choice of the vessel was dictated by the release mechanism. The cylindrical vessel,

which was also used in the previous study (Grosshandler et al., 1994), was made from a stainless steel SS 304 tube with an internal diameter of 50.8 mm and a wall thickness of 6.4 mm. One end of the

tube was welded with an end cap in which four holes were drilled and tapped to provide access for

mounting a piezoelectric dynamic pressure transducer (Kistler Model 603B1), a pressure gage, two Ktype thermocouples (Omega CASS-18E-12 and KMQSS-125U-6), and a needle valve (Whitey

SS-1RS4) for agent filling and subsequent nitrogen pressurization. Occasionally, a fast response static pressure transducer (Druck Model PDCR 330) was also mounted to the vessel in order to validate the

readings from the dynamic pressure transducer. A flange for mounting the release mechanism was

welded to the other end of the tube. The internal volume of the vessel without the release mechanism

in place was found to be 4.97 x lO"'^ m-^, determined by the volume of water required to fill the

vessel. Although the vessel was hydrostatically tested to 17.2 MPa, it was never operated above

1 1 MPa for safety reasons. A schematic of the vessel is shown in Figure 7.
440 8. OPTIMIZATION OF SYSTEM DISCHARGE

Pressure transducer Needle valve for agent filling

and nitrogen pressurization

Access holes

for thermocouples

\andnitrog(

pVilVi II II

240 mm

J Mil II II 'III U.

50.8 mm

Threads for mounting

Marotta, HTL, or Kidde-Graviner valve

Figure 7. Schematic of the cylindrical vessel.
8. OPTIMIZATION OF SYSTEM DISCHARGE 441

The spherical vessel was custom-made for the National Institute of Standards and Technology by

Pacific Scientific HTL Kin-Tech Division. The vessel was made of metal alloy with an internal

volume of 5.38 x 10"^ m^. There was a three-outlet manifold welded to the top of the vessel. The

three access holes were for mounting a K-type thermocouple to monitor the liquid temperature, a dynamic or static pressure transducer, and a needle valve. The vessel was also equipped with a one- way fill valve (not used in this study). At 21 °C, the vessel had a proof pressure of 12.4 MPa and a

burst pressure of 18.6 MPa. Figure 8 shows the schematic of the spherical vessel. The cylindrical vessel with sight gages was an industrial liquid level gage (Jerguson 14TL10). It was made of stainless steel SS 316 and had two opposing slots on the side wall of the vessel. The

slots were machined to provide plane seats for the gaskets and the glass assemblies. The gage glasses were specially processed, heat treated, and tempered. The glass assemblies were secured in place with

steel covers and specially designed bolts. There were a 3/4 NPT top end connection for mounting

instrumentation and a 600# RF slip-on flanged bottom connection for mounting the release mechanism.

A cross was attached to the top end connection to provide three access holes for a dynamic pressure

transducer, a pressure gage, and a needle valve. The vessel was hydrostatically tested by the manufacturer to 10.2 MPa at 38 'C. The internal volume of the vessel without the release mechanism was

estimated to be 6.10 x 10"^ m'^. The schematic of the vessel is shown in Figure 9. Four types of release mechanisms were used in this study: two solenoid valves (from Marotta

Scientific Model MV121KJ-2 and HTL Pacific Scientific PN 36400036), a pyrotechnic piston-actuated

valve (Kidde-Graviner, Jordan Crew Bay System), and a squib (HTL Pacific Scientific). There are some slight differences in the operating procedures for the two solenoid valves. Upon activation, the

Marotta valve with an outlet of 4.45 cm remains open and requires re-arming if it is to be used again,

whereas the HTL valve, which has an outlet size similar to the Marotta valve, will close automatically

and no re-arming of the valve is needed. Note that both valves have a smaller passage upstream of

the outlet. The activation of the Kidde-Graviner valve with a straight-through opening of 4 cm is achieved by using a fast acting beam assembly held initially in position by a spindle. A pyrotechnic

piston actuator, when it is energized, causes the spindle to rotate and free the beam assembly. Rearming of the valve is also required. The squib assembly consists of an explosive cartridge

(HTL PN 30903827), a rupture disc, and a discharge outlet port (Military Standard MS33649-08) with

a diameter of 2 cm. Opening of the rupture disc is initiated by the explosive cartridge. The squib was

used solely with the spherical vessel. A DC power source was used to activate the release mechanism.

Approximately 25 V and 10 A were required for the Marotta and HTL solenoid valves, approximately

10 V and 1 A for the piston-actuated Kidde-Graviner valve, and approximately 20 V and 3.5 A for the

HTL squib. The response times of all the release mechanisms were on the order of 10 ms. The experimental preparation involved the following steps. The vessel was filled approximately

two-thirds full (by volume); in some experiments a half-filled condition was used. The total mass of

the agent dispensed to the vessel was obtained by weighing on an electronic scale with an uncertainty

of 1 g. The vessel was then pressurized with nitrogen to a specified equilibrium pressure at room

temperature. The attainment of the final equilibrium pressure could be facilitated by shaking the

vessel intermittently and vigorously and by repetitive nitrogen pressurization. The vessel was then

mounted and was ready to be discharged at room temperature.

Figure 10 shows the experimental set-up for the vessel with a Marotta or HTL solenoid valve.

Because of the configuration of the valve outlet, the normal discharge orientation is horizontal in this case. For the Kidde-Graviner valve and the HTL squib, the normal discharge direction is vertically

downward. The five lasers and photodiode detectors shown in Figure 10 were used to measure the

average dispersion velocities of the leading edge of the evaporating spray propagating downstream

from the vessel exit. The measurement technique was based on the principle of laser attenuation,

details of which can be found in the previous report by Grosshandler et al. (1994). A piezoelectric
442 8. OPTIMIZATION OF SYSTEM DISCHARGE

Pressure

transducer

Figure 8. Schematic of the spherical vessel.
OPTIMIZATION OF SYSTEM DISCHARGE

Dynamic pressure

transducer

re

Needle valve

23.5 cm

1

5.08 cm O

O

O

O

O

O

o

o

o

\

Static pressure

transducer (optional)

Sight gage

Release mechanism

Figure 9. Schematic of the cylindrical vessel with sight gages.
444 8. OPTIMIZATION OF SYSTEM DISCHARGE

Lasers

Dynamic

pressure transducer

Pressure

gauge

Thermocouple

Needle iSIIq' Pressure

valve i" U D n. transducer

Vessel

/

Valve

(Marotta or HTL)

Figure 10. A schematic of the experimental set-up.
8. OPTIMIZATION OF SYSTEM DISCHARGE 445

dynamic pressure transducer (Kistler Model 603B1) oriented normal to the flow direction was placed

along the centerline of the valve outlet at a downstream position of approximately 1 .4 m and was used

to measure the impact pressure and to obtain some qualitative two-phase behavior of the flashing

spray. The mounting of the pressure transducer was also discussed in detail in Grosshandler et al. (1994).

All the piezoelectric dynamic pressure transducers were regulated by dual mode charge amplifiers

(Kistler Model 5004) with 180 kHz frequency filters. The amplifiers were operated in the "long" time

constant mode in order to track the temporal variation of pressure over the relatively long discharge

time interval (on the order of 100 ms).

The outputs from the amplifiers, the static pressure transducer (when used), and the photodiode

detectors were recorded using a 16 channel, 12 bit, high-speed data acquisition board (Strawberry Tree

FLASH- 12TM Model 1) at a rate of 25 kHz per channel for 0.6 s. The total number of samples

collected per channel was 15000. Because of the speed and memory size, a 1 Mbyte daughter board was required for direct memory access. The data acquisition system was controlled by menu-driven

software (WorkBench 4.0), and the data obtained were stored in a personal computer (PC-486) for

subsequent data analysis.

A high-speed movie camera (Photec IV) operating at 2000 frames per second was used to document the flashing behavior of agent/nitrogen mixtures at the exit of the vessel and the events

occurring inside the vessel during discharge when the cylindrical vessel with sight gages was used.

Kodak Ektachrome high-speed daylight films (ASA 400) together with two flood lights for front

lighting and an intense flood light for backlighting were used.

The experimental sequence was controlled by a timing circuit. At the initiation of the timing

sequence, the high-speed camera was first triggered to ensure that the full framing rate had been

attained before photographing the discharge process, and the camera was on for 2 s. At 0.8 s, the data

acquisition was initiated for a duration of 0.6 s. At 1 s, the timing circuit sent a control signal to turn on the DC power supply for 0.2 s to activate the release mechanism.

In this section, unless otherwise stated, all the discharge experiments were conducted at room

temperature with a nitrogen charge pressure of - 4.12 MPa. The vessel was initially filled with an amount of pure liquid agent equivalent to approximately two-thirds of the vessel volume. The

discharge orientation was the normal configuration of the release mechanism, as discussed above. This

condition is termed standard discharge in the discussion.

The effects of several experimental parameters on the discharge characteristics were also

examined. These parameters included: (1) discharge orifice size, (2) initial amount of agent, (3)

nitrogen charge pressure, (4) initial temperature of agent, (5) discharge orientation, and (6) degree of

nitrogen saturation in the liquid agent. Table 16 shows the experimental matrix.

The effect of the orifice size on the standard discharge process was evaluated by mounting a

stainless steel orifice plate at the Marotta or HTL solenoid valve outlet. The orifice diameter

(2.38 cm) used in this work was about half of the valve outlet diameter.

The initial amount of agent in the vessel was also varied from two-thirds fill (used in all the

standard discharge tests) to half fill. Only two agents (CF3I and HFC- 125) were used in this part of

the parametric studies.

The effect of initial nitrogen charge pressure on the discharge was examined by performing a

series of experiments at room temperature with ~ 2.75 MPa, 4.12 MPa, and 5.50 MPa nitrogen

pressures. Because the operating pressures of the Marotta and HTL valves were rated to be less than

5.50 MPa, only two charge pressures (2.75 MPa and 4.12 MPa) were used, whereas all three pressures were evaluated using the spherical vessel equipped with a squib.

Some experiments were performed to determine the effect of initial agent temperature on the

discharge characteristics. In the low temperature experiments, the vessel was initially filled with agent,
446 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 16. Test matrix for discharge experiments

CF3I FC-218 HFC- 125 CF,Br

HTL valve'' /

HTL valve'' HTL valve^'^

Marotta valve^ / i

Marotta valve''

Marotta valve^'^

Marotta vaive"'*^ /

Marotta valve'''^ /

KG valve' / /

KG valveS / /

Squib'' /

Squib'' /

Squib*" /

Squib'

Squib"

'initial nitrogen charge pressure -4.12 MPa

''two-third fill condition

"^half fill condition

''initial nitrogen charge pressure ~ 2.75 MPa

^with an orifice plate at the valve outlet 'cylindrical vessel with sight gages

^high temperature experiments

""initial nitrogen charge pressure - 5.50 MPa

'cold temperature experiments

^effect of discharge orientation

pressurized with nitrogen to - 4.12 MPa at room temperature, and then cooled with dry ice to below - 50 °C before a discharge was initiated. Due to the large thermal inertia of the cylindrical vessel and

the limitations for the operating temperatures of the Marotta and HTL solenoid valves, only the

spherical vessel with a squib was used in this set of experiments. In the high temperature experiments,

the vessel was heated with a heating tape wrapped around it after it was filled with agent and

pressurized with nitrogen to ~ 4.12 MPa at room temperature. Depending on which agent was used in the test, the vessel was heated to a temperature (90 °C for CF3I) at which the corresponding pressure

inside the vessel did not go beyond 9.62 MPa; this was considered to be the maximum operating

pressure of the closed cylindrical vessel. Because of the limitation of the operating pressure, the

cylindrical vessel with sight gages and the spherical vessel were not used in this study. In addition,

only the Kidde-Graviner valve was used due to the fact that the final attainable pressure at the elevated

temperature was above the operating pressure of the Marotta and HTL solenoid valves. Note that,

since no shaking or agitating of the vessel was performed during the cooling or heating process, the
8. OPTIMIZATION OF SYSTEM DISCHARGE 447

liquid and gas phases might not be in equilibrium before discharge. This condition is similar to that of

an actual fire extinguishing vessel in which no mixing mechanism to maintain equilibrium between the

liquid and vapor phases exists at cold or hot ambience.

The study of the effect of orientation on the discharge dynamics was conducted at room

temperature with a nitrogen charge pressure of ~ 4.12 MPa by using the spherical vessel and varying

the discharge direction from vertically downward (standard discharge) to vertically upward, horizontal,

and downward with an angle of 30° or 45° with the vertical. The spherical vessel was chosen because

it was lightweight and easy to mount. Only HFC- 125 was used in this parametric study. In order to shed some light on the effect of the degree of nitrogen saturation on the discharge

process, a limited set of experiments was performed by partially saturated the liquid agent with

nitrogen. Partial saturation was achieved by simply bubbling nitrogen through the liquid agent to a

specified pressure without attaining the final equilibrium pressure and without agitating the vessel. In other words, if one were to agitate the vessel, an equilibrium pressure lower than the specified pressure

would be obtained because more nitrogen would be dissolved in the liquid agent.

8.4.3 Results and Discussion

8.4.3.1 General Visual Observations. The events occurring inside the vessel during a discharge

of CFjBr were conceptually described by Elliot et al. (1984) without any confirmation from visual

observations. Based on the experimental pressure decay curves during discharge and a simple

theoretical flow model, he suggested that there were several sequential events occurring inside the

vessel. When the release mechanism was activated, there was a short duration during which the pressure continuously decreased, and the liquid agent/nitrogen mixture was expelled as a superheated

liquid through the valve exit. Nitrogen remained dissolved in the liquid agent as a non-equilibrium

supersaturated solution, and the liquid/nitrogen mixture remained a clear solution. As the pressure

inside the vessel dropped further, degassing of nitrogen in the solution occurred, and nitrogen bubbles

started to form. Once the bubbles nucleated, the dissolved nitrogen quickly came out of the solution

into the bubbles causing the bubbles to grow. Due to rapid bubble expansion, the liquid level swelled,

and the ullage was compressed to a higher pressure which was noted as a short transient pressure

recovery in the pressure decay history. The pressure then began to decrease again until an inflection

point in the pressure decay curve was reached. A low quality (defined as the fraction of the mass flow

rate which is gas) bubbly two-phase CFjBr/nitrogen mixture was assumed to be discharged from the

vessel during this time interval. The inflection point signified the depletion of the low quality twophase mixture and the initiation of the discharge of the ullage gas mixture (CF3Br vapor and nitrogen)

from the vessel. The rate of pressure decay increased because the gas mixture was less dense than the

low quality bubbly two-phase mixture and experienced a higher volumetric discharge rate. The visual observations made in this study by using the cylindrical vessel with sight gages and . high-speed photography revealed the following phenomena which were slightly different from the

description given by Elliot et al. (1984). For CFjBr, HFC-125, and FC-218, the liquid agent/nitrogen

mixture remained clear for a short duration (~ 5 ms) after the initiation of the discharge. This duration was found to be agent dependent, with CF3Br the shortest and FC-218 the longest. During this period,

the receding liquid/vapor interface was clearly visible. The interior of the vessel as seen through the

sight gages then became completely foggy, and the liquid/vapor interface was no longer visible. It was not possible to determine from the movies whether the fogginess prevented the observation of the

liquid/vapor interface or the degassing of the dissolved nitrogen induced frothing of the liquid mixture

throughout the vessel, thus causing the liquid/vapor interface to disappear. The fogginess was

sustained for a period of time during which the liquid mixture was assumed to be completely
448 8. OPTIMIZATION OF SYSTEM DISCHARGE

discharged. The sight gages on the vessel then became clear once again. This fogginess is conjectured to be due both to vapor condensation as a result of cooling by the adiabatic expansion of the

ullage vapor, and the evolution of the dissolved nitrogen as bubbles in the liquid mixture. The

existence of the nitrogen bubbles could not be visually confirmed from the high speed movies due to the limited spatial resolution of the pictures. For CF3I, the receding liquid/vapor interface and the

liquid mixture remained clear for a relatively long time (at least half of the liquid had been depleted)

after the initiation of the discharge. The fogginess then appeared throughout the vessel, and the

liquid/vapor interface vanished. The interior of the vessel became clear again after a period of

fogginess. It was postulated that nitrogen degassing did not occur in the CFjI/nitrogen mixture (see

discussion in Section 8.4.3.2). Selected photographic sequences of events for CFgBr, HFC-125,

FC-218, and CF3I are given in Figures 11, 12, 13, and 14 respectively.

The flashing behavior of CFjBr external to the vessel was also noted by Elliot et al. (1984).

They reported that to the eye the discharge was a puff of white vapor. Their high speed movies

showed an opaque white plume that steadily diminished in size but did not change in density. The

appearance of the plume did not seem to change, with either front lighting or back lighting, even when

the vapor mixture discharge began because a white cloud continued to form due to vapor condensation. The observations obtained in the current study were very similar to those made by Elliot et al. (1984) although the visualization of the process using the naked eye was almost impossible because

the discharge time was much shorter in this work. Due to the difficulty in discerning the transition

from the liquid discharge to the remaining vapor discharge from the appearance of the spray at the

valve exit, the time for the depletion of the liquid from the vessel could not be determined with great

certainty from the high-speed movies. In addition, even when the vessel with sight gages was used,

the fact that the liquid/vapor interface inside the vessel was not visible during the course of the

discharge also hindered the determination of the liquid depletion time. Figures 15, 16, 17, and 18,

respectively, show individual photographic shots taken at the same instant (5 ms after the first appearance of the spray at the valve exit) during discharges of CFgBr, HFC-125, FC-218, and CF3I

from the cylindrical vessel equipped with a Marotta solenoid valve. When other release mechanisms

were used, the general appearances of the flashing sprays (other than the spray angles) were also

observed to be very similar to Figures 15 to 18.

8.4.3.2 Marotta Solenoid Valve. This subsection discusses the experimental results obtained

using a Marotta solenoid valve. Figure 19 through Figure 22 show respectively the pressure decay

curves obtained inside the vessel during standard discharges of CFjBr, HFC-125, FC-218, and CF3I.

Pressures reported in the figures are gauge pressures. The pressure is nondimensionalized by the

initial pressure {P^ in the vessel before the initiation of a discharge. The time f = 0 s is defined as the

time when the solenoid valve is first opened partially. This time, with an estimated uncertainty of

0.08 ms, is inferred from the instant when the initial vessel pressure starts to decrease by 34 kPa and

continues to decrease. Note that the baseline fluctuation of the initial vessel pressure is less than

34 kPa peak-to-peak.

For the CFjBr (Figure 19) standard discharges, the pressure decay curves show four regions as suggested by Elliot et al. (1984). The first corresponds to the discharge of metastable liquid

agent/dissolved nitrogen mixture, followed by a sudden pressure recovery which presumably implies

the degassing of nitrogen from the liquid mixture. The third region corresponds to the discharge of a low quality two-phase bubbly mixture, and the last region signifies the discharge of the remaining

vapor mixture from the vessel. In Figure 19, the transition from the third to the fourth region in one

case {Pi = 4.05 MPa) is not very distinct. For the HFC-125 and FC-218 tests, the aforementioned four regions are clearly shown in Figures 20 and 21 respectively. However, it is unclear, at present, why the rate of pressure decay in
8. OPTIMIZATION OF SYSTEM DISCHARGE 449

t = 5.0 ms t = 7.5 ms

Figure 11. Photographic sequence of events as seen through the sight gages of the cylindrical

vessel during a discharge of CFjBr.
450 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 12 Photographic sequence of events as seen through the sight gages of the cylindrical

vessel during a discharge of HFC- 125.
8. OPTIMIZATION OF SYSTEM DISCHARGE 451

Figure 13. Photographic sequence of events as seen through the sight gages of the cylindrical

vessel during a discharge of FC-218.
452 8. OPTIMIZATION OF SYSTEM DISCHARGE

t = 5.0 ms / = 7.5 ms

Figure 14. Photographic sequence of events as seen through the sight gages of the cylindrical

vessel during a discharge of CF3I.
8. OPTIMIZATION OF SYSTEM DISCHARGE 453

Figure 15. A snapshot of CF3Br discharge (taken 5 ms after the first appearance of the spray at

the valve exit).
Figure 16.

A snapshot of HFC- 125 discharge (taken 5 ms after the first appearance of the spray

at the valve exit).
OPTIMIZATION OF SYSTEM DISCHARGE 455

Figure 17. A snapshot of FC-218 discharge (taken 5 ms after the first appearance of the spray at the valve exit).
456 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 18. A snapshot of CF3I discharge (taken 5 ms after the first appearance of the spray at the valve exit).
OPTIMIZATION OF SYSTEM DISCHARGE 457

Figure 19. Pressure decay curves during standard discharges of CFgBr using a Marotta valve.
458 8. OPTIMIZATION OF SYSTEM DISCHARGE

0 25 50 75 100 125

Time (ms)

Figure 20. Pressure decay curves during standard discharges of HFC- 125 using a Marotta valve.
8. OPTIMIZATION OF SYSTEM DISCHARGE 459

Figure 21. Pressure decay curves during standard discharges of FC-218 using a Marotta valve.
460 8. OPTIMIZATION OF SYSTEM DISCHARGE

0 25 50 75 100 125 150

Time (ms)

Figure 22. Pressure decay curves during standard discharges of CF3I using a Marotta valve.
8. OPTIMIZATION OF SYSTEM DISCHARGE 461

the fourth region is slower than that in the third region, which is opposite to the trend observed in the

CFjBr tests. For the CF3I experiments, the pressure recovery before the third region is not noted in Figure 22.

In addition, there is a pressure recovery in the fourth region. This pressure recovery was subsequently

found to be caused by some artifact from the dynamic pressure transducer. This is substantiated by

Figure 23 which compares the two pressure decay curves obtained using two different pressure

transducers, one static and one dynamic, in the same test. Note that this pressure recovery artifact was

only observed in the CF3I tests. If the pressure recovery is assumed to be related to the degassing of the dissolved nitrogen in the

liquid agent and if the degassing process can presumably be explained in terms of homogeneous

nucleation theory and can be modeled by an isothermal decompression process, then the homogeneous

nucleation pressure, according to the work by Forest and Ward (1977), should depend on the initial pressure, the surface tension of the liquid mixture, and the vapor pressure of the agent and the

solubility of nitrogen in the liquid agent at the temperature of interest. The homogeneous nucleation

pressure, which is the pressure at which homogeneous nucleation occurs, can be estimated by the

following set of equations whose derivations were given in detail by Forest and Ward (1977). To

simplify the analysis, it is further assumed that the effect of pressure and concentration of the

dissolved nitrogen on the vapor pressure of the liquid is negligible and that the gas mixture inside the

bubble is ideal. It is expected that under the experimental conditions used in this study, a very good

estimate of the homogeneous nucleation pressure can be obtained by invoking the above assumptions

and without resorting to the detailed calculations. The homogeneous nucleation pressure (P/^) is given

by

CP 1/2 P +

sat (19)

3kT]n{Z/J)

with

In Cs

= ]nH + (20)

and

fn = ^/.exp (21)
462 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 23. Comparison of the pressure decay curve obtained by a dynamic pressure transducer

with that obtained by a static pressure transducer.
8. OPTIMIZATION OF SYSTEM DISCHARGE 463

where P^^, is the vapor pressure of the agent at temperature T, C is the ratio of moles of nitrogen

dissolved to moles of liquid agent, is the concentration of nitrogen which would be present in the

liquid agent if the pressure of the gas above the liquid were at the total pressure P^, a is the surface

tension, Z (cm"^ s"^) is the rate constant, J (nuclei s~' cm~^) is the nucleation rate per volume, k is the

Boltzmann constant (= 1.38048 x 10~^^ dyne cm K~'),/„'' is the fugacity of the pure gas, H is the

Henry's constant, is the partial molar volume of the dissolved nitrogen, 5,^^ is the second virial coefficient, and R is the universal gas constant. The mole ratio C can be determined from the solubility (mole fraction) of nitrogen in the liquid

agent at T and by the following equation:

C =

^" (22)

where is the solubility of nitrogen and can be calculated using the vapor-liquid equilibrium

computer code (PROFISSY).

Following the procedure outlined in Forest and Ward (1977), Henry's constant H can be estimated

using Equation (19) without the last term and Equations (20) and (21), with replaced by P- in all three equations.

The surface tension for pure agent was estimated by the method of Brock and Bird (Reid et al., 1987); the effect of the dissolved nitrogen on the surface tension was assumed to be negligible. The

vapor pressure was estimated by the method proposed by Gomez-Nieto and Thodos (Reid et al., 1987). The partial molar volume of the dissolved nitrogen was calculated using the method in Chueh

and Prausnitz (1967). The second virial coefficient B^^ was obtained from the tabulation in Dymond

and Smith (1969). For all the agents used in this study, the rate constant Z was estimated by the

method described in Volmer (1945) and found to be on the order of 10 cm s . For the nucleation

rate J, a value of 10"* nuclei s~^ cm~^ was used. The value of the pressure determined was found to be very insensitive to the value of J chosen because J is in the logarithmic term in Equation (19). The

determination of P^ requires an iteration procedure because of the nonlinear nature of Equation (19).

The Newton-Raphson numerical scheme (Carnahan et al., 1969) was used to obtain Pf^. The initial pressure was used as an initial guess for Pf^ in the iteration. Table 17 shows the calculated results based on the experimental conditions used to obtain Figures 19 to 22. The calculated homogeneous

nucleation pressures for CFjI/nitrogen mixtures were found to be negative (i.e., in tension); negative

homogeneous nucleation pressures are possible according to Reid (1976) although at present no

experimental data on CF3I/nitrogen mixtures are available to substantiate the predictions.

The pressure ratio (P/P,) at which the degassing of nitrogen initiates or not can now be qualitatively explained in terms of the aforementioned, simplified homogeneous nucleation theory. Based on

the calculated homogeneous nucleation pressures in Table 17, the degassing of nitrogen from

CFjI/nitrogen liquid mixtures, which was not observed experimentally as shown in Figure 22, was

predicted not to occur under the same experimental conditions. No attempt was made to compare

quantitatively the calculated homogeneous nucleation pressures to the experimental degassing pressures because of the uncertainties associated with the thermophysical property predictions used in the

calculations. It should be noted that although homogeneous nucleation theory appears to provide some

qualitative explanations with regard to the degassing process, heterogeneous bubble nucleation (a more

complicated process) from the nucleation sites on the vessel wall could also play a role. However, it can be argued that the ease with which the agents wet the inner surface of the container due to their

small contact angles (low surface tensions) during vessel filling favors homogeneous nucleation
464 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 17. Calculated homogeneous nucleation pressure (Pf^) at T

Agent Mass (g) T{K) P, (MPa) Ph (MPa)

CF3Br 591 295 4.10 1.90 0.46

CFjBr 564 295 4.05 1.85 0.46

HFC- 125 438 294 4.03 1.63 0.41

HFC- 125 428 294 4.16 1.76 0.42

FC-218 450 295 4.04 2.12 0.53

FC-218 435 294 4.10 2.11 0.52

CF3I 723 294 4.12

CF3I 719 294 4.21

CF3I 755 294 4.17

because less effective vapor trapping in the irregular cavities on the imperfect container wall surface

will be expected (Carey, 1992).

A comparison of the pressure decay curves for all the agents is made in Figure 24. It is clear that different agents exhibit different discharge rate and degassing characteristics. Based on the pressure

decay curves, CF3I discharges the fastest and CFjBr the slowest under similar initial conditions.

These observations were also confirmed qualitatively by the high-speed movies.

The average dispersion velocities of the spray obtained by the laser attenuation technique are summarized in Figure 25. The dispersion of CF3I has the slowest average downstream velocities. All

the other agents have similar but higher spray dispersion velocities than CF3I.

The measurements from the downstream dynamic pressure transducer can be used to provide a

qualitative description of the two-phase behavior of the spray downstream. According to Whalley

(1990), if the two-phase flow is assumed to be homogeneous, then the momentum flux M, which in this case is the measured dynamic pressure, is given by

M = = (23)

with

p, = ap + (1 - a)p, (24)

where G is the mass flux, u is the average velocity, a is the void fraction, p^^ is the homogeneous

mixture density, and and p^ are the saturated vapor and liquid densities. The void faction is defined
OPTIMIZATION OF SYSTEM DISCHARGE 465

Figure 24. Comparison of pressure decay curves during standard discharges using a Marotta

valve.
466 8. OPTIMIZATION OF SYSTEM DISCHARGE

C/3

o

200

180

160

140

120

^ 100

>

CD^ 80

M

(D

< 60

. 40

20

0

o CF3Br (Pj = 4. 10 MPa)

CFjBr (Pj = 4.05 MPa)

HFC-125 (Pj = 4.03 MPa)

HFC-125 (Pj = 4.16 MPa)

FC-218 (Pj = 4.04 MPa)

FC-218 (P. = 4. 10 MPa)

I I

Marotta

CF3I (P. = 4. 12 MPa)

CF3I (P. = 4.21 MPa)

CF3I (Pj = 4. 17 MPa)

CFJ (P. = 4. 12 MPa)

±

0 20 40 60 80 100 120

Average downstream distance (cm)

Figure 25. Average downstream velocities during standard discharges using a Marotta valve.
8. OPTIMIZATION OF SYSTEM DISCHARGE 467

as the average fraction of the volume which is occupied by the gas phase. Using Equations (23) and

(24), one can solve for a,

..2 _ V (25)

Since M and u were measured in the experiments, an estimate of a could be made. Table 18

summarizes the calculated results. The saturated vapor and liquid densities were assumed to be those

of the pure agents at their normal boiling points. For CFjBr, HFC-125, and FC-218, the saturated

densities were obtained from the REFPROP database (Gallagher et ai, 1993). For CF3I, the saturated

liquid density was estimated using the modified Rackett method described in Reid et al. (1987), and

the vapor density was calculated by assuming the vapor was ideal. The average velocities u were

obtained by taking the ratio of the distance between the transducer and the valve exit to the time when

the spray hit the transducer. The dynamic pressures (Af) used in the calculations corresponded to the

maximum pressures recorded in the experiments. Among all the agents, CF3I has the lowest estimated

void fraction, which implies that a fair amount of CF3I has not been evaporated at this downstream

location. The low Jakob number (see Grosshandler et al., 1994) of CF3I and its high normal boiling

point also provide some qualitative support of the calculated low void fraction. Since the other three

agents have higher Jakob numbers and lower boiling points than CF3I, it is not surprising to find

higher calculated void fractions for all three agents.

The maximum angles sustained by the sprays at the valve exit during discharges were also

measured and are tabulated in Table 18. Note that the spray angle is a function of time during

discharge. Since it was very difficult to define clearly the boundaries of the sprays because of their

o flashing characteristics, an uncertainty of 4 was typical. Compared to the other agents, the maximum

spray angles in the selected CF3I tests were found to be the largest.

8.4.3.3 HTL Solenoid Valve. For comparison purposes, selected pressure decay curves obtained

during standard discharges of CF3Br, HFC-125, FC-218, and CF3I using the HTL solenoid valve are shown in Figure 26. The initial, relatively flat plateaus in the figure suggest that the opening times of

the solenoid valve are slow and differ, depending on which agent is used. However, no general trends were observed when all of the experimental data were examined although Figure 26 shows the opening

time of the valve for CF3I is the slowest. The initial, slow, and continuous decrease in pressure could

not be attributed to the drift of the baseline initial pressure P- before the opening of the valve because

the baseline value for P- remained relatively constant except for some small fluctuations. Note that such initial slow pressure decays were not observed in all the tests using a Marotta valve.

Similar to the CF3I discharges using a Marotta valve, the pressure recovery artifact in the CF3I

pressure trace caused by the dynamic transducer is also noted in Figure 26, and no degassing of CF3I

is suggested. If the opening time of the valve were to be ignored, CF3I would discharge the fastest,

and CF3Br the slowest under similar initial conditions.

Figure 27 summarizes the average downstream dispersion velocities of the sprays for all the

agents. Similar to the observations made using a Marotta valve, the dispersion of CF3I exhibits the

slowest downstream velocities. All the other agents have similar but higher spray dispersion velocities

than CF3I.
468 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 18. Calculated void fractions and measured maximum spray angles (standard discharges using

a Marotta valve)

Agent Mass (g) u (m/s) M (MPa) max a Maximum

spray angle

CFjBr 591 103 0.30 0.99 100°

CFjBr 564 ...* ... ... —

HFC- 125 438 114 0.84 0.96 110°

HFC- 125 428 104 0.41 0.98

FC-218 450 102 0.43 0.98 108°

FC-218 435 94 0.25 0.99

CF3I 723 64 1.58 0.84 142°

CF3I 719 64 1.65 0.83 141°

CF3I 755 61 1.68 0.81

*not recorded

Table 19 tabulates the calculated void fractions at the location of the downstream dynamic

pressure transducer. Similar to the results for the Marotta valve, the void fraction for CF3I is estimated to be the smallest, whereas all the other agents show comparable and higher void fraction

values. The measured maximum spray angles are also summarized in Table 19. The CF3I spray

angles still appear to be the largest.

8.4,3.4 Kidde-Graviner Valve The results of the standard discharges of all the agents for this valve are shown in Figure 28. The pressure recovery artifact occurring close to the end of the

discharge in the CF3I measurements is again noted in the figure. According to the pressure decay

curves, the discharge rate of CF3I is the fastest and CF3Br the slowest. The pressure trace also

suggests that the degassing of dissolved nitrogen does not occur in the CF3I test, consistent with the

prediction using the homogeneous nucleation theory (see Table 17). For the other three agents, the

pressure recoveries due to the degassing of dissolved nitrogen are not very pronounced, that is the

pressures do not rebound significantly. Since this valve has a bigger opening, one possible explanation for these flat pressure plateaus is that the rate of the ullage compression due to the degassing is offset by the rate of ullage expansion resulting from the faster liquid discharge.

The average downstream dispersion velocities are summarized in Figure 29. Once again, the

dispersion velocities of CF3I are the slowest among the other three agents. The high temperature

results in the figure will be discussed later in Section 8.4.3.9. The calculated void fractions at the location of the downstream pressure transducer, together with

the measured maximum spray angles, are tabulated in Table 20. Care should be exercised in interpreting the calculated void fraction for HFC- 125 because the unusual high M, which results in a low void fraction, could have been caused by the accidental impact of the flying debris (from the

actuation of the pyrotechnic piston) on the pressure transducer; the debris was frequently observed

coming from the valve exit in the high-speed movies. Excluding the result of HFC-125, CF3I again
OPTIMIZATION OF SYSTEM DISCHARGE 469

Figure 26. Comparison of pressure decay curves during standard discharges using a HTL

solenoid valve.
470 8. OPTIMIZATION OF SYSTEM DISCHARGE

200

Average downstream distance (cm)

Figure 27. Average downstream velocities during standard discharges using a HTL valve.
8. OPTIMIZATION OF SYSTEM DISCHARGE 471

Table 19. Calculated void fractions and measured maximum spray angles (standard discharges using

a HTL valve)

Agent Mass (g) u (m/s) M^^ (MPa) a Maximum

spray angle

CFjBr 586 91 0.83 0.95 103°

HFC- 125 431 98 0.58 0.97 105°

HFC- 125 426 98 0.65 0.96 105°

FC-218 444 83 0.55 0.95 108°

CF3I 707 46 1.65 0.67 115°

CF3I 713 51 1.67 0.73 120°

shows the lowest calculated void fraction at a fixed downstream location. In addition, CF3I spray has

the largest maximum spray angle at the valve exit.

8.4.3.5 HTL Squibs. Figure 30 summarizes the results of the standard discharges using HTL

squibs. Three salient points are noted in the figure. First, no degassing of the dissolved nitrogen in CF3I is suggested. Second, the inflection points in the pressure decay curves of CFjBr, HFC- 125, and

FC-218 are barely distinguishable. Third, the initial CF3I discharge rate is the fastest, and CFjBr has

the slowest discharge rate. The dynamic pressure transducer still caused a pressure recovery artifact close to the end of the discharge in some of the CF3I tests; however, the pressure rebound was not as pronounced as in the tests using Marotta, HTL, and Kidde-Graviner valves. All of the pressure curves have much more fluctuation during the initial discharge period; this could be the result of the violent

vibration caused by the activation of the explosive cartridge.

The measurements of average spray dispersion velocities could not be carried out because the

brightness of the flash caused by the explosion of the cartridge completely saturated the laser detectors. However, estimates of the average velocities could still be made using the time at which the

downstream pressure transducer detected the arrival of the spray. Table 21 summarizes the average

velocities, the calculated void fractions, and the measured maximum spray angles. The average

dispersion velocity u of the CF3I spray is still the slowest. In Table 21, the unusual high values of M

could again be due to the flying debris (resulting from the action of the explosive cartridge) hitting the pressure transducer. The measured maximum spray angles of all the agents are very similar within the measurement uncertainty. Based on the discussion so far, the spray angle appears to be dependent not only on the agent considered but also on the release mechanism used.

8.4.3.6 Effect of an Orifice Plate. The pressure decay curves obtained using an orifice plate

and a Marotta valve during standard discharges are respectively shown in Figures 31, 32, 33, and 34

for CF3Br, HFC-125, FC-218, and CF3I. The low and half-fill measurements in the figures will be

discussed in later sections. There are three important features in these figures.
472 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 28. Comparison of pressure decay curves during standard discharges using a KiddeGraviner (KG) valve.
OPTIMIZATION OF SYSTEM DISCHARGE 473

200 T

180

160

140

120

100

80

60

40

20

KG

73 °C

O CF3Br(Pi = 4.14MPa) • CF3Br {?. =9.24 MPa)

V HFC-125 (Pj = 4.13 MPa)

T HFC-125 (Pj = 9.10 MPa)

HFC-125 (Pj = 9.49 MPa) @ 65 °C

70 °C

FC-218 (Pj = 4.14 MPa)

A FC-218 (Pj = 8.58 MPa) @ 75 "C

A FC-218 (Pj = 8.81 MPa)

O CF3I (Pj = 4. 15 MPa)

CF3I (Pj = 8.64 MPa) @ 90 °C

77 °C

0

I 20 40 60 80 100 120 140

Average downstream distance (cm)

Figure 29. Average downstream velocities during standard discharges using a Kidde-Graviner

valve.
474 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 20. Calculated void fractions and measured maximum spray angles (standard discharges using

a Kidde-Graviner valve)

Agent Mass (g) u (m/s) M„„ (MPa) a Maximum

spray angle

CF^Br 625 89 1.28 0.92 132°

HFC- 125 497 92 2.20 0.83 145°

FC-218 492 75 1.18 0.87 134°

CF3I 800 54 1.21 0.83 152°

1. The pressures at which the degassing occurs (see Figures 31, 32, and 33) do not appear to depend on the presence or absence of the orifice plate under similar initial conditions of the

vessel. This is consistent with the calculations of the homogeneous nucleation pressures which

depend only on the initial conditions of the vessel (see Section 8.4.3.2).

2. No degassing is noticeable in the CF3I pressure decay curves. In Figure 34, the pressure

rebound at the end of the pressure decay curves is due to the artifact of the pressure trans- ducer.

3. Compared to the standard discharges, the discharge rates are much slower when a restricted

orifice plate is present.

Figure 35 compares the results of all the agents under similar initial conditions. The tendency to degas also conforms qualitatively with the predictions using homogeneous nucleation theory. The

initial discharge rate of CF3I is the fastest, whereas CF3Br again shows the slowest discharge rate. Figure 36 summarizes the results obtained from the HTL solenoid valve with the same orifice plate.

Note that CF3I was not used in this parametric study. The results are very similar to those obtained

using the Marotta valve; this is not surprising because both valves are geometrically similar. In Figure 36, it is unclear why the pressure traces of all the CFgBr runs show jagged appearances when

an orifice plate is used.

Figures 37, 38, 39, and 40 show the average spray dispersion velocities of individual agents using

the Marotta valve under various experimental conditions. Only the standard discharges with and

without an orifice plate will be discussed here. The other parametric effects will be considered in subsequent sections. In the presence of the orifice plate, the measured average velocities closest to the

exit (i.e., at 20 cm downstream) are comparable to or lower than those observed in the standard

discharges. Similar trends are noted in Figures 41, 42, and 43 when using the HTL solenoid valve.

Figure 44 compares the average velocities obtained in the standard Marotta discharges of all the agents

with the orifice plate. The average dispersion velocities of CF3I again are the slowest among all the

other agents.

Table 22 tabulates the calculated void fractions and the measured maximum spray angles. There

are no significant differences in the spray angles for the different cases. However, CF3I still has the

lowest calculated void fraction at the specified downstream location.
OPTIMIZATION OF SYSTEM DISCHARGE 475

Figure 30. Comparison of pressure decay curves during standard discharges using HTL squibs.
476 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 21. Calculated void fractions and measured maximum spray angles (standard discharges using

HTL squibs)

Agent Mass (g) u (m/s) ^n,a. (MPa) a Maximum

spray angle

CFjBr 581 93 2.03 0.89 140°

HFC- 125 454 89 2.37 0.81 131°

HFC- 125 453 95 2.61 0.81 146°

FC-218 476 83 1.72 0.84 144°

CF3I 773 62 1.52 0.84 136°

8.4.3.7 Effect of Fill Density. Figures 32 and 34 also compare respectively the discharge

characteristics of HFC- 125 and CF3I under two fill conditions, the standard (2/3) versus one-half,

using the Marotta valve. Notice the pressure rebound, an artifact, at the end of the CF3I discharge,

and the pressure decay curves of CF3I do not suggest any degassing of the dissolved nitrogen, an

observation consistent with homogeneous nucleation theory. Figure 45 shows the results for HFC- 125

from the HTL valve tests. The pressures at which the degassing occurs under two different fill conditions with similar initial nitrogen charge pressures at the same temperature appear to be not

significantly different from each other although the results in Figure 45 show some scatter. This

observation can also be explained qualitatively by homogeneous nucleation theory. According to the

theory, the bubble nucleation pressure only depends on the initial total pressure and temperature.

Since the amount of agent required is less under one-half fill condition, the overall discharge rates, based on the inflection points in the curves, are faster than the discharges under two-thirds fill condition.

The calculated void fractions and the measured spray angles are summarized in Table 23. The

calculated void fractions of CF3I are still the lowest irrespective of the fill conditions. The average

spray dispersion velocities under one-half fill conditions are also shown in Figures 38, 40, and 42.

Comparing Figures 38 and 40 for the one-half fill condition, CF3I still has lower average velocities

than HFC- 125. However, no general conclusions can be made when the average velocities under the

two fill conditions are compared at various downstream locations because of the scatter in the results.

8.4.3.8 Effect of Initial Nitrogen Charge Pressure. Visual observations of the behavior of the

agent/nitrogen mixtures inside a vessel were also made using the cylindrical vessel with sight gages

and the low initial charge pressure (~ 2.75 MPa). For the same agent, the high-speed movies revealed

similar phenomena as observed in the standard discharges, except that the liquid/vapor interface

remained visible and clear for a much longer duration.

Figures 31, 32, 33, and 34 compare the discharge characteristics of individual agents using the

Marotta valve under two initial nitrogen charge pressures. At the low initial nitrogen charge pressure,

there is no indication of degassing from the pressure decay curves using any of the agents. The

homogeneous nucleation theory, discussed above, also predicts that the degassing does not occur. Table 24 shows the calculated homogeneous nucleation pressures of all the agents based on the

experimental conditions used in the Marotta valve tests. The nucleation pressure of CF3I is not shown

in the table because it is calculated to be negative (i.e., in tension). It is clear from the pressure decay
OPTIMIZATION OF SYSTEM DISCHARGE 477

Figure 31. Pressure decay curves during discharges of CFjBr using a Marotta valve under

various conditions.
478 8. OPTIMIZATION OF SYSTEM DISCHARGE

J I I L 0 25 50 75 100 125 150

Time (ms)

Figure 32. Pressure decay curves during discharges of HFC- 125 using a Marotta valve under

various conditions.
OPTIMIZATION OF SYSTEM DISCHARGE 479

Figure 33. Pressure decay curves during discharges of FC-218 using a Marotta valve under

various conditions.
480 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 34. Pressure decay curves during discharges of CF3I using a Marotta valve under various

conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 481

0 50 100 150 200 250

Time (ms)

Figure 35. Comparison of pressure decay curves during standard discharges using a Marotta

valve with an orifice plate.
482 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 36. Comparison of pressure decay curves during standard discharges using a HTL valve

with an orifice plate.
OPTIMIZATION OF SYSTEM DISCHARGE 483

200

180

160

140

S

^ 120

o

100

80

60

40

20

0

0

CF^Br (Marotta)

o Pi = 2.76 MPa (partial N2 saturation) • Pi = 2.76 MPa (partial saturation)

V Pi = 2.75 MPa

T Pi = 4.10 MPa

Pi = 4.05 MPa

Pi = 4. 16 MPa (orifice)

20 40 60 80

Average downstream distance (cm)

100 120

Figure 37. Average downstream velocities during discharges of CFgBr using a Marotta valve

under various conditions.
484 8. OPTIMIZATION OF SYSTEM DISCHARGE

200

180

160

140 -

120 -

V. 100

80 -

60 -

40 -

20 -

0

HFC- 125 (Marotta) o Pi = 2.85 MPa (partial N2 saturation) _ • Pi = 2.68 MPa (partial N2 saturation)

V Pi = 2.66 MPa (partial saturation)

T = 4.11 MPa (partial saturation)

p.

1 = 4.16 MPa (partial saturation)

Pj = 2.74 MPa

A p. = 4.03 MPa

p. = 4.16 MPa

0 p. = 4.13 MPa (1/2 fill)

p.

1 = 4.11 MPa (orifice)

0 20 40 60 80 100 120

Average downstream distance (cm)

Figure 38. Average downstream velocities during discharges of HFC- 125 using a Marotta valve

under various conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 485

200

180

160

o p • p

V P

T P

= 2.74 MPa (partial saturation) = 2.75 MPa (partial saturation) = 4. 10 MPa (partial saturation) = 4.13 MPa (partial saturation)

FC-218 (Marotta)

o

140

120

100

>

^ 80

>

< 60

40

20

0

p

p

A P

A P

= 2.67 MPa

= 4.04 MPa

= 4. 10 MPa

= 4.02 MPa (orifice)

0 20 40 60 80 100 120

Average downstream distance (cm)

Figure 39. Average downstream velocities during discharges of FC-218 using a Marotta valve

under various conditions.
486 8. OPTIMIZATION OF SYSTEM DISCHARGE

200

180

160

140 -

120 - I,

o

o

>

^ 80 f- < 60 h

40

20

0

o P = 2.71 MPa (partial Nj saturation) • P = 2.70 MPa (partial Nj saturation)

V P = 2.78 MPa

T P = 4.12 MPa

P = 4.21 MPa

P = 4.17 MPa

A P = 4.12 MPa

P = 4. 12 MPa (orifice) P = 4.29 MPa (1/2 fill) o

CFJ (Marotta)

0 20 40 60 80 100 120

Average downstream distance (cm)

Figure 40. Average downstream velocities during discharges of CF3I using a Marotta valve

under various conditions.
OPTIMIZATION OF SYSTEM DISCHARGE 487

200

180

160

CF-Br (HTL)

140

120

-§ 100

>

0)

^ 80

>

< 60

40

20

0

0 p.

1 = 4.11 MPa (partial N2 saturation) • Pi = 4.11 MPa (partial N2 saturation)

V p.

1 = 2.76 MPa

T Pi = 4.15 MPa

Pi = 4.13 MPa

p.

1 = 4.06 MPa (orifice)

0 20 40 60 80 100 120

Average downstream distance (cm)

Figure 41. Average downstream velocities during discharges of CFgBr using a HTL valve under

various conditions.
488 8. OPTIMIZATION OF SYSTEM DISCHARGE

200

180

160

140

120

100

80

60

40

20

0

HFC- 125 (HTL)

O Pj = 4.13 MPa (partial N2 saturation) Pj = 2.78 MPa

• Pj = 4. 15 MPa (partial saturation) P- = 4.00 MPa

V Pj = 4.18 MPa (partial N2 saturation) (orifice) Pj = 3.99 MPa

A p. = 4.10 MPa (orifice) _

A Pj = 4.09 MPa (1/2 fill)

0 20 40 60 80 100 120

Average downstream distance (cm)

Figure 42. Average downstream velocities during discharges of HFC- 125 using a HTL valve

under various conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 489

200

180

160

o

V

4.17 MPa (partial saturation)

4.18 MPa (partial saturation)

4.12 MPa (partial saturation) (orifice)

FC-218 (HTL)

C/3 140

>. 120

o

o

>

(L> ^ 80

>

< 60 h

40

20

0

Pj = 2.84 MPa

Pj = 2.83 MPa

?.= 4.05 MPa

A P = 4.10 MPa (orifice)

0 20 40 60 80 100 120

Average downstream distance (cm)

Figure 43. Average downstream velocities during discharges of FC-218 using a HTL valve

under various conditions.
490 8. OPTIMIZATION OF SYSTEM DISCHARGE

200

180

Marotta (orifice)

(/3

160

140

^ 120

o

-§ 100

>

^ 80

(L> 60 -

40 -

O CFjBr (Pj = 4. 16 MPa) • HFC- 125 (Pj = 4.11 MPa)

V FC-218 (Pj = 4.02 MPa)

T CFjHPj =4.12MPa)

20 h

0

0 20 40 60 80 100 120

Average downstream distance (cm)

Figure 44. Average downstream velocities during standard discharges using a Marotta valve

with an orifice plate.
8. OPTIMIZATION OF SYSTEM DISCHARGE 491

Table 22. Calculated void fractions and measured maximum spray angles (discharges with orifice plate)

Agent Mass (g) u (m/s) (MPa) a Maximum

spray angle

CFjBr* 577 66 0.19 0.98 143°

HFC- 125* 452 76 0.41 0.96 142°

FC-218' 442 68 0.21 0.98 145°

CFjl' 695 54 1.32 0.81 147°

CF3Br" 577 59 0.35 0.95

HFC- 125" 425 61 145°

FC-218" 447 58 0.34 0.94 145°

Marotta valve 'HTL valve

curves that by the time the Pf/Pj ratio is reached, all the liquid has been discharged from the vessel;

therefore, no homogeneous nucleation or degassing is possible.

The discharge rates, based on the inflection points, are slower for a lower initial nitrogen charge

pressure. The results obtained by using the HTL valve were very similar. Figures 46 and 47 compare

the discharge characteristics among the agents. Note that CF3I was not used in the HTL valve tests. In Figure 46, the discharge rate of CF3I is the fastest, and CFgBr the slowest.

The comparisons of the average dispersion velocities of the leading edges of the sprays between

the low initial charge pressure tests and the standard discharges are also shown in Figures 37 to 43 for individual agents. In general, the velocities near the valve exit are higher in the standard discharges

than in the low pressure tests. The comparisons of the velocities among the four agents in the low

pressure tests are given in Figure 48. Note again that CF3I has the lowest dispersion velocities. Table 25 tabulates the calculated void fractions and maximum spray angles from the tests using

low initial charge pressure. For CF3I, the calculated void fraction is the lowest.

A set of experiments using HFC- 125 and the Marotta valve was performed in order to examine

the effect of a range of initial charge pressures. The results are summarized in Figure 49. In the

figure, "pure" refers to no nitrogen pressurization. Curve e in the figure was obtained using the static pressure transducer which, in general, exhibited more noise than the dynamic transducer in the

transient pressure measurements. From the figure, degassing appears to occur only for the two highest

initial charge pressures. These two runs have the abrupt pressure recovery associated with the

degassing of dissolved nitrogen. This observation can also be explained qualitatively in terms of the

homogeneous nucleation theory. Table 26 shows the calculated homogeneous nucleation pressures for

various initial charge pressures. At pressures below 2.74 MPa, the calculated homogeneous nucleation

pressures are negative (i.e., in tension). In other words, homogeneous nucleation is predicted not to occur under the experimental conditions used in this study. The prediction agrees with the observations in Figure 49. Although homogeneous nucleation is predicted for P- = 2.14 MPa, the nucleation

pressure is so low by the time this pressure is reached that all the liquid in the vessel has been

depleted; this prediction is in agreement with the observation for = 2.74 MPa in Figure 49 that degassing is not indicated. The dependence of discharge rate on initial charge pressure can also be
492 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 45. Pressure decay curves during discharges of HFC- 125 using a HTL valve under two

fill conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 493

Table 23. Calculated void fractions and measured maximum spray angles (discharges with different

fill densities)

Agent Mass (g) u (m/s) ^ma. (MPa) a Maximum

spray angle

HFC- 125' 428 104 0.41 0.98

HFC- 125' 354" 114 0.33 0.99 114°

CFjl' 719 64 1.65 0.83 141°

CFjl' 547" 57 1.20 0.85 133°

HFC- 125^ 431 98 0.58 0.97 105°

HFC- 125^ 342" 94 1.07 0.92 108°

Marotta valve one-half fill condition ^HTL valve

seen in Figure 49. Based on the inflection points of the curves for similar fill densities, higher initial charge pressures result in higher discharge rates. Figure 50 summarizes the average spray dispersion velocities of HFC- 125 for various initial charge pressures. At the location closest to the valve exit (i.e., 20 cm in the figure), it appears that the

velocity is slower for lower initial charge pressures. The estimated void fractions and the measured

maximum spray angles are given in Table 27. The maximum spray angle initially increases with

increasing initial charge pressure, reaches a maximum, and then decreases with increasing initial charge pressures. Despite the scatter in the measured average velocities and momentum fluxes, the

calculated void fractions appear to be higher at high initial charge pressures than at low charge

pressures. This implies that the high initial charge pressure facilitates the dispersion and evaporation

of the agent downstream.

The effect of initial charge pressure was also studied using HTL squibs. Figures 51, 52, 53,

and 54 show the discharges of CFjBr, HFC- 125, FC-218, and CF3I respectively. The low temperature

condition in the figures will be discussed later. In all the figures, it is clear that there is no indication

that degassing occurs at low initial charge pressures. At the highest initial pressure (P- ~ 5.48 MPa),

the initial fluctuation in the pressure traces in Figures 52 and 53 due to the initiation of the explosive

cartridge could have masked the pressure recovery during degassing; there is no apparent reason why

degassing should not occur in HFC-125/nitrogen and FC-218/nitrogen liquid mixtures at higher initial charge pressures. Once again for CF3I, no degassing is indicated in Figure 54 even at the highest

initial charge pressure. Figure 55 compares the results of all the agents at low initial charge pressures. The overall discharge rate of CF3I is the fastest and that of CF3Br the slowest. Figure 56 compares

the results of HFC- 125 and CF3I at high initial charge pressure. Although the two pressure decay

curves cross over at P/P^ < 0.3, the overall discharge rate of CF3I is still faster than that of HFC- 125

because for P/P- < 0.3, most of the liquid should have been expelled.

Table 28 compares the calculated void fractions and measured maximum spray angles under

various initial charge pressures for all the agents. Note that the scatter in the velocities and momentum flux measurements could have been caused by the flying debris from the explosive cartridge. It appears that the initial charge pressure has little or no effect on the maximum spray angle and that high initial charge pressures result in high average velocities and vice versa.
494 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 24. Calculated homogeneous nucleation pressure (P^j) at T

Agent Mass (g) T(K) P,. (MPa) Ph (MPa)

CFjBr 586 294 2.75 0.51 0.19 HFC- 125 448 294 2.74 0.39 0.14

FC-218 442 295 2.67 0.69 0.26

CF3I 725 294 2.78

8.4.3.9 Effect of Initial Vessel Temperature. Except the spray angles, visual observations from

high-speed movies reveal a similarity in the appearance of the spray between the standard discharges

and the discharges where an agent is initially chilled before discharging into the simulated dry bay

which is at room temperature. However, after most of the agent/nitrogen mixture had been discharged,

large liquid drops were observed dripping out of the vessel outlet. Figures 51, 52, 53, and 54 also show the pressure decay curves of individual agents obtained at cold temperatures. Figure 57 summarizes the results of all the cold agent discharges. Irrespective of

agent, degassing is not noted in the discharges of cold agents. The discharge rates are slower in cold

discharges than in standard discharges because the initial pressures are lower.

In Table 29, the calculated void fractions and measured maximum spray angles and average

velocities are compared to those obtained from standard discharges. Despite the potentially inaccurate

measurements of and u caused by the flying debris from the explosive cartridges, the very low

calculated void fractions based on the release of a cold agent, when compared to the standard

discharges, indicate that the liquid phase still constitutes a significant volume fraction at a location ~ 1 .4 m downstream from the vessel exit. This appears to be consistent with the fact, based on

droplet evaporation theory (Kanury, 1977), that the liquid agent droplets, formed as a result of

flashing, vaporize much slower when the initial droplet temperature is much closer to or lower than its normal boiling point than the droplets whose initial temperature is much higher than its normal boiling

point, as in the case of a standard discharge. The average velocities of cold agent discharges are much

slower than those of standard discharges, and there is also a significant, in some cases a factor of two,

reduction in the maximum spray angle for cold agent discharges. The observations are very similar to the critical discharges of a pure liquid under different degrees of subcooling (Celata et al, 1982, 1983)

in that the spray angle decreases with increasing liquid subcooling.

From the high-speed movies, the spray appearance, other than the spray angle, of a hot-agent

discharge is similar to that of a standard discharge spray. Figures 58, 59, 60, and 61 compare the

internal pressure traces for standard discharges of CFjBr, HFC- 125, FC-218, and CF3I to high

temperature discharges using a Kidde-Graviner valve. Note that the temperatures in these figures are

the measurements from the thermocouple initially located in the liquid phase. In Figures 58 and 59,

the high temperature curves suggest the occurrence of degassing, thus implying that the critical points

of the agent/nitrogen mixtures may not have been reached; otherwise smooth pressure decay curves of

supercritical fluid discharges should have been observed, as in Figure 60 for FC-218. Note that the

calculated mixture critical points from the PROFISSY code are very close to (slightly lower than) the

two high temperatures used in Figures 58 and 59. However, since there is no mixing to ensure thermal equilibration between the liquid and vapor phases during heating, a temperature gradient in the vessel exists. Therefore, it is not possible to know precisely the temperatures of the mixtures inside
OPTIMIZATION OF SYSTEM DISCHARGE 495

Figure 46. Comparison of the pressure decay curves during discharges using a Marotta valve

under low initial charge pressures.
496 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 47. Comparison of pressure decay curves during discharges using a HTL valve under low

initial charge pressures.
OPTIMIZATION OF SYSTEM DISCHARGE 497

200

180

160

Marotta

o

>

>

<

140 -

120 -

100 -

80 -

60 -

40

20 h

0

O CFjBr {? = 2.75 MPa) • HFC-125 (Pj = 2.74 MPa)

V FC-218 (Pj = 2.67 MPa)

CF3I (P. = 2.78 MPa)

0 20 40 60 80 100 120

Average downstream distance (cm)

Figure 48. Average downstream velocities during discharges using a Marotta valve under low

initial charge pressures.
498 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 25. Calculated void fractions and measured maximum spray angles (discharges with P- ~

2.75 MPa)

AgGnt Mass ( u ( m/^)

'"max Uvira; (X K/ln IViaAllllLlllI Y imiim

spray angle

CFjBr* 586 75 1.19 0.90 135°

HFC- 125* 448 73 1.37 0.83 140°

FC-218* 442 85 0.22 0.99 136°

CFjl' 725 47 1.80 0.66 127°

CFjBr" 586 77 1.14 0.91 107°

HFC- 125" 433 88 2.41 0.80 140°

FC-218" 438 78 0.58 0.94 123°

FC-218" 450 77 0.69 0.93 115°

Marotta valve HTL valve

the vessel. On the contrary, in the case of FC-218, the achievable temperature, dictated by the

working pressure of the vessel, in the experiment is much higher (at least 20 °C) than the calculated

mixture critical point. Despite the existence of a temperature gradient, it is reasonable to expect that the condition inside the vessel in this case may still reach supercritical. In Figures 58 and 59, the degassing process can be partly explained using homogeneous

nucleation theory. The differences in the nucleation pressures (at which the plateaus occur) observed

in the standard and high temperature cases can be attributed to differences in the thermodynamic

pathways that lead to the metastable states (at which homogeneous nucleation occurs) of the mixtures

from their initial states. In the standard discharges, the pathways are likely to be isothermal depressurization, whereas in the high temperature discharges, due to the temperature gradients inside the vessel

the pathways are very complicated. For the purpose of illustration, if one assumes the pathways in the

high temperature cases to be isentropic (not an unrealistic assumption), then the homogeneous

nucleation pressures can be lower (possibly in this case) or higher than those observed during

isothermal decompression, depending on the initial entropies of the mixture (Kim-E and Reid, 1983).

The occurrence of a plateau instead of an observable pressure recovery in the figures could be due to the resultant effect of the rate of degassing and the rapid discharge rate. Contrary to the standard discharges, the high temperature result of CF3I in Figure 61 indicates some degree of degassing, evident from the occurrence of a plateau in the pressure trace. Note that the temperature used in the test is still below the calculated mixture critical point. If one recalls that the homogeneous nucleation pressure under standard discharge conditions was calculated to be

negative (i.e., in tension), no degassing was observed in the standard discharges of CF3I. The

occurrence of degassing in the high temperature discharge can be explained based on the different

thermodynamic pathways leading to the metastable states as discussed above in the cases of CFjBr and

HFC- 125. That is, in this case, the homogeneous nucleation pressure of a high temperature discharge

is higher (positive) than that (negative) of a standard discharge. Note also that the disappearance of

the pressure rebound, an artifact, at the end of the high temperature discharge in Figure 61, implying

that temperature has an important role in the performance of the dynamic pressure transducer.
8. OPTIMIZATION OF SYSTEM DISCHARGE 499

HFC- 125 (Marotta)

0 25 50 75 100 125 150

Time (ms)

Figure 49. Comparison of pressure decay curves during discharges of HFC- 125 using a Marotta

valve under various initial charge pressures.
500 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 26. Calculated homogeneous nucleation pressure {Pf^) at T under various initial charge

pressures for HFC- 125

Mass (g) 7"(K) P, (MPa) Ph (MPa)

428 294 4.16 1.76 0.42

456 294 3.39 1.02 0.30

448 294 2.74 0.39 0.14

454 294 2.16

454 294 1.50

453 294 1.17

A comparison of the average spray dispersion velocities for all the agents studied in the standard

and high temperature discharges is given in Figure 29, and CF3I again exhibits the slowest velocities.

For the same agent, the measured dispersion velocities nearest to the valve exit are higher due to higher initial vessel pressure as a result of higher temperature than those measured in the standard

discharges.

The measured maximum spray angles of standard and high temperature discharges using a KiddeGraviner valve are compared in Table 30. The void fractions are not calculated and are not listed in Table 30 because an estimate of a representative downstream temperature cannot be made in the high

temperature discharges. For the same agent, the maximum spray angle is smaller in the standard

discharges than in the high temperature discharges. The very low values of the maximum downstream

dynamic pressures in the case of high temperature discharges of FC-218 also provide some evidence

that FC-218 was discharged as a supercritical fluid (gas) at the temperature of interest.

8.4.3.10 Effect of Vessel Orientation. Based on visual observations from the high-speed movies,

the direction of vessel discharge does not generally seem to affect the spray appearance at the squib

exit although the sprays resulted from vertically upward discharges are narrower downstream. Figures

62, 63, and 64 show snapshots taken at the same instant (5 ms after the activation of the squib) of

vertically downward and upward, and horizontal discharges of HFC- 1 25 using HTL squibs. Figure 65

compares the pressure decay curves during discharges of HFC- 125 using HTL squibs under various

discharge orientations. Two important features are noted in the figure. First, irrespective of the

discharge orientation and as long as the initial conditions {i.e., total pressure, initial temperature,

amount of agent) remain the same in all cases, the degassing occurs approximately at the same P/Palthough the time to reach P/P- differs. This is not surprising because bubble nucleation should be

independent of any direction relative to the gravitation vector. Second, the discharge rates do not seem to depend strongly on the orientation of the vessel, a contrast to the observations made in Grosshandler et al. (1994). In that study, nitrogen was in contact with the agent for a very short time

before the agent was discharged; however, in this work nitrogen was fully saturated with the agent

before the experiments began. The dissolved nitrogen clearly plays an important role in the situation

where the discharge orientation is a limiting factor in the application. However, extreme care should

be exercised when one tries to extrapolate this observation to other fill conditions. For example, in the

situation where the initial liquid level is very low (a case which may not be encountered in practical

applications) and the discharge orientation is upwards, the maximum swelling of the liquid level may
OPTIMIZATION OF SYSTEM DISCHARGE 501

200

180

160

140

120

o

-§ 100

>

as

;-i >

<

80 -

60 -

40 -

20 -

0

o Pi = 1.17 MPa (pure) • Pi = 1.50 MPa

V p.

1 = 2.16 MPa

T P. = 2.74 MPa

p.

1 = 3.39 MPa

Pi = 4.03 MPa

A p. = 4.16 MPa

HFC- 125 (Marotta)

0 20 40 60 80 100 120

Average downstream distance (cm)

Figure 50. Average downstream velocities during discharges of HFC- 125 using a Marotta valve

under various initial charge pressures. -
502 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 27. Calculated void fractions and measured maximum spray angles of HFC- 1 25 discharges

under various initial charge pressures (using a Marotta valve)

Mass (g) P| (MPa) u (m/s) M (MPa) max ^ ' a Maximum

spray angle

438 4.03 114 0.84 0.96 110°

428 4.16 104 0.41 0.98

456 3.39 121 2.41 0.90 124°

448 2.74 73 1.50 0.83 140°

454 2.16 60 2.34 0.57 133°

454 1,50 67 2.30 0.67 115°

453 1.17 65 2.30 0.64 102°

not reach the top of the vessel during depressurization. In this case, a very high quality two-phase

mixture is expected to be discharged from the vessel. Therefore, it is recommended that irrespective

of the fill conditions, the discharge orientation should be downwards so that a very low quality twophase mixture can be obtained.

Table 31 summarizes the measured maximum spray angles and the calculated void fractions.

Because of the flying debris resulted from the cartridge explosion, it is very difficult to make any

reasonably meaningful conclusions in the calculated void fractions. Within the uncertainty, the

measured maximum spray angles appear to be similar in all cases; however, the overall appearance of

the vertically upward discharge spray is different (less dispersed) from that of the standard discharge

spray (c/ Figures 62 and 63).

8.4.3.11 Effect of Partial Nitrogen Saturation. Figures 66 and 67 compare respectively the

discharge characteristics of complete nitrogen saturation (standard discharge) to partial nitrogen

saturation for HFC- 125 and FC-218 at f -4.12 MPa and room temperature using a Marotta valve. In the two figures, the degassing of nitrogen is not noted when it is only partially saturated with the

agent. Another interesting observation is that the discharge rate is slower when the agent is completely saturated with nitrogen.

Figures 68, 69, 70,and 71 compare the two cases under a lower initial total pressure {P- ~

2.75 MPa) for CFgBr, HFC- 125, FC-218, and CF3I respectively. No degassing is noted in either case, and the discharge rate is again slower in the case of complete nitrogen saturation. Note that in Figure 71, the pressure rebound at the end of the discharge is an artifact from the pressure transducer,

as discussed above. The reason for not recording such pressure rebound in one of the runs is not known. When a HTL valve was used for similar tests (complete versus partial), similar trends were

observed with regard to the degassing and the discharge rates. Figures 37, 38, 39, 40, 41, 42, and 43 also display the average downstream spray velocities for

the various agents studied when nitrogen is partially saturated with the agent. Compared to the case of

complete nitrogen saturation, the average spray velocities nearest to the valve exit are consistently

slower.
OPTIMIZATION OF SYSTEM DISCHARGE 503

Figure 5 1 . Comparison of pressure decay curves during discharges of CFjBr using HTL squibs

under various initial conditions.
504 8. OPTIMIZATION OF SYSTEM DISCHARGE

0 25 50 75 100 125 150 175 200 225 250

Time (ms)

Figure 52. Comparison of pressure decay curves during discharges of HFC- 125 using HTL

squibs under various initial conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 505

Figure 53. Comparison of pressure decay curves during discharges of FC-218 using HTL squibs

under various initial conditions.
506 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 54. Comparison of pressure decay curves during discharges of CF3I using HTL squibs

under various initial conditions. . . . . ,
OPTIMIZATION OF SYSTEM DISCHARGE 507

0 50 100 150 200 250 300

Time (ms)

Figure 55. Comparison of pressure decay curves during discharges using HTL squibs at low

initial charge pressure.
508 8. OPTIMIZATION OF SYSTEM DISCHARGE

0 50 100 150 200 250 300

Time (ms)

Figure 56. Comparison of pressure decay curves during discharges using HTL squibs at high

initial charge pressure.
8. OPTIMIZATION OF SYSTEM DISCHARGE 509

Table 28. Calculated void fractions and measured maximum spray angles under various initial charge

pressures (using HTL squibs)

Agent Mass (g) (MPa) u (m/s) Mmax (MPa) a Maximum

spray angle

CF3Br 4.23 93 2.03 0.89 140°

CF3Br 586 2.95 71 2.63 0.74 136°

HFC- 1 25 A CO 458 5.49 93 2.22 0.83 145°

HFC- 125 454 4.12 89 2.37 0.81 131°

HFC- 1 25 453 4.14 95 2.61 0.81 146°

HFC- 125 454 2.63 69 2.24 0.69 140°

FC-218 476 4.21 83 1.72 0.84 144°

FC-218 458 2.79 69 1.68 0.77 147°

CF3I 773 5.48 79 1.49 0.90

CF3I 111 4.17 62 1.52 0.84 136°

CF3I 753 2.81 47 1.39 0.74 130°

8.4.3.12 Effect of Vessel Geometry. In all the tests conducted, no qualitative differences in the

discharge behaviors were observed when the cylindrical vessel equipped with a Marotta, HTL, or Kidde-Graviner valve or a spherical vessel equipped with a HTL squib was used. Although no

experiments have been performed using cylindrical and spherical bottles with identical internal

volumes due to logistics in the experimental hardware design, it is expected that the discharge

behaviors should also be strikingly similar, based on the following arguments. First, nitrogen

degassing should be independent of the vessel geometry assuming that the process can be explained

according to homogeneous nucleation theories. As discussed above, the homogeneous nucleation

theories seem to provide a reasonably good, qualitative prediction of the occurrence of the degassing

process. Second, the discharge characteristics of all the agents studied, using a cylindrical or spherical

vessel with a slightly different internal volume and different release mechanism, are very similar. Third, the discharge process depends mainly on the ullage pressure, and pressure is an intensive

variable and independent of the shape of the vessel; therefore, the effect of vessel geometry will likely

have an insignificant effect on the discharge rate. It should be pointed out that the final drainage of

the remaining liquid through a valve attached to the bottom of one of the two flat heads of a

cylindrical vessel and from a spherical vessel will be different. However, in practical vessel design, a

cylindrical pressure vessel does not normally have two flat heads, but rather has two hemispherical

heads. Consequently, the final drainage problem is expected to be dynamically similar to that of a

spherical vessel.

8.4.3.13 Model Predictions. Based on the above discussion and experimental observations, a model that predicts the discharge times of an agent under various experimental conditions is proposed.

The intent here is to determine whether the proposed model can provide similar qualitative trends

when compared to the experimental observations. The major difference between the proposed model
510 8. OPTIMIZATION OF SYSTEM DISCHARGE

0 50 100 150 200 250 300

Time (ms)

Figure 57. Comparison of pressure decay curves during cold agent discharges using HTL squibs.
8. OPTIMIZATION OF SYSTEM DISCHARGE 511

Table 29. Calculated void fractions and measured maximum spray angles of standard and cold agent

discharges using HTL squibs

Agent Mass T P. u a Maximum

(S) ( ^) (.Mra; (mjs) (WlrH) spray angle

CFjBr 58

1

2 1 4.23 93 2.03 0.89 140°

CFjBr 584 -52 2.03 43 2.25 0.39 82°

HFC- 125 454 21 4.12 89 2.37 0.81 131°

HFC- 125 453 iz 4.14 95 2.61 0.81 146°

HFC- 125 457 -52 2.13 47 2.45 0.27 73°

HFC- 125 455 -55 2.05 85°

FC-218 476 22 4.21 83 1.72 0.84 144°

FC-218 475 -51 2.21 47 2.00 0.41 70°

CF3I 773 21 4.17 62 1.52 0.84 136°

CF3I 771 -58 2.54 48 1.74 0.68 51°

and the model developed by Elliot et al. (1984) is how the process of degassing of dissolved nitrogen

is treated. In Elliot's model (for CFjBr discharge), there is no provision to determine whether

degassing will or will not occur. The degassing of dissolved nitrogen is always assumed to occur, and

the critical bubble size is fixed at 15 nm. Degassing may not occur when CF3I is used or when the

initial vessel pressure is below 4.12 MPa, as discussed above. Note that in Elliot's experiments, the

total initial vessel pressures all exceeded 4.1 MPa. The proposed model, under certain assumptions,

can predict the occurrence or the non-occurrence of degassing. In the following, the assumptions and

the equations used in the model are presented.

The discharge configuration of the vessel is assumed to be vertically downward, i.e., the liquid

mixture will be discharged first, followed by the ullage vapor mixture. During discharge, no heat and mass transfer between the liquid and the vapor phases is assumed to occur. As discussed in Grosshandler et al. (1994), the liquid is assumed to undergo an isothermal depressurization during

discharge, and the degassing of dissolved nitrogen is modeled by homogeneous nucleation theory, as described in Forest and Ward (1977). The occurrence of degassing is determined by comparing the

ullage pressure to the calculated homogeneous bubble nucleation pressure during the liquid discharge

period. If the ullage pressure at some instant in this time interval is equal to or less than the homogeneous nucleation pressure, degassing is assumed to occur. When no degassing is predicted by the

homogeneous nucleation theory, i.e., during the liquid discharge period the ullage pressure never goes

below the homogeneous nucleation pressure, the discharge is modeled as a two-step process, a superheated liquid discharge followed by an emptying of ullage vapor. When degassing occurs, the

model is formulated as a four-step process, a superheated liquid discharge followed by swelling of the

liquid level due to degassing, a two-phase mixture discharge, and finally a vapor discharge.

The control volume under consideration is the ullage above the liquid. Since no heat and mass

transfer is assumed to occur between the liquid and the vapor phases, the control volume is an

adiabatic closed system. Furthermore, due to the rapidity of the discharge, the ullage expansion or
512 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 58. Comparison of pressure decay curves during discharges of CFjBr using a KiddeGraviner valve under two initial agent temperatures.
OPTIMIZATION OF SYSTEM DISCHARGE 513

Figure 59. Comparison of pressure decay curves during discharges of HFC- 125 using a KiddeGraviner valve under two initial agent temperatures.
514 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 60. Comparison of pressure decay curves during discharges of FC-218 using a KiddeGraviner valve under two initial agent temperatures.
8. OPTIMIZATION OF SYSTEM DISCHARGE 515

Figure 61. Comparison of pressure decay curves during discharges of CF3I using a KiddeGraviner valve under two initial agent temperatures.
516 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 30. Comparison of measured maximum spray angles under standard and high temperature

conditions using a Kidde-Graviner valve

Agent Mass T Pi u Maximum

(g) (°C) (MPa) (mis) (MPa) spray angle

CFjBr 625 21 4.14 89 1.28 132°

CFjBr 620 73 9.24 106 1.84 167°

HFC- 125 477 21 4.13 92 2.20 145°

HFC- 125 481 70 9.10 114 1.05 180°

HFC- 125 484 65 9.49 117 1.66 176°

FC-218 492 21 4.14 75 1.18 134°

FC-218 505 75 8.58 111 0.08 170°

FC-218 508 77 8.81 115 0.12 168°

CF3I 800 21 4.15 54 1.21 152°

CF3I 817 90 8.64 68 1.56 161°

momentary compression (caused by degassing) is assumed to be reversible. Therefore, the processes

occurring in the ullage can be assumed to be isentropic. From the first law of thermodynamics for an

isentropic process and assuming an ideal vapor phase,

PV = constant , (26)

where y is the ratio of the constant pressure and volume heat capacities of the vapor mixture, P is the

ullage pressure, and V is the ullage volume. Since nitrogen is the dominant species in the vapor

phase, the y of nitrogen with a value of 1.4 is used in the model calculations. If a time derivative of

Equation (26) is performed, then

^.ll^ = 0 (27)

dt V dt

Before degassing, the initial saturated liquid agent/dissolved nitrogen mixture discharging through

a sharp-edged orifice behaves in a completely metastable manner (Henry, 1979), and the discharge rate can be predicted using incompressible flow formulation. Assuming that the flow through the discharge
8. OPTIMIZATION OF SYSTEM DISCHARGE 517

Figure 62. A snapshot of vertically downward discharge of HFC-125 (taken 5 ms after the

activation of the squib).
518 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 63. A snapshot of vertically upward discharge of HFC- 125 (taken 5 ms after the activation of the squib).
8. OPTIMIZATION OF SYSTEM DISCHARGE 519

Figure 64. A snapshot of horizontal discharge of HFC- 125 (taken 5 ms after the activation of

the squib).
520 8. OPTIMIZATION OF SYSTEM DISCHARGE

0 50 100 150 200 250 300 350

Time (ms)

Figure 65. Comparison of pressure decay curves during discharges of HFC- 125 using HTL

squibs under various discharge orientations.
8. OPTIMIZATION OF SYSTEM DISCHARGE 521

Table 31. Calculated void fractions and measured maximum spray angles of discharges of HFC- 125

using HTL squibs at room temperature, P- -4.12 MPa, and different discharge orientations

Orientation Mass u

(m/s) (MPa) a Maximum

Vertically downward454 89 2.37 0.81 131°

453 95 2.61 0.81 146°

Vertically upward453 83 2.49 0.77 137°

Downward 45° 449 146°

Downward 60° 456

Horizontal 450 99 0.70 0.96 145°

valve can be approximated by an orifice flow of a superheated liquid, the rate of change of the ullage

volume can be expressed in terms of the volumetric flow rate of the liquid as

dt \

2(P - PJ (28)

where is the discharge coefficient for the liquid, with the measured value found to be 0.61 in the

literature (Henry, 1979), A is the opening of the orifice, is the ambient pressure, and p^,^ is the

liquid mixture density. In writing Equation (28), quasi-steadiness is implicitly assumed, that is the

discharge is treated as a steady-state process at any instant in time. In the calculations, the value of

0.61 was used for C^, and the liquid mixture density was obtained from the computer program

PR0F7SSY. Equations (27) and (28) were solved numerically using a fourth-order Runge-Kutta

method (Camahan et al, 1969) to obtain the ullage pressure and volume as a function of time. At

each time step, the calculated pressure was compared to the calculated homogeneous nucleation

pressure in order to determine whether degassing occurred or not. In the case of degassing, the

swelling of the liquid level is treated with the following simple model. At the onset of bubble

nucleation, all the dissolved nitrogen is assumed to form bubbles with a critical radius, and any

subsequent bubble growth and rise to the liquid surface is not considered in the formulation. That is

all the bubbles, once formed, stay in the liquid. This approach was also taken by Elliot et al. (1984).

According to Forest and Ward (1977) who assumed negligible effects of pressure and concentration of

the dissolved nitrogen on the vapor pressure of the liquid and an ideal gas mixture inside the bubble,

the critical bubble radius can be calculated as
522 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 66. Comparison of pressure decay curves during discharges of HFC- 125 using a Marotta

valve at P- -4.12 MPa under complete and partial nitrogen saturation.
8. OPTIMIZATION OF SYSTEM DISCHARGE 523

Figure 67. Comparison of pressure decay curves during discharges of FC-218 using a Marotta

valve at -4.12 MPa under complete and partial nitrogen saturation.
524 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 68. Comparison of pressure decay curves during discharges of CFjBr using a Marotta

valve at P- ~ 2.75 MPa under complete and partial nitrogen saturation.
8. OPTIMIZATION OF SYSTEM DISCHARGE 525

Figure 69. Comparison of pressure decay curves during discharges of HFC- 125 using a Marotta

valve at P- ~ 2.75 MPa under complete and partial nitrogen saturation.
526 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 70. Comparison of pressure decay curves during discharges of FC-218 using a Marotta

valve at P- ~ 2.75 MPa under complete and partial nitrogen saturation.
8. OPTIMIZATION OF SYSTEM DISCHARGE 527

Figure 7 1 . Comparison of pressure decay curves during discharges of CF3I using a Marotta

valve at P- ~ 2.75 MPa under complete and partial nitrogen saturation.
528 8. OPTIMIZATION OF SYSTEM DISCHARGE

CP

K = 2a(P„, - - P,) (29)

sat

where is the critical bubble radius. The volume of each bubble can then be calculated as

y^ubbu = l^^c • (30)

The pressure inside the bubble can be obtained from the Young-Laplace equation (Carey, 1992):

Pbubble ~ ~ • (31)

The amount of nitrogen mass in each bubble can be calculated assuming an ideal gas. The temperature used in the calculations is the initial temperature of the liquid because the thermodynamic path is isothermal depressurization. The total number of nitrogen bubbles formed (Nhutbie^ obtained

from the ratio of the remaining mass of dissolved nitrogen before degassing, which can be calculated

from Equations (27) and (28), to the amount of nitrogen vapor in a bubble. The swelling of the liquid

level, which is assumed to be primarily due to the volume displacement of the nitrogen bubbles, can now be estimated by

^Lswell ^ ^bubble ybubble ^32)

where V^^^^g// is the increase in liquid volume due to swelling. The expansion of the liquid level causes compression of and reduction in the ullage volume. If the process is assumed to be isentropic.

the pressure {Precovery) in the compressed ullage volume (Vcompress) can be calculated using

P = P recovery bd ^bd

V

\ compress /

(33)

where and are the immediate pressure and ullage volume before degassing respectively, which

can be obtained in solving Equations (27) and (28). The above approach used to estimate the pressure

recovery in the pressure decay curves will result in an abrupt rise in pressure rather than a smooth

recovery as noted in the experimental results. The smooth pressure rebound is probably due to the

combined effects of bubble inception and growth, bubble coalescence, bubble rise to the liquid surface,

disengagement of rising bubbles from the liquid surface, and the discharge of the agent during the

degassing period. Because of their complexity, all these effects have not been taken into account in the present model.
8. OPTIMIZATION OF SYSTEM DISCHARGE 529

To simulate the two-phase (vapor/liquid) discharge after degassing, a homogeneous frozen flow

model (Whalley, 1990) is assumed. That is, the two-phase mixture is assumed to be homogeneous,

and the quality, defined as the mass fraction of gaseous nitrogen in the mixture, is assumed to be

constant during the flow through the orifice. The liquid is assumed not to vaporize. Furthermore, no

heat transfer between the gas and the liquid is assumed to occur, and the two-phase mixture is treated

as incompressible, a reasonable assumption for a low quality two-phase flow. Under these assumptions, Equation (28) can be used as an approximation to describe the homogeneous two-phase flow

with being the homogeneous two-phase mixture density in this case. This density can be

estimated using the volume and the mass of the two-phase mixture immediately after degassing. A

discharge coefficient with a value of 0.61 was assumed in the calculations. The initial pressure during

the two-phase discharge period is Precovery- After complete depletion of liquid (in the case of no degassing) or two-phase mixture (in the case

of degassing), the remaining ullage vapor discharge can be modeled using the following equation

(Kim-E and Reid 1983):

1/2 fi

-

[ m j [ 2y I

2y

(34)

with

K = Y - U

y * 1 (35)

where is the discharge coefficient, m is the molecular weight of the vapor (assuming to be

nitrogen), P^i and T^j are the pressure and temperature in the vessel at the instant of complete

depletion of liquid or two-phase mixture, respectively, and is the vessel volume. A value of 0.61 was also used for in the calculations T^l can be calculated by

Tel = T

ullage (36)

where V^^^^^ and T are the initial vapor ullage volume and temperature, respectively.

If the initial vessel temperature is above the critical temperature of the agent/nitrogen mixture, the

discharge of the supercritical fluid may simply be modeled using Equations (34) and (35) with T^f and

P^i being the initial vessel temperature and pressure. Table 32 summarizes the predicted and observed complete discharge times for all the agents

under the standard discharge conditions. Although the time corresponding to the inflection point in the

pressure trace could be used to indicate the end of liquid discharge, this approach was not used due to
530

Table 32.

8. OPTIMIZATION OF SYSTEM DISCHARGE

Comparison between predicted and observed discharge times for standard discharges

Agent Mass Observed discharge time Predicted discharge time

(g) (ms) ± 5 ms (ms)

CFjBr* 591 89 55

CFjBr' 564 54

CFjBr^ 625 72 21

CFjBr'' 581 265 79

HFC-125* 438 77 51

HFC- 125* 428 50

HFC-125^ 477 64 20

HFC-125'' 454 257 75

FC-218* 450 73 50

FC-218* 435 49

FC-218'' 492 60 19

FC-2181' 476 250 74

CFjI* 723 75 56

CF3I* 719 75 56

CF3I* 755 59

CF3I* 800 78 22

CF3I' 773 295 86

Marotta valve ^Kidde-Graviner valve •htl squibs

the uncertainties in discerning the inflection points in most of the pressure traces. Instead, the

complete discharge times were used in the comparison.

In Table 32, the observed complete discharge times were measured from the high-speed movies.

These times aie defined as the interval between the instant when the flashing spray first appears at the

vessel exit and the time when the spray can no longer be observed. Using this definition, the

discharge times for CF3I are the longest, which seems to contradict the observations from the

measured pressure decay curves. However, the high-speed movies show that the denseness of the

flashing CF3I spray diminishes much sooner than those of the other agents, implying that the bulk of

CF3I discharges faster. The calculated discharge times do not show such a distinct trend among

agents. This is probably due to the assumptions used in the model. When all the dissolved nitrogen

is assumed to degas, it will cause an artificially high liquid swelling level, resulting in an artificially

high pressure recovery in the ullage and an overall faster discharge rate. Therefore, the agent/nitrogen

mixtures with a tendency to degas will have shorter discharge times. In addition, the calculated

homogeneous nucleation pressures in the mode! agree only qualitatively with the observed degassing

pressures. Although one can use the experimentally determined degassing pressures as input
8. OPTIMIZATION OF SYSTEM DISCHARGE 531

parameters in the model and good predictions can be obtained if the measured degassing pressures and

the pressure recoveries are used as input in the model, it is better to predict these pressures using

homogeneous nucleation theories so that no a priori information is required and the versatility of the

model can be demonstrated. Despite these simplified descriptions of the liquid swelling process and

nitrogen degassing, the discharge times can still be estimated within a factor of four (usually better when a solenoid valve was used) by the model. It is conjectured that the action of a squib or pyrotechnic device could render some of the assumptions used in the model invalid.

8.4.4 Summary and Conclusions. Experiments have been conducted in order to delineate the effects of various parameters on agent discharge into an enclosure. The major experimental observations are summarized in the following.

1 . Irrespective of the release mechanisms (two solenoid valves, one piston-actuated valve, and a

squib), CF3I exhibits the fastest volumetric discharge rate from the pressure vessel at room

temperature; however, the average dispersion velocity of the spray is the slowest once the

agent leaves the vessel. All the other agents have comparable volumetric discharge rates and

spray dispersion velocities.

2. For a given initial vessel pressure, reducing the orifice opening prolongs the discharge time.

Based on the experimental results, the discharge time is approximately inversely proportional

to the area of the orifice opening. Reducing the liquid fill volume decreases the discharge

time because less agent mass is available in the vessel. For a given fill condition, increasing

initial nitrogen charge pressure decreases the discharge time and vice versa. However, due to the dependence of the degassing process on the initial charge pressure, no simple, empirical

correlation between the discharge time and the initial vessel pressure was apparent from the

experimental data.

3. The average dispersion velocities of cold agent discharges are much slower than those of

discharges at room temperature. In addition, the discharge rates of cold agents are slower, and

the spray angles are narrower.

4. Based on the results of nitrogen saturated HFC- 125 discharges {2/3 liquid fill), the vessel

orientation does not affect the discharge rate significantly. The dissolved nitrogen in the

liquid agent plays an important role. .-

: .;•

5. When the agent is partially saturated with nitrogen, instead of complete nitrogen samration, the

volumetric discharge rate is much faster.

6. A discharge model incorporating homogeneous nucleation theories to account for nitrogen

degassing predicts all qualitative aspects of the discharge process and the discharge times

within a factor of four.

8.5 Computer Simulation of Agent Dispersion

8.5.1 Introduction. The objective of this study is to explore a computational tool which will provide

detailed analysis on the dispersion process of a flashing spray in a simulated dry bay. Since the
532 8. OPTIMIZATION OF SYSTEM DISCHARGE

discharge process under this scenario is extremely rapid (events typically lasting less than 100 ms),

detailed spatial and temporal measurements of temperatures, velocities and concentrations of the

flashing spray are very difficult. Therefore, numerical simulation may prove valuable for studying

such a highly transient process. This work was conducted in connection with the experimental studies on the agent discharge and dispersion processes described in Section 8.4.

8.5.2 Modeling of Agent Dispersion Process. The computer codes KIVA-II and CONCHASSPRAY developed at the Los Alamos National Laboratory (LANL) (Amsden et al, 1989;

Cloutman et al., 1982) were used to simulate the agent dispersion of a flashing spray which was

obtained from a liquid agent being discharged from a nitrogen-pressurized vessel through a simple,

nozzle/orifice-like exit. The system under consideration involves a vessel volume of the order of

10 m , with a discharge time of the order of 10 s. To achieve such a rapid discharge, the pre- pressurization level is typically on the order of several megapascals.

The dispersion of a flashing spray into a confined space is a complicated process which involves

many physical phenomena occurring simultaneously and interacting with one another. As the agent

exits from the vessel, thermodynamic and fluid-dynamic instabilities lead to flashing and break-up of

the agent into a two-phase gaseous/droplet jet mixture. This occurs in a relatively short transition

region which starts at the vessel exit and ends at a section of the two-phase jet where thermodynamic

and fluid-dynamic instabilities have ceased.

Downstream of this section, called the "initial section," the flow begins to develop as a mixed,

two-phase agent/air jet. Beyond the initial section, thermodynamic equilibrium is maintained and

droplet collision and agglomeration do not play an important role in either the ensuing jet dynamics or

the dispersion of the agent throughout the protected space. Within the initial section, the spray

consists of a mixture of liquid-agent fragments of various shapes and sizes and gaseous agent. The

spray begins to entrain initially quiescent air from the surrounding environment as it moves away from

the bottle exit. Because the temperature of the discharged liquid agent is far above its normal boiling

point, it very quickly evaporates and creates a cloud of agent vapor surrounding the moving spray. The entrainment, mixing and evaporation continue as the spray is dispersed throughout the space. The

above processes are included in the present analysis.

The dispersion process is described as a two-phase, three-component, turbulent, compressible,

dissipative flow. It is assumed that the gas phase consists of the following three components: agent

vapor, nitrogen, and oxygen. The liquid phase only consists of pure liquid agent. The vapor pressure

equation is a highly non-linear function of temperature. Heat conduction and convection, diffusive and

convective mass transfer, momentum transport, and turbulent flow are taken into account in the gas- phase formulations. In addition, the interphase transport processes, such as momentum exchange in the form of Stokes forces and aerodynamic drag between the droplets and the gas phase, heat

exchange in the form of forced convection, and mass exchange in the form of evaporation, are considered. It is assumed that the spray droplets can be represented by a Sauter mean diameter. The

droplet evaporation process is assumed to be quasi-steady. The effect of turbulence on droplet motion

within the spray is also included, and the turbulence in the gas phase is described by relating the

velocity gradients to the turbulent viscosity.

The transport coefficients (heat conductivity, mass diffusivity, and viscosity) and the specific

heats are temperature dependent. The effect of the gravitational field is also included. No chemical

reactions and their associated thermal effects {e.g., heat release) are assumed in the formulation.

The governing equations are the conservation equations of mass, momentum and energy for the

medium under investigation. This set of equations constitutes an initial/boundary value problem within

a region bounded by the initial section and the far-field boundaries of the protected space. The set of
8. OPTIMIZATION OF SYSTEM DISCHARGE 533

time-dependent equations is expressed in a three-dimensional form although the discharges were assumed to be axisymmetric.

The total mass density of the fluid, excluding the mass of the spray droplets, is defined in terms

of a partial mass density of the component species. The mass balance equates the change in species

concentration with time to the sum of the advection, diffusion and source terms due to evaporation of

that particular component. The total fluid density satisfies the total mass balance equation, recognizing

that the net diffusion must sum to zero. The axial and radial momentum balance equations for the fluid mixture equate the time rate of

change of momentum to the sum of the advection, pressure, viscous stress, droplet drag, and body

force terms. The energy equation for the gas phase accounts for the local change in specific internal

energy from the sum of the net advection of sensible heat, heat conduction, enthalpy transport

accompanying the diffusion of individual species, flow work, viscous heating, and energy transfer

between the two phases. The state relations are assumed to be those of an ideal gas mixture. The

transport coefficients include the laminar and turbulent contributions to the momentum, heat, and mass

transfer. The turbulent viscosity is broken into a constant uniform background term and a variable

turbulent viscosity. The thermal and mass turbulent diffusivity are related to the viscosity through the

local Prandtl and Schmidt numbers of the gas mixture.

The motion of the droplets in the spray is governed by ordinary differential equations in Lagrangian form. It is assumed that the spray is sufficiently thin, and, as a consequence, the volume

displaced by the droplets may be neglected. Velocity components in the axial, radial and swirl

directions are tracked, and the transport between phases is dependent upon the Reynolds number

computed from the slip velocity and the film temperature and composition. Internal droplet motion is ignored, leading to uniform droplet temperatures.

The model developed by Dukowicz (1979) was used to calculate droplet mass loss due to evaporation. The particle velocity is computed from the drag associated with the slip velocity. The

latent heat cools the droplet during evaporation, and the heat transfer between the drop and the

environment is determined using the Ranz-Marshall correlation (1952), modified to account for mass

leaving the surface. The fluid-particle interactions are all two-way coupled to ensure overall

conservation of mass, momentum and energy.

8.5.3 Solution Method. The gas-phase solution procedure is based on a finite volume method called

the ALE (Arbitrary Lagrangian-Eulerian) method. Spatial differences are formed on a finite-difference

mesh that subdivides the computational region into a number of small cells that are hexahedrons. The

comers of the cells are called vertices, and the positions of the vertices may be arbitrarily specified

functions of time, thereby allowing a Lagrangian, Eulerian, or mixed description. The arbitrary mesh

can conform to curved boundaries and can move to follow allowable changes in geometry. An

advantage of the method is that the mesh need not be orthogonal. The transient solution is marched

out in a sequence of finite time increments called cycles or time steps. In each cycle the values of the

dependent variables are calculated from those in the previous cycle.

8.5.4 Numerical Results and Discussion. Computations using the two-dimensional axisymmetric

approach were performed using the LANL CONCHAS-SPRAY and KIVA-II codes. CONVEX and

CRAY supercomputers with the UNIX operating system at NIST were used for the calculations.

The results for a generic agent simulant with thermophysical properties very similar to HCFC-22

are presented graphically as contours of constant velocity, density, pressure, temperature, species

concentration, vorticity, kinematic viscosity, mass flux, and liquid phase. The only difference in thermophysical properties between the simulant and HCFC-22 is the vapor pressure curves. As will be

discussed below, the actual vapor pressure curve of HCFC-22 could not be used in the calculations
534 8. OPTIMIZATION OF SYSTEM DISCHARGE

because of convergence problems. The shape of the curve was modified slightly so that calculations

could be performed. The rationale for selecting this generic simulant is that its thermophysical

properties lie within those of the selected agents.

The performance of the two codes, CONCHAS-SPRAY and KIVA-II, were compared. It was

found that in one test case, CONCHAS-SPRAY and KIVA-II gave essentially the same results. Since

CONCHAS-SPRAY requires less computing time, the computational results reported herein were

obtained using CONCHAS-SPRAY It should be emphasized that the original codes were designed

for the use with the LANL computer facility. In addition, the codes employed graphics libraries installed in that system. To apply the codes using the NIST computing facility, it was necessary to configure them to the NIST operating system. The codes have also been modified to write their

output into organized ASCII files. The output files were transferred onto a PC computer where they were further graphically processed with a graphics program written in the C language. The program

converted the numerical output into postscript files. The configuration investigated is the discharge of a droplet spray, with properties similar to one

of the halon alternatives, into a totally enclosed space. The transitional region between the nozzle exit

and a fully-developed spray was not considered but has been discussed by Cooper (1994). Therefore,

it is assumed that the spray has been already established at the discharge orifice exit and is injected

into the area at a prescribed rate. Computational parcels representing the droplets are introduced at the injector with an axial

velocity of 61.4 m/s, directed at an angle of 0° relative to the symmetry axis. A zero azimuthal

component of velocity is assumed for the droplets at the injector. All the gas and liquid parameters

are assumed to be uniform across the orifice exit. All the parameters within the compartment are also

assumed to be uniform initially.

The surrounding walls of the compartment are treated as solid boundaries maintained at a constant temperature of 21 °C. Free-slip conditions are assumed on the walls for the velocities. Thus the

subsequent penetration and shape of the spray result solely from the interactions of the liquid droplets

with the ambient air. The numerical solution of the balance equations associated with the initial and

boundary conditions and the material properties provides the time evolution of the spray in the

compartment.

The nozzle is located on the centerline of the 0.42 m diameter cylindrically shaped volume. The

length of the compartment is either 1.0 m or 2.5 m. Agent inflow boundary conditions, including

temperature, density, and velocity are estimated from the results of the discharge vessel experiments

and specified in the calculation. The ambient temperature is either 21 °C or - 30 °C. The ambient

pressure is 1 x 10^ kPa. The liquid spray mean diameter is either 200 |im or 600 pm. Table 33 lists the set of the input data used as initial parameters in the computations and the figure numbers corre- sponding to the figures showing the computational results. Figures 72 to 76% illustrate the computational mesh used in the calculations and the temporal

variations of liquid spray outline, agent vapor density, temperature, pressure and gas density contour

under the following conditions: initial droplet diameter of 200 pm, enclosure length of 1 m, and initial ambient temperature equal to the 21 °C.

The effect of droplet diameter is demonstrated in Figures 77 to 80, where the initial droplet size

varies from 200 pm to 600 pm; all other parameters remained constant. Figures 81 to 84 illustrate the evolution of the spray with an initial Sauter mean diameter of

200 pm and an initial ambient temperature of - 30 °C. Figures 85 to 88 show the results for the

larger initial mean diameter (600 pm) at - 30 °C.

Figures 89 to 92 show the calculated results using a Sauter mean diameter of 200 pm and a

different compartment length (2.5 m) at 21 °C. Figures 93 to 96 display the results for a Sauter mean

diameter of 600 pm under the same initial ambient temperature.
8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 33. Input data used in the computations

535

Figure number Spray mean diameter (|im) Ambient temperature (°C) Compartment height (m)

72-76 200 21 1.0

77-80 600" 21 1.0

81-84 200, -30 . 1.0

85-88 . 600 -30 1.0

89-92 200 21 2.5

93-96 600 21 2.5

Based on the computational results, the penetration distance of the spray (around 0.9 m after 14 ms, which roughly corresponds to the values measured in the totally open discharge apparatus) is about the same for the 200 |jm and 600 pm drops; however, the spray structure and thickness differ

significantly. For the small droplet diameter, the spray is clearly thicker, especially in the head of the

spray, the distribution of the parameters within the spray is more uniform, and the higher density area of the gas phase (containing agent vapor) is much larger.

An additional aerodynamic effect is the tendency to form a vortex in the vicinity of the head of

the spray with the small droplet mean diameter. This phenomenon enhances the area filled with the

agent vapor. The spray with the large droplet mean diameter is significantly narrower and without the

fat head.

An unexpected effect is the generation of local temperatures which exceed the initial ambient

temperature. This is caused by the strong shock waves created during the discharge process which

compress the gas. The possibility of creating such waves is indicated by Thompson et al. (1986) and

Kurschat et al. (1992). Although the vaporization of the agent cools the chamber, it is possible that some local temperature rise due to compression of the strong shock waves can offset the temperature

decrease due to evaporation. The maximum gas phase temperature is significantly higher for the spray

with the small droplet mean diameter (there is an increase of 40 K relative to the initial ambient

temperature), whereas for the large droplet mean diameter, the temperature only increases 10 K. This

influences the interphase heat transfer, which depends on the surface area and temperature difference.

Similar observations, with one exception, were made at a lower initial ambient temperature (- 30 °C).

For the small droplet diameter, the maximum gas-phase temperature is just slightly higher than the

ambient value, but for the large droplet diameter, it is lower than the ambient value and nearly equal

to the liquid phase temperature (boiling point).

Similar observations were also made when a longer compartment was used in the simulation.

Again, there is one exception in the gas-phase temperature behavior. For the large droplet mean

diameter, the maximum gas-phase temperature is slightly higher than the ambient value. For the small

droplet diameter, the temperature is lower by 5 1 ° relative to the ambient value. The conclusion is that the dynamics of the interphase heat exchange phenomena depends strongly on the spray Sauter mean

diameter, ambient temperature, and geometry. . r The effect of compartment geometry on the dispersion process can be assessed by comparing the

'

results obtained using two different compartment heights. In the case of a 1.0 m high compartment, a

pressure wave is generated by the fast moving agent just after 1 ms of the process. The pressure wave

reflects alternately from the bottom and top walls of the compartment, thereby influencing the spray
536 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 72. Agent discharge. Computational mesh and liquid phase outlines at: 8.3 ms, 10.7 ms,

and 14.3 ms (upward). Initial spray angle 0°. Spray mean diameter 200 |im. Ambient

T = 294 K and P = 101 kPa. Compartment 1 m x 0.42 m.
8. OPTIMIZATION OF SYSTEM DISCHARGE 537

Figure 73. Vapor density maps at: (3.5, 8.3, 10.7, 14.3) ms (upward); the outline value

0.032 kg/n?. Initial spray angle 0°. Spray mean diameter 200 |im. Ambient T =

294 K and P = 101 kPa. Compartment 1 m x 0.42 m.
538 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 74. Gas phase temperature maps at: (3.5, 8.3, 10.7, 14.3) ms (upward); the maximum

value 335 K. Initial spray angle 0°. Spray mean diameter 200 |im. Ambient T =

294 K and P = 101 kPa. Compartment 1 m x 0.42 m.
8. OPTIMIZATION OF SYSTEM DISCHARGE 539

Figure 75. Gas phase pressure maps at: (3.5, 8.3, 10.7, 14.3) ms (upward); the outline value

105 kPa. Initial spray angle 0°. Spray mean diameter 200 |im. Ambient T = 294 K

and P = 101 kPa. Compartment 1 m x 0.42 m.
540 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 76. Gas phase density maps at: (3.5, 8.3, 10.7, 14.3) ms (upward); the outline value

1.27 kg/nv'. Initial spray angle 0°. Spray mean diameter 200 |im. Ambient T = 294 K

and P = 101 kPa. Compartment 1 m x 0.42 m.
8. OPTIMIZATION OF SYSTEM DISCHARGE 541

Figure 77. Vapor density maps at: (3.5, 8.3, 10.7, 14.3) ms (upward); the outline value

0.026 kg/m?. Initial spray angle 0°. Spray mean diameter 600 pm. Ambient T =

294 K and P = 101 kPa. Compartment 1 m x 0.42 m.
Figure 78. Gas phase temperature maps at: (3.5, 8.3, 10.7, 14.3) ms (upward); the maximum

value 304 K. Initial spray angle 0°. Spray mean diameter 600 \xm. Ambient T =

294 K and P = 101 kPa. Compartment 1 m x 0.42 m.
8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 79. Gas phase pressure maps at: (3.5, 8.3, 10.7, 14.3) ms (upward); the outline value

113 kPa. Initial spray angle 0°. Spray mean diameter 600 |jm. Ambient T = 294 K

and P = 101 kPa. Compartment 1 m x 0.42 m.
544 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 80. Gas phase density maps at: (3.5, 8.3, 10.7, 14.3) ms (upward); the outline value

1.37 kg/m^. Initial spray angle 0°. Spray mean diameter 600 |im. Ambient T = 294 K

and P = 101 kPa. Compartment 1 m x 0.42 m.
8. OPTIMIZATION OF SYSTEM DISCHARGE 545

Figure 81. Vapor density maps at: (3.5, 8.3, 10.7, 14.3) ms (upward); the outline value

0.036 kg/m^. Initial spray angle 0°. Spray mean diameter 200 \im. Ambient T =

243 K and P = 101 kPa. Compartment 1 m x 0.42 m.
546 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 82. Gas phase temperature maps at: (3.5, 8.3, 10.7, 14.3) ms (upward); the maximum

value 249 K. Initial spray angle 0°. Spray mean diameter 200 [im. Ambient T =

243 K and P = 101 kPa. Compartment 1 m x 0.42 m.
. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 83. Gas phase pressure maps at: (3.5, 8.3, 10.7, 14.3) ms (upward); the outline value

89 kPa. Initial spray angle 0°. Spray mean diameter 200 |im. Ambient T = 243 K

and P = 101 kPa. Compartment 1 m x 0.42 m.
548 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 84. Gas phase density maps at: (3.5, 8.3, 10.7, 14.3) ms (upward); the outline value

1.29 kg/m^. Initial spray angle 0°. Spray mean diameter 200 \xm. Ambient T = 243 K

and P = 101 kPa. Compartment 1 m x 0.42 m.
8. OPTIMIZATION OF SYSTEM DISCHARGE 549

J

Figure 85. Vapor density maps at: (3.5, 8.3, 10.7, 14.3) ms (upward); the outline value

0.029 kg/m^. Initial spray angle 0°. Spray mean diameter 600 |jm. Ambient T =

243 K and P = 101 kPa. Compartment 1 m x 0.42 m.
550 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 86. Gas phase temperature maps at: (3.5, 8.3, 10.7, 14.3) ms (upward); the maximum

value 233 K. Initial spray angle 0°. Spray mean diameter 600 ^m. Ambient T =

243 K and P = 101 kPa. Compartment 1 m x 0.42 m.
8. OPTIMIZATION OF SYSTEM DISCHARGE

J

1

Figure 87. Gas phase pressure maps at: (3.5, 8.3, 10.7, 14.3) ms (upward); the outline value

0.93 kPa. Initial spray angle 0°. Spray mean diameter 600 \xm. Ambient T = 243 K

and P = 101 kPa. Compartment 1 m x 0.42 m.
552 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 88. Gas phase density maps at: (3.5, 8.3, 10.7, 14.3) ms (upward); the outline value

1.35 kg/m^. Initial spray angle 0°. Spray mean diameter 600 |am. Ambient T = 243 K

and P = 101 kPa. Compartment 1 m x 0.42 m.
8. OPTIMIZATION OF SYSTEM DISCHARGE 553

J

Figure 89. Vapor density maps at: (3.5, 8.3, 9.4, 9.5) ms (upward); the outline value

0.024 kg/m^ Initial spray angle 0°. Spray mean diameter 200 |im. Ambient T =

294 K and P = 101 kPa. Compartment 2.25 m x 0.42 m.
554 8. OPTIMIZATION OF SYSTEM DISCHARGE I

Figure 90. Gas phase temperature maps at: (3.5, 8.3, 9.4, 9.5) ms (upward); the maximum value

243 K. Initial spray angle 0°. Spray mean diameter 200 |jm. Ambient T = 294 K and

P = 101 kPa. Compartment 2.25 m x 0.42 m.
OPTIMIZATION OF SYSTEM DISCHARGE

Figure 91. Gas phase pressure maps at: (3.5, 8.3, 9.4, 9.5) ms (upward); the maximum value

121 kPa. Initial spray angle 0°. Spray mean diameter 200 jjm. Ambient T = 294 K

and P = 101 kPa. Companment 2.5 m x 0.42 m.
556 8. OPTIMIZATION OF SYSTEM DISCHARGE

J

Figure 92. Gas phase density maps at: (3.5, 8.3, 9.4, 9.5) ms (upward); the outline value

1.39 kg/n?. Initial spray angle 0°. Spray mean diameter 200 |am. Ambient T = 294 K

and P = 101 kPa. Compartment 2.25 m x 0.42 m.
8. OPTIMIZATION OF SYSTEM DISCHARGE 557

Figure 93. Vapor density maps at: (3.5, 8.3, 10.7, 13.8) ms (upward); the outline value

0.024 kg/m^. Initial spray angle 0°. Spray mean diameter 600 |im. Ambient T =

294 K and P = 101 kPa. Compartment 2.5 m x 0.42 m.
558 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 94. Gas phase temperature maps at: (3.5, 8.3, 10.7, 13.8) ms (upward); the maximum

value 298 K. Initial spray angle 0°. Spray mean diameter 600 jam. Ambient T =

294 K and P = 101 kPa. Compartment 2.5 m x 0.42 m.
8. OPTIMIZATION OF SYSTEM DISCHARGE 559

Figure 95. Gas phase pressure maps at: (3.5, 8.3, 10.7, 13.8) ms (upward); the outline value

102 kPa. Initial spray angle 0°. Spray mean diameter 600 ^im. Ambient T = 294 K

and P = 101 kPa. Compartment 2.5 m x 0.42 m.
560 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 96. Gas phase density maps at: (3.5, 8.3, 10.7, 13.8) ms (upward); the outhne value

1.19 kg/m^. Initial spray angle 0°. Spray mean diameter 600 (im. Ambient T = 294 K

and P = 101 kPa. Compartment 2.5 m x 0.42 m.
8. OPTIMIZATION OF SYSTEM DISCHARGE 561

structure and dynamics. The spray becomes thicker and slower. In the case of a 2.5 m long

compartment, the spray penetration is greater, but at the same time the spray and its head are thinner.

After 1 ms, the wave generated by the spray still travels downward, approaching the bottom wall. It eventually reaches the wall at a later moment and becomes weaker due to dissipation and collision

with the head of the spray. The speed of the pressure wave for the two different compartment heights

is about the same. Because of the difference in the compartment height, the pressure wave influences

the spray structure 2.5 times more frequently in the short compartment. Moreover, the interactions are

shifted in time, thereby affecting the spray dynamics and structure. A similar effect was observed

when a different initial droplet mean diameter was used.

By comparing the results for the two different ambient temperatures (with the spray mean

diameter fixed at 200 jam), the penetration of the spray is about the same; however, the spray

structures differ significantly. This is evident especially in the temperature, pressure, and density

fields. At an ambient temperature of 21 °C, the density outlines of the gaseous phase are significantly

broader The maximum temperature of the gas phase is higher than the initial ambient temperature by

about 40 °C, thereby enhancing the interphase heat exchange process. At - 30 °C, the maximum

temperature is about the same as the ambient one. This phenomenon may be associated with the much

faster evaporation process, which depends on the interphase heat exchange rate and the compressibility

effect in the gas phase. To further substantiate these observations, the behavior of the reflected pressure waves from the walls is examined. In the case of the higher ambient temperature, the pressure wave travels faster. After 10.7 ms in the 1.0 m long compartment, the pressure wave

interacts with the spray, while in the other case the pressure wave reaches the spray only after 14.3 ms. Thus, the evaporation process affects the phenomena associated with the compressibility

effects. The effect of initial ambient temperature depends, to some extent, on the spray mean diameter.

For the 600 pm diameter, all the phenomena described above occur, with the exception of the

maximum gas-phase temperature. In this case, the peak temperature is only slightly higher than the

ambient value. This indicates that the process of cooling via evaporation dominates the compression

heating. In all cases, the maximum values of the velocities in the gas phase are between 80 m/s and

290 m/s, the lower values being comparable to those measured in the discharge experiments (refer to Section 8.4). In the computations, the initial velocity of the agent droplets was assumed to be

61.4 m/s, and the ambient air was initially stagnant. In the near-field, momentum is transferred to the

air, increasing the air velocity at the expense of the droplets. In the far-field, the velocities of both

phases would be close to each other if vaporization were not present. However, the velocity of the gas

phase increases dramatically. Under certain conditions the agent evaporation process becomes so fast

that the vapor leaving the droplet surface contributes to the resultant gas velocity. The phenomenon

occurs in the vicinity of the injector orifice and also downstream, well beyond the injector area. The

maximum values occur first in the vicinity of the injector orifice during the first 3 ms of the process, and then in a later stage, between 10 ms (smaller diameter) and 14 ms (bigger diameter) downstream.

Such high velocities lead to the creation of pressure waves, reflecting further between the walls. When the local temperature of the gas phase approaches the agent boiling point, the Mach number

may exceed the sonic condition because the speed of sound decreases with the square root of absolute

temperature. This means that a weak shock wave is created. The wave initially appears close to the

injector exit and then travels downstream toward the end wall. It is also stronger in the area downstream, where higher gas velocities occur. A typical maximum increase of pressure at the wave

is about 20 %.

It is necessary to remember that the dynamics of the dispersion process depends on the initial conditions in the transitional region where we have no experimental data to use as input for the
562 8. OPTIMIZATION OF SYSTEM DISCHARGE

calculations. In the above simulations, the droplets were assumed to have a specific size and the

temperature was set equal to the agent boiling point, - 41 °C. In reality, the droplet size distribution

is unknown, and the theories leading to droplet formation under such dynamic conditions have not yet

been established. Likewise, the temperature may be higher in the transition zone. Because the evaporation model approximates the agent vapor pressure as an exponential function of the surrounding temperature, small changes in temperature have a large impact on the spray dynamics.

It should be emphasized that realistic results from such models as CONCHAS-SPRAY/KIVA-II

will be dependent on the validity of the transition-region concept and on the ability to provide reasonable estimates of the state of the spray at the initial section (i.e., the boundary conditions). In spite of

the fact that the initial droplet size distribution may play an important role in the agent dispersion

process, a method for estimating such a distribution does not seem to be forthcoming. In the present

calculations, the droplet size distribution could only be hypothesized.

A numerical problem was encountered when solving the set of equations describing the droplet

evaporation process. The problem arises when the agent vapor pressure equation is described by an

exponential function of temperature and when the temperature difference between the liquid and gas

phases is over 10 °C; this problem becomes worse with increasing temperature difference between the two phases. Under such conditions, the required time step to obtain stable computation is on the order

of 10 s when the original vapor pressure equation is used. If the vapor pressure equation is analytically modified in such a way to mitigate the sharp pressure increase, the time step is increased by six orders of magnitude to 10~^ s. However, this modification is associated with the change of the

original thermophysical properties of the compound. Therefore, it is impossible at the present time to compare the computational results with the dispersion characteristics determined experimentally. Note

that the LANL codes are originally designed for applications to internal combustion engines, where

liquid hydrocarbons are used. It is possible that a separate subroutine could be written to solve the

vapor pressure equation with a numerical method appropriate for this highly non-linear behavior. One

possible solution to this problem is to adopt the numerical techniques that are used in solving

exponential chemical kinetics equations.

A further comparison between KIVA-II and CONCHAS-SPRAY for the same set of conditions is needed to determine the impact of the turbulence and droplet breakup model on the results. If the

outcome is independent of the code, additional analyses with CONCHAS-SPRAY could be conducted.

If the turbulence and droplet models influence the results, KIVA-II or preferably KIVA-III

(Amsden et al., 1993) should be used for future calculations.

The KIVA codes have the additional advantage that fully three-dimensional spaces can be

handled. Extended resources on a CRAY computer will be required for these investigations. The

spectrum of parameters influencing the spray dynamics should be broadened to include additional

compartment geometries, discharge orifice geometry, thermodynamic and gas dynamic parameters of

the agent and ambient air, agent injection and atomization parameters, as well as the conditions of the

transition region of spray.

8.5.5 Conclusions. Numerical simulation was carfied out using the LANL CONCHAS-SPRAY and

KIVA-II computer codes in an attempt to better understand agent dispersion processes. The following

conclusions can be drawn based on the computational results obtained so far:

1 . CONCHAS-SPRAY is at least an order of magnitude more efficient, in terms of the computational time, than KIVA-II when applied to the dispersion process of a simulated halon

alternative. Both codes give the same qualitative and quantitative results using the same input

data.
8. OPTIMIZATION OF SYSTEM DISCHARGE 563

2. It was found from the calculations that the geometry of a compartment into which the agent was released is an important parameter governing the spray dynamics. The propagation of

pressure waves generated by the fast moving agent affects the shape and penetration of the

spray and eventually affects the agent vapor mass fraction field due to pressure wave reflections. The smaller the compartment, the thicker the spray, the lesser the spray penetrates, and

the broader the agent density field.

3. The liquid agent atomization process, expressed in terms of Sauter mean diameter, influences

the spray structure significantly. The smaller the spray mean diameter, the better the agent

disperses.

4. For a higher ambient temperature, the spray is broader, and the penetration of the spray is farther.

5. The occurrence of higher gas velocities corresponds to the presence of pressure waves. The

phenomenon may be associated with the enhanced evaporation process which contributes to the total gas velocity.

8.6 Pipe Flow Characteristics of Alternative Agents

8.6.1 Introduction. Three potential near-term halon alternatives for engine nacelle fire protection

applications have been identified by the Technology Transition Team (Grosshandler et ai, 1994):

HFC-227ea, CF3I, and HFC- 125. In engine nacelle applications, the fire suppression agent is located

remotely from the nacelle, and a piping system delivers the agent to the nacelle injection location(s).

It is crucial to deliver the proper amount of agent in a timely manner to achieve the desired agent

concentration in the nacelle for the required time interval. The flow regime in the piping is characteristically a two-phase, two-component gas/liquid system. Since no universally accepted flov/ prediction

method is available for two-phase flows, an experimental study was undertaken to investigate the

effects of selected parameters on the flow of alternative agents. The work here focused on the

performance of alternative agents based on flow time through piping systems. Rapid transport to, and

effective mixing in the nacelle are important for an efficient design utilizing an alternative agent.

The initial selection criteria of the alternative agents included more than just environmental

concerns and suppression efficiency. NIST Special Publication 861 (Grosshandler et al., 1994) details the measurements and analyses performed by NIST that assisted in the selection of a reduced list of

alternatives. Detailed studies of pipe flow characteristics were not included, primarily because it represents a portion of the overall system performance which does not, to some extent, depend only on

the chemical properties of the agent. However, design parameters such as pipe diameter, length, and

storage conditions play a significant role in the system's discharge performance. In this study, the

effects of the following design parameters were examined: bottle fill condition and temperature, pipe

diameter, length, elbows, tees, contractions, and expansions. Steady pipe flow experiments were

performed by adding make-up nitrogen to the storage bottle as the agent was discharged. From, these

experiments, mass flow rate and pipe pressure drop as functions of fixed bottle pressure were obtained.

Optical access to the flowing agent in the pipe verified the nature of the flow.

8.6.2 Agent Discharge in Piping. The fact that there is two-phase flow of halon or of alternative

agents stems from the "optimized design" of a generic suppression system, expressed as the desire to
564 8. OPTIMIZATION OF SYSTEM DISCHARGE

get most of the agent to the nacelle very quickly, while minimizing the storage space. First, the bulk

of the agent is in the liquid phase in the storage bottle which drastically cuts down on the storage

volume. The liquid phase is subject to flashing during depressurization. The portion of the bottle volume that contains vapor is referred to as the ullage. Second, nitrogen gas is added to the storage

bottle, raising the ullage pressure above the saturation vapor pressure of the agent; the ullage pressure

is the main driving force for the expulsion of the agent from the bottle. Nitrogen dissolves in each of

the agents to some extent and can come out of solution during depressurization, initiating two-phase

flow. At this point, a qualitative description of halon or alternative agent pipe flow sets the stage for the more detailed discussion to follow. For now consider that the bottle is very large relative to the piping

so that conditions in the bottle do not change during the ensuing pipe flow. Thus, upstream conditions

are given by stagnation properties and the pipe flow is steady. Traversing down the long horizontal

pipe (with no significant entrance effects, valves, etc.) the pressure is dropping due to wall friction losses and acceleration of the expanding fluid. If the agent is originally saturated with nitrogen, a pressure decrease results in a supersaturated liquid and bubbles can form. Ignoring the effects of

bubble inception (nucleation) and growth, a differential amount of nitrogen will come out of solution

after a differential drop in pressure from the stagnation conditions (assuming an isothermal liquid

phase, so that Henry's law constant does not change). These bubbles will contain both nitrogen and

agent vapor even though the total pressure may still be above the saturation vapor pressure of the

agent. The quality of the flow (defined as the mass fraction in the gas phase) increases as the pressure

drops along the pipe. The two-phase flow is compressible; thus it is possible to exhibit choking. In gaseous flows, choking is the realization of a thermodynamic limit on the mass flux as the flow

achieves its sonic velocity; critical flow and choked flow have the same meaning. In two-phase gas/liquid flows choking may be realized at velocities below that of the gas-phase sonic velocity. If the

flow is choked, the mass flow rate is independent of the discharge pressure (here, the pressure of the

ambience).

An actual bottle discharge exhibits highly transient flow. At first, as the pipe is filling up with

agent, the pipe pressure at any point is rising rapidly while the bottle pressure is falling rapidly. The

pipe pressure achieves a maximum when the pipe just fills, then starts to fall along with the bottle pressure. As the bottle pressure drops, the dissolved nitrogen in the liquid phase may not come out of

solution immediately, resulting in a metastable, supersaturated solution. After some time delay,

degassing of the dissolved nitrogen can produce a net pressure rise inside the bottle due to the

expansion of a bubbly two-phase fluid. The bottle pressure may reach a local maximum then start to drop again as this two-phase fluid continues to empty. At the point where the last of the two-phase

mixture is being expelled from the bottle, a pressure rise is observed in the piping for a short period of

time. The fluid in the piping is changing from a low quality (mostly liquid) fluid to a high quality

(mostly gaseous) fluid, which exhibits a different pressure drop in the piping. At that point, most of

the agent has been discharged from the bottle and piping. The pressure in the piping starts dropping

again as the ullage contents discharge.

8.6.3 Single-Phase Flow. Single-phase steady pipe flows are well understood, and accurate flow

calculations can be readily performed. Single-phase flows are treated as either incompressible (i.e., liquids) or compressible (i.e., gases). As a rule of thumb, if the change in fluid density is greater than

10 % over the flow section of interest, the flow is treated as compressible. Incompressible flow is the

simplest case and is described below to illustrate the significant parameters affecting the flow

characteristics.
8. OPTIMIZATION OF SYSTEM DISCHARGE 565

As is characteristic of all fluid flows, the driving force is the pressure gradient. The pressure

force is needed to overcome the resistance to flow arising from fluid drag at the pipe walls. For both

laminar and turbulent flow, the following equation holds for flow in a straight horizontal pipe.

AP _ 2fpu^ (37^

I D

AP is the pressure drop over the pipe length of interest. L is the pipe length, D is the diameter, p is the fluid density, u is the mean velocity and /is the Fanning friction factor. The friction factor is a parameter that describes the fluid-wall interaction; it is possible to calculate the friction factor for

laminar flow from pipe geometry, flow conditions and fluid properties. For turbulent flow, the friction factor is treated empirically and is correlated with pipe geometry, flow conditions and fluid properties.

Correlations for the friction factor are cast in terms of the Reynolds number (a dimensionless group

formed by the ratio of the inertia force to the viscous force) and the relative roughness of the pipe,

e/D where e is the surface roughness and has been measured for many typical piping materials. The

Reynolds number also is indicative of the flow regime with laminar flow existing below Re ~ 2300

and turbulent flow above Re ~ 4000 with a transition region between 2300 and 4000 for pipe flows.

The concept of equivalent length is very useful in determining flow rates or pressure drops in realistic piping systems with fittings, bends, valves, etc. Simply put, the concept is to equate the observed

pressure drop in a pipe fitting with an equivalent length of pipe (L^^) with the same diameter and

fnction factor as the piping of interest. A complicated piping system is reduced to an equivalent

(fictitious) total pipe length. L^^/D can be obtained from head loss coefficients tabulated for many

different pipe fittings (SFPE Handbook, 1988). A loss coefficient divided by (2/) yields L^^D.

If the fluid is compressible, the density is a function of pressure, and the equation above is not

appropriate. The flow depends on the thermodynamic path the fluid follows. Usually, to simplify

analysis, a compressible flow is assumed to be either isothermal, adiabatic isentropic or adiabatic

isenthalpic, which fixes the thermodynamic path. The same friction factor correlations hold for compressible flow as those for incompressible flow, though the friction factor may change locally as

the flow progresses.

8.6.4 Two-Phase Flow. Two-phase gas/liquid flows are common in the chemical process industries

and in electric power plants (steam/water system). Design calculations for many steady two-phase

flow conditions are readily available, although the accuracy of two-phase flow correlations can be poor

(Olujic, 1985); all two-phase flow correlations rely on some empiricism. The lack of a generalized

physical model stems in part from the number of flow regimes possible in two-phase gas/liquid flows.

These flow regimes are descriptive of the phase separation and include bubbly, annular, wavy, slug

and mist flows. A correlation which is reasonable for wavy flow may be extremely poor for bubbly

flow, etc. A flow map, which is a graphical depiction of the conditions giving rise to the various flow

regimes, can be used to assist in selecting the best correlation equations, though it is best to have

direct knowledge of the flow regime if the conditions are uncertain. Depending on the flow regime, a

particular model may be preferred to describe the flow. The homogeneous flow model assumes that the two phases are well mixed and traveling with the same velocity which is reasonable for bubbly

flows. There are equilibrium and non-equilibrium homogeneous flow approximations. Separated flow

models, where the liquid and gas phases are not well mixed and may be travelling at different speeds,

characterize the rest of the cases. A good discussion of two-phase gas/liquid flow is presented by

Whalley (1990).
566 8. OPTIMIZATION OF SYSTEM DISCHARGE

The literature on transient two-phase flows mainly involves the high-pressure water/steam system.

This is due to the importance of nuclear reactor safety assurance during a loss-of-coolant-accident

(LOCA). Most of the literature on halon 1301 systems deals with design of systems for total flooding

applications with agent discharge times which are an order of magnitude longer than that required for engine nacelle fire protection application. A review of halon discharge testing and system design for

total flooding applications is presented by DiNenno and Budnick (1988). They also include an

extensive bibliography on two-phase flow in general and halon systems in particular. The most

comprehensive study detailing transient discharge of halon 1301 through piping systems reported in the literature is by Elliot et al. (1984).

Design calculation schemes for halon systems are for the most part proprietary. System designers

only have to provide a design that works for a given case, meaning design concentrations of agent are maintained for the appropriate times. This requires an experimental test to verify a specific design.

Three models of halon 1301 pipe flow described in the literature are mentioned below. The National

Fire Protection Association's Standard 12A "Halon 1301 Fire Extinguishing Systems" (National Fire

Codes, 1995), gives a simplified design method for systems that should meet the performance criteria for extended discharge into enclosed spaces. The pipe flow equation is an application of the steadystate mechanical energy and mass balances with a specified friction factor and data specific to halon 1301 inserted; homogeneous, equilibrium two-phase flow is assumed. Pipe flow calculations are based on a fixed average bottle pressure condition and adiabatic, isenthalpic flow. DiNenno and

Budnick (1988) point out potential accuracy and uncertainty problems in this method. This method was designed for extended discharges (discharge times of about 10 s). Elliot et al. (1984) use a homogeneous, equilibrium model of two-phase flow to solve for transient pipe and nozzle flow in halon systems. They also examined a Los Alamos National Laboratory code "Sola-Loop" developed

for the steam/water system which uses the "drift-flux model," a separated flow model, to predict the

transient discharge of halon. Both of their calculations compare favorably to some limited halon 1301

tests.These documented flow calculations do well for certain situations, but rely on data explicit to halon 1301, and make various assumptions about the flow. The observations and data from this study

verify some of the assumptions made, and their validity for the alternative agents. Here, a model that simulates the transient discharge of an agent through piping has been developed. It is similar to the

NFPA 12A methodology and Elliot's model in that it is based on a homogeneous, equilibrium twophase flow description. Thermodynamic properties and fill conditions are obtained from a separate

vapor-liquid equilibrium computer program called "PROFISSY" (see Section 8.3).

8.6.5 Experimental Apparatus. A range of test conditions was examined to better understand the nature of the alternative agent flows under conditions that may be encountered in suppression system

piping. Both transient and quasi-steady flow cases were investigated. The quasi-steady flow

experiments were performed to obtain the mass flow rate and steady pipe pressure drop for a fixed

discharge pressure. The experimental apparatus was designed so that different configurations and conditions could be

studied. A schematic diagram of the experimental apparatus is shown in Figure 97 and each of its components is described below. System constraints, specific to engine nacelle applications, limited the range of conditions investigated. Pipe diameters and lengths were selected based on a review of

military aircraft specifications. Also the fill condition and nitrogen pressurization were selected based on knowledge of typical halon 1301 engine nacelle systems.

8.6.5.1 Discharge Vessel and Piping Design. The discharge vessel was constructed from a

cylindrical tube of stainless steel with an internal diameter of 100 mm and a wall thickness of
8. OPTIMIZATION OF SYSTEM DISCHARGE 567

NITROGEN

MAKE-UP

TANK

ELECTRIC

SOLENOID

VALVE AGENT FILLING

NITROGEN

CHARGE

THERMOCOUPLES

COMMERCIAL

FREEZER

PRESSURE

TRANSDUCER

RECOVERY

TANKS

RELEASE

VALVE

TRANSFER

CYLINDER

Figure 97. Schematic of the apparatus for pipe flow experiments.
568 8. OPTIMIZATION OF SYSTEM DISCHARGE

12.7 mm. Flanges were welded to the two ends of the tube so that top and bottom plates, 19.1 mm

thick, could be attached with bolts (Figure 98). The seal between a flange and plate was made with an

0-ring. The vessel volume with the flanges attached is 4.06 x 10"^ m^. This volume was verified by

water displacement. The combined standard uncertainty in the volume measurement is estimated to be

less than ±10 cm^. The top plate had five threaded access holes for agent filling, two thermocouples,

a pressure transducer, and make-up nitrogen. The bottom plate had one threaded access hole for the

release mechanism. The vessel was designed to operate at pressures up to 14 MPa. There was a

pressure relief valve connected to the filling port to protect the vessel from unexpected high pressure

conditions.

The agent release mechanism was a quick-opening solenoid valve (Marotta Scientific Controls

Inc., Model MV121KJ-2) with an inlet diameter of 31.8 mm and an exit diameter of 44.5 mm. The

valve exit is at a 90° angle to the inlet. This valve can be easily re-armed for quick turn-around of

tests. The valve was activated by a DC power source of approximately 25 V and 10 A; the opening

time was on the order of 10 ms. Since the valve exit was much larger than the internal pipe diameters

used in this study, a smooth tapered connector was fabricated to make the internal diameter transition

from 44.5 mm to 15.9 mm (Figure 99).

The nitrogen make-up tanks were two standard high-pressure gas cylinders with a combined

internal volume of 88 x 10 m . The cylinder valves were removed and 19.1 mm I.D. tubing leading

to a large-orifice solenoid valve was used to deliver the make-up nitrogen. The valve was a pistonplunger type solenoid valve which required a pressure difference of 30 kPa to open after the solenoid was activated. The opening time is on the order of 50 ms for nitrogen flow at test condition pressures. The make-up nitrogen entered the vessel from an access hole in the top flange. There were two

reasons for using a solenoid valve here: first, the bottle contents must be separated from the nitrogen

make-up tanks or else agent vapor will mix with the nitrogen, and second, it was desired to shut down

the nitrogen flow after the agent discharge to avoid losing most of the make-up nitrogen and

pressurizing the entire system.

The piping was drawn stainless steel tubing. Two diameters were investigated, 15.9 mm and

9.5 mm inner diameter with tube wall thicknesses of 3.2 mm. Brass pressure taps were constructed

with the same inner diameter as each of the pipes (Figure 100). These taps allowed the pipe pressure

transducers to be installed flush to the tube wall. The transducers were mounted and maintained in the

horizontal orientation. An extra access port was provided for thermocouple installation. The piping

sections were connected with threaded fittings for easy reconfiguration, and rigidly attached to a bench

to limit motion during agent discharge.

The recovery tanks contained the agent after it flowed through the piping. They consisted of four

separate high pressure sample cylinders with a combined volume of 15 x 10"^ m'^, located in a commercial freezer and surrounded by dry ice. This allowed most of the agent to condense in the

sample cylinders after a test. Piping connected to the bottom of the cylinders led to a transfer cylinder

which was used to transfer the agent back to the discharge vessel for subsequent tests. Usually, 80 %

to 90 % of the agent was recovered with this system after a given test. This significantly reduced the

total amount of agent required to perform the series of tests.

8.6.5.2 Experimental Measurements. The liquid and gas temperatures in the vessel were

recorded by type E sheathed thermocouples prior to a test. Temperatures were not taken during tests because the thermal time lag through the sheath was longer than the discharge time. Attempts to measure the temperature of the fluid in the piping during discharge were unsuccessful because the

unsheathed fine wire (13 |jm diameter) thermocouple probes would break in the flow. The pressure in the storage vessel was recorded with a strain-gage-type transducer (Druck

Model 330) with a total accuracy of 0.75 % and an operating and compensated temperature range of
8. OPTIMIZATION OF SYSTEM DISCHARGE 569

Top

Plate

Flange (2)

Bottom

Plate

0-ring

Groove

Storage Vessel

100 mm

H H

Threaded for

Discharge Valve

500 mm

JO,

Figure 98. Schematic of the pressure vessel.
570 8. OPTIMIZATION OF SYSTEM DISCHARGE

flattened surface

tapped for pressure

transducer

63.5 38.1

mm

tapped

3/4" FNPT

tapped for

thermocouple

203 mm

Faced off

to seat gasket

1.875 -12 UNJ -3B

threaded

View a-a View b-b

Faced off

Figure 99. Schematic of valve reducer.
OPTIMIZATION OF SYSTEM DISCHARGE

Pressure Transducer

Figure 100. Schematic of pipe tap for pressure transducer and thermocouple.
572 8. OPTIMIZATION OF SYSTEM DISCHARGE

- 54 °C to 150 °C. The transducer's pressure range was 0 MPa to 14 MPa with a response time on

the order of 1 ms. Pressure in the piping was recorded by strain-gage type, flush-mount pressure

transducers (Omega Model PX-600) with a diaphragm face diameter of 7.9 mm. The transducers had

a nominal range of 0 MPa to 3.4 MPa or 0 MPa to 6.9 MPa, depending on the gage. The stated accuracy was 1 % over the compensated temperature range of 16 °C to 71 °C. The operable temperature range was - 54 ''C to 150 °C; a temperature correction must be applied to achieve the nominal

accuracy outside the compensated temperature range. The gage response time was on the order of

1 ms. A fast-response (2 kHz) signal conditioner/amplifier provided the boosted voltage signal that was recorded by the data acquisition system. In early trial experiments, piezoelectric-type dynamic

transducers were used in the piping. Pipe pressure traces from the dynamic transducers and the strain- gage type transducers were essentially the same for nitrogen flow, but for agent flow the results deviated near the end of the liquid two-phase expulsion. The deviation was attributed to a significant

non-compensated temperature effect experienced by the dynamic transducers. That problem eliminated

dynamic transducers from consideration for use.

Initially, the pipe pressure transducers were calibrated against the more accurate and stable storage

vessel transducer. This was done by filling the vessel and the piping with nitrogen to a number of

pressures and recording the output from each pipe transducer and comparing it to the vessel transducer

output. Periodic re-cahbrations indicated zero shifts for the pipe pressure transducers on the order of

± 30 IcPa, though very little change in calibration-line slopes was evident. The pipe pressure

transducers were subjected to water-hammer-type shock waves during flow start-up, which can sometimes cause a zero-shift. Therefore, the pipe pressure traces are offset to initial values of

0.101 MPa in the reported results. Transparent acrylic tube test sections with the same inner diameter as the larger diameter pipe

(15.9 mm) were fabricated and one transparent test section was hydrostatically tested to failure at a

pressure of 1 1 MPa. A section was connected to the piping at the upstream location right after the

valve reducer and was replaced after one or two tests. A high speed movie camera (Photec IV)

operating at 500 frames per second recorded the flow through the transparent section. Kodak

Ektachrome high-speed daylight film with intense lighting at a 90° angle to the camera was used. The

camera was turned on approximately 1 second before the discharge valve was activated to allow the camera motor to come up to the stated framing rate. The liquid two-phase flow portions of selected

tests were filmed and analyzed.

Data were acquired with a Strawberry Tree Flash- 12 Model 1 data acquisition board installed in a DOS-based personal computer. This board can scan a single channel at a rate of 1 MHz. Here the

multiple channels were scanned at 1 kHz since the response time of the pressure transducers is only

1 ms. A timing circuit triggered the data acquisition and sent signals to solid state relays to activate

the valves and movie camera if a high speed movie was taken. The raw data were stored in the

computer for subsequent analysis.

8.6,5.3 Test Matrix. The series of tests included a number of transient discharges with various

initial vessel conditions. The effects of agent temperature, fill volume, and vessel pressure were

investigated. Piping configurations included 3.5 m of straight pipe, straight pipe with a 90° angle

bend, tee, pipe expansion, and pipe contraction. Figure 101 details the piping configurations.

Constant-head tests were performed with fixed 3.5 m of piping at 3 constant vessel pressures:

3.0 MPa, 2.5 MPa, and 2.0 MPa. Table 34 details the nominal conditions for the separate tests. A

total of 21 different tests were performed for each agent. A limited number of repeated tests at slightly different fill conditions were also performed. The initial pipe and recovery tanks pressure were at atmospheric pressure except for one test where the pipe and recovery tanks were evacuated.
8. OPTIMIZATION OF SYSTEM DISCHARGE 573

Flow

Configuration

P P P P

3.5 m straight

Contraction from 15.9

to 9.5 mm I.D.

Figure 101. Piping configurations. Pressures were recorded at tap locations marked with a

"p"
574

Table 34. Test Matrix

8. OPTIMIZATION OF SYSTEM DISCHARGE

raralllClcr ICdlcU

Pipe geometry Initial storage vessel conditions

Diameter (mm)Nominal

length (m)

Pressure (MPa)

Liquid fill volume

(%)

Temperature

(°C)

Constant-head 9.5 and 15.9 3.0 2.0, 2.5 and 3.0 N.A. 23

Straight pipe 9.5 and 15.9 3.0 2.75 and 4.12 50 and 67 23

Storage temperature15.9 3.0 4.12 50 - 45 and 71

90° Bend 9.5 and 15.9 3.0 and 4.0 4.12 50 23

Tee 15.9 4.0 4.12 50 23

Expansion 9.5 to 15.9 3.0 4.12 50 23

Contraction 15.9 to 9.5 3.0 4.12 50 23

Pressure taps with installed transducers were placed at four locations in the piping. The first pipe

pressure tap was in the 15.9 mm diameter section of the valve reducer. For straight piping the

pressure was recorded at 1 .00 m intervals from the valve reducer. Since the distance from the valve

reducer to the flow split in the recovery tanks was 3.50 m, the last pressure recording was at the

3.00 m location. For the 3.50 m of piping with a 90° angle bend, the taps were located at 1.00 m

intervals starting from the valve reducer. For the 4.50 m of piping with a 90° angle bend, the taps were located at the valve reducer, and 1 .00 m, 2.00 m, and 4.00 m downstream. For the flow-splitting

tee experiments, the taps were located at the valve reducer and at 2.00 m downstream before the tee. The tee was located 2.10 m downstream, and the last two taps were place in the opposing pipes

0.50 m from the tee. Pressure in the recovery tanks was monitored in about half of the tests. Cold storage vessel tests were performed with the large diameter straight piping configuration.

An insulated sheet-metal sleeve was placed around the vessel and filled with dry ice to cool the liquid contents to about - 45 °C (Figure 102). First the vessel was filled to 1/2 liquid volume at 23 °C and

pressurized with nitrogen to a total pressure of 4.12 MPa, then dry ice was added to the cooling jacket.

When the temperature of the liquid reached about - 45 °C, the agent was discharged.

High storage vessel temperature tests were also performed with the large diameter straight piping

configuration. Electrical heating tape was wrapped around the vessel to heat up the vessel and

contents (Figure 102). First, the vessel was filled to 1/2 liquid fill volume at 23 °C and pressurized

with nitrogen to 4.12 MPa, then power was supplied to the heating tape. When the temperature of the

liquid reached about 70 °C, the agent was discharged. Pipe pressures were recorded at the valve

reducer section and the last pressure tap 3.00 m downstream. The pipe pressure transducers for the

elevated temperature discharge have a nominal range of 0 MPa to 6.9 MPa.

8.6.6 Results and Discussion. The initial fill conditions for each test are given in Tables B-1 to B-5

(see Appendix B). The temperatures of the liquid and gas phase in the storage vessel were between

21 °C and 23° C (23 °C being the ambient room temperature) for the room temperature fill conditions.

The mass of agent and pressure were recorded prior to each test. The combined standard uncertainty

in the mass determination is estimated as ± 0.01 kg based on the load cell uncertainty of ± 0.005 kg
Figure 102. Agent cooling and heating techniques for the cooled and heated agent discharge tests.
576 8. OPTIMIZATION OF SYSTEM DISCHARGE

and the fact that typically three agent transfers were required to fill the storage vessel. The combined

standard uncertainty in the tabulated pressures is ± 0.005 MPa due to rounding off the third digit past

the decimal and not due to the stated uncertainty of the transducer. The computer program

"PROFISSY" was used to estimate the fill condition and mass of the liquid and gas phase assuming

equilibrium conditions were achieved at a temperature of 23 °C.

Temperature equilibration was normally achieved in about 1 hour. Slow bubbling of the nitrogen

through the liquid phase (typically 30 min in duration) greatly facilitated nitrogen dissolution. For some storage vessel fills, the agent/nitrogen mixture was allowed to sit for 60 h or longer. The

pressure drop over such a long time is indicative of the level of nitrogen saturation achieved by the

bubbling technique. Two equilibrium calculations (using PROFISSY) were performed based on the

agent mass and the pressure achieved just after bubbling and after a long time storage. Assuming the

system had equilibrated after the long storage, the nitrogen saturation level achieved from bubbling

alone was approximately (90 ± 5) % of the mass required for saturation. HFC- 125 achieved the

lowest level of nitrogen saturation from the bubbling technique, while HFC-227ea and CF3I nearly

achieved nitrogen saturation from bubbling alone. The test results did not appear to be strongly

influenced by differences in nitrogen saturation level on the order of 10 %.

Pressure traces for all tests are shown in Appendix B. The results from the different configurations are discussed below.

8.6.6.1 Constant-Head Discharge. The three selected pressures (3.0 MPa, 2.5 MPa and

2.0 MPa) cover the bottle pressure range where most of the liquid would be discharged from the bottle for a typical transient discharge. Figure 103 (Figure B-2 in Appendix B) is a typical constant-head

pressure trace for the small diameter piping. The agent was halon 1301 with the vessel pressure fixed

at 2.5 MPa, and the pipe pressure was initially atmospheric. At approximately 1000 ms the discharge

valve opened. The vessel pressure dropped sharply then recovered due to the nitrogen flow from the

make-up tanks. The make-up tank solenoid valve was set to open 20 ms after the discharge valve

opened (the solenoid valve requires a pressure difference to open). The make-up tanks were filled to the initial vessel pressure and quickly re-pressurized the vessel. The vessel pressure remained nearly constant until the make-up nitrogen valve was closed about 2.5 s after it was opened. The pipe

pressure traces rose sharply after the discharge valve opened, reaching a maximum about 100 ms later and remained nearly constant until the liquid ran out of the vessel. As the liquid runs out of the

vessel, the pressure starts to rise in the pipe, then falls as the pipe empties the two-phase fluid contents. After the make-up tank valve was closed, the bottle and pipe pressures continue to drop

until the pressure equilibrated about 3.5 s after the beginning of the discharge.

Figure 104 (Figure B-14 in Appendix B) is a typical constant-head pressure trace for the large

diameter piping. The agent was halon 1301 with a vessel pressure fixed at 2.0 MPa, and the initial pipe pressure was atmospheric. The pipe pressure traces rise fast and take about 300 ms to reach

steady values. The crossover of the third pressure tap's trace with that of the fourth tap's trace is due

most likely to an incorrect time response of the gage at the third tap during the transient pipe filling

period. The recovery tank pressure starts to climb much slower than the pipe pressures and ends up at the final pressure. Notice that the increasing recovery pressure has no noticeable effect on the pipe

pressure trace until after the two-phase fluid runs out of the bottle and gas is being discharged. The

pressure equilibrated at about 2.0 MPa.

Experimental time-averaged mass flow rates were estimated by dividing the initial liquid-phase mass (calculated by PROFISSY) by the time increment from the initial rise in the pipe pressure to the

first indication that the liquid has run out of the vessel. The combined standard uncertainty in the

experimental mass flow rate was estimated by the RSS method (root-sum-of-the-squares) for the

standard uncertainty of the time increment and the liquid mass. Here the standard uncertainty of the
OPTIMIZATION OF SYSTEM DISCHARGE 577

T—I—I—I—

I—I—I—I—I—

I—I—

r

Discharge

^ valve open

End of liquid

Pressures

equilibrate

J I I I I I I I I I I I I I I I I I I I I I I I I I I I L

1000 2000 3000 4000 5000 6000

Time (ms)

Figure 103. Small diameter piping (9.5 mm I.D.) constant-head discharge test. The agent was

halon 1301 with a vessel pressure of 2.5 MPa, and piping initially at atmospheric.
578 8. OPTIMIZATION OF SYSTEM DISCHARGE

J I I I I I I I I I I I I I I I I I I I I I I I I I I I L

0 500 1000 1500 2000 2500 3000

Time (ms)

Figure 104. Large diameter piping (15.9 mm I.D.) constant-head discharge test. The agent was

halon 1301 with a vessel pressure of 2.0 MPa, and piping initially at atmospheric.
8. OPTIMIZATION OF SYSTEM DISCHARGE 579

liquid-phase mass is estimated to be equal to the standard uncertainty of the total agent mass. The

results, including the uncertainty estimates are reported in Table 35. The large diameter pipe mass flow rates are 3 to 4 times greater than the small diameter pipe

results. Theoretically, during the filling and emptying of the pipe, the mass flow rate should be higher

than the steady value; there is a shorter length of pipe for the contents to flow through. Since the

filling and emptying times of the large pipe tests are a significant fraction of the total flow time, the

estimated average flow rate over the entire test will be higher than the flow rate during the steady-flow

portion of the test. The constant-head discharge tests yield steady pipe flow conditions from 0.5 s to over 4 s, depending on the fill condition and pipe diameter. There are no observed effects of the recovery tanks

pressure on the two-phase flow during the tests. The flatness of the pressure traces suggests that transient heat transfer from the pipe walls to the two-phase fluid was not significant. Nonlinearity of

the pressure drop in the piping is indicated by the variation in the pressure difference between the

equally spaced pipe taps, which implies compressible flow. Another indication that the flow is compressible is that if the flow is assumed to be incompressible, calculated mass flow rates from

Equation (37) using the initial densities are on the order of 2-3 times greater than the experimentally

determined mass flow rates.

8.6.6.2 Transient Discharge. The transient discharge tests simulated two-phase pipe flow of

realistic systems. A typical ambient temperature, straight piping test result is shown in Figure 105

(Figure B-41 in Appendix B). Initially, the vessel was filled with HFC-125 to 1/2 liquid fill volume,

pressurized to 4.12 MPa with the pipe (0.0159 m I.D.) and recovery tanks at atmospheric pressure. At

approximately 1000 ms the discharge valve was opened, and the vessel pressure started to drop while

the pipe and recovery tanks pressures started to rise. The pipe pressures peak, then start to fall with

the vessel pressure. Notice that one trace, representing the second transducer, crossed over the third

and fourth traces. As stated above, this effect is most likely due to that transducer's incorrect response

to the highly transient conditions; it is apparent in the large diameter tests, though not apparent in the

small pipe diameter tests. The vessel and pipe pressures stop decreasing after about 1400 ms, achieve

a local maximum, and then resume decreasing. This pressure recovery is thought to be due to degassing of nitrogen in the storage bottle. At about 1800 ms the vessel pressure starts decreasing at a

faster rate, which is attributed to the discharge of the ullage contents after the two-phase mixture

leaves the vessel. At the same time, pipe pressures start to increase, peak, then decrease at a much

faster rate which is indicative of the two-phase mixture leaving the piping followed by discharge of the

ullage contents. All pressure traces equilibrate to about 1.2 MPa. The final equilibrated pressure for

the transient tests ranged from 0.5 MPa to 1.5 MPa, depending on the agent, fill conditions, and piping

configuration.

Experimental liquid discharge times were estimated for each test as the time interval from the

initial vessel pressure drop to the final hump in the last pipe pressure trace. There is some uncertainty

in the liquid discharge time since the last amount of liquid draining from the vessel will not flow out

as a plug, but will be entrained into the discharging ullage gas. The bracketed time is estimated to represent lower and upper limits of the liquid discharge time, with the best estimate given by the mean

value, and the standard uncertainty given by the half width of the time interval divided by the square

root of 3 (Taylor and Kuyatt, 1994). The results for the experimental liquid discharge times are presented in Tables 36-39. The discharge times for the large diameter pipe were approximately 2.5

times shorter than the small diameter pipe.

The effects of the increasing recovery tanks pressure on the transient flow are identifiable, given

the pipe and recovery tanks pressure traces. If the pipe exit pressure is above the discharge pressure

(recovery tanks pressure), the flow remains choked, and the flow rate is independent of the discharge
580 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 35. Constant-head mass flow rates

Agent/

figure number

Pressure

(± 0.005 MPa)

Pipe

I.D. (mm)

Liquid phase mass

(+ 0 01 ko)

Flow time

(± 0.025 s)

Mass flow

rate (ks/s)

Calculated mass flow

rate (kg/s)

CF^Br/Bl 2.00 9.5 3.00 2.95 1.02 (0.01) 1.09

CF3BR/B2 ^ 2.50 9.5 3.15 2.50 1.26 (0.01) 1.28

CF:,Br/B3 3.00 9.5 3.09 2.0,1 1.55 (0.02) 1.50

CF,Br/B14 2.00 15.9 3.87 1.00 3.89 (0.10) 3.39

CFjBr/BlS 2.50 15.9 3.57 0.85 4.20 (0.12) 3.92

CFjBr/Bie 3.00 15.9 4.23 0.93 4.55 (0.12) 4.51 HFC 125/B26 2.00 9.5 2.31 2.05 1.13 (0.01) 1.06 HFC 125/B27 2.50 9.5 2.75 2.50 1.10 (0.01) 1.14 HFC 125/B28 3.00 9.5 2.32 1.80 1.29 (0.02) 1.33 HFC 125/B37 2.00 15.9 2.54 0.77 3.30 (0.11) 3.07 HFC 125/B38 2.55 15.9 3.15 0.70 4.50 (0.16) 3.51 HFC 125/B39 3.00 15.9 2.63 0.60 4.38 (0.18) 3.98 HFC 227ea/B51 2.00 9.5 3.24 2.45 1.32 (0.01) 1.36 HFC 227ea/B52 2.50 9.5 3.10 2.05 1.51 (0.02) 1.51 HFC 227ea/B53 3.00 9.5 3.16 2.00 1.58 (0.02) 1.70 HFC 227ea/B62 2.00 15.9 3.68 0.70 5.25 (0.19) 4.25 HFC 227ea/B64 2.50 15.9 3.59 0.70 5.13 (0.18) 4.81 HFC 227ea/B65 3.00 15.9 3.49 0.65 5.37 (0.21) 4.93

CF3I/B77 2.00 9.5 5.32 3.50 1.52 (0.01) 1.61

CF3I/B78 2.50 9.5 5.50 3.00 1.83 (0.02) 1.95

CF3I/B79 3.00 9.5 5.13 2.60 1.97 (0.02) 2.11

CF3I/B88 2.00 15.9 3.92 0.60 6.53 (0.27) 5.16

CF3I/B89 2.50 15.9 4.10 0.60 6.83 (0.29) 5.94

CF3I/B9I 3.00 15.9 4.20 0.60 7.00 (0.29) 6.45

'Bracketed numbers are combined standard uncertainties (2o) for experimental mass flow rates.

pressure. If, however, the pipe exit pressure is equal to the discharge pressure, then the flow rate depends on the discharge pressure, and it will not achieve the maximum choked value. For these tests, the pipe exit pressure is somewhat lower than that recorded by the last pressure tap (the pipe

terminates at the recovery tanks some distance downstream from the last pressure tap), so if that trace follows the trend of the recovery tanks pressure it is evidence that the flow is not choked at the exit.
8. OPTIMIZATION OF SYSTEM DISCHARGE 581

1—I—I—I—

J—I—I—I—I—

I—I—I—I—I—

I—I—I—I—I—I—I—I—I—r-|—I—I

I I I I I III I I I I I I I I I

I I I I I I I I—I—I—I—

L

0 500 1000 1500 2000 2500 3000

Time (ms)

Figure 105. Large diameter piping (15.9 mm ID.) transient discharge test. The configuration was

straight piping. The vessel was filled to 1/2 liquid fill volume with HFC- 125 and

pressurized to 4.12 MPa.
582 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 36. Liquid discharge times for halon 1301

Pipe geometry Fill conditions Figure td number (exp.) (calc.)

(s) (s)

Piping Diameter Pressure Liquid Equil. Non-Eq.

fill vnl ('7r,^ 1111 VUl. \ 'C

)

Straight 9.5 2.77 46.6 B-4' 3.3 - 3.7 3.41 3.41

15.9 2.75 51.1 B-17* 1.4 - 1.7 1.46 1.46

9.5 4.12 50.4 B-8* 2.5 - 2.8 2.79 3.02

15.9 4.12 49.6 B-19 1.0 - 1.2 1.03 1.08

9.5 4.12 65.7 B-9* 3.6 - 4.0 4.23 4.48

15.9 4.12 65.5 B-20' 1.4 - 1.6 1.55 1.66

3.5 m,90° 9.5 4.12 48.9 B-10* 2.9 - 3.1 2.75 2.95 Bend

15.9 4.12 48.7 B-21 1.1 - 1.3 1.02 1.05

4.5 m, 90° 9.5 4.12 49.1 B-lT 3.0 - 3.4 3.09 3.32 Bend

15.9 4.12 48.7 B-22 1.1 -1.3 1.14 1.16

Expansion 9.5/15.9 4.12 48.3 B-13 2.1 - 2.4 1.84 1.98

Contraction 15.9/9.5 4.12 48.3 B-12 2.4 - 2.8 2.30 2.43

tee 15.9 4.12 47.9 B-23 1.0 - 1.2 1.00 1.05

straight 15.9 2.20^ 41.8 B-24 1.4 - 1.5 1.09 1.10

*Two-phase flow appears effected by recovery tanks pressure ^Coid temperature discharge

To simulate the flow of the two-phase mixture discharging to the atmosphere, it is only necessary to delay the effect of the recovery tanks pressure increase until the ullage contents start to discharge,

ensuring that the two-phase flow is always choked. In some cases it appears that the recovery tanks

pressure is high enough to affect the two-phase flow from the piping, occurring at the later stage of

the two-phase discharge. Figure 106 (Figure B-54 in Appendix B) is an example of this effect. The

agent is HFC-227ea and the fill conditions are 1/2 liquid fill volume, pressurized to 2.75 MPa. The

configuration was straight, small diameter piping. The pressure trace for the last tap location shows a gradual pressure increase starting at about 4500 ms. This behavior is most evident for the following

conditions: high fill volume, small diameter piping and agents with lower saturation vapor pressures

(HFC-227ea and CF3I). If the flow is not always choked at the pipe exit, the flow time will be

slightly longer than the flow time for an atmospheric discharge. The tests that appear to be affected

by the finite size of the recovery tanks are indicated in Tables 36-39. This effect probably results in not more than a 10 % increase in flow time for any affected test. In some experiments, like the one described in Figure 105, the storage vessel pressure experiences

a recovery to a local maximum. It is most pronounced for halon 1301 and HFC- 125 2/3 fill
8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 37. Liquid discharge times for HFC- 125

583

Pipe geometry Fill conditions Figure

'd number (exp.) (calc.)

(s) (s)

Piping Diameter Pressure Liquid Equil. Non-Eq.

config. (mm) (MPa) fill vol. (%)

Straight 9.5 2.76 50.4 B-30* 3.1 - 3.3 3.35 3.48

15.9 2.75 54.2 B-40 1.3 - 1.5 1.40 1.43

9.5 4.13 56.4 B-31* 2.7 - 2.9 2.79 3.34

15.9 4.12 52.7 B-42 0.9 - 1.1 0.99 1.08

9.5 4.12 69.8 B-32 3.4 - 3.7 4.19 4.65

15.9 4.12 65.9 B-44' 1.4 - 1.5 1.39 1.59

3.5 m,90° 9.5 4.12 47.1 B-33 2.6 - 2.8 2.28 2.53 Bend

15.9 4,12 48.8 B-46 0.9 - 1.0 0.90 0.99

4.5 m, 90° 9.5 4.12 48.1 B-34 2.6 - 2.8 2.63 2.92 Bend

15.9 4.12 46.6 B-47 0.9 -1.0 0.95 1.02

Expansion 9.5/15.9 4.12 49.6 B-35 1.9 - 2.1 1.67 1.91

Contraction 15.9/9.5 4.12 49.8 B-36 2.2 - 2.4 2.10 2.34

tee 15.9 4.12 54.4 B-48 0.9 - 1.0 1.04 1.18

straight 15.9 2.55' 39.1 B-49 0.9 - 1.0 0.80 0.82

'Two-phase flow appears effected by recovery tanks pressure ^Cold temperature discharge

conditions. It is not strong for HFC-227ea and CF3I and is only indicated in one trace each for those

agents. No visual observations of the fluid inside the vessel were made during these tests. It has been

surmised that the pressure recovery is due to nitrogen coming out of a supersaturated solution, which

in turn increases the liquid two-phase volume compressing the ullage gas and raising the vessel

pressure. The exact mechanism(s) for bubble formation is not known and it is possibly a combination

of heterogeneous and homogenous nucleation. There is scatter in the lowest pressure reached before

the pressure rise for any agent. It appears to depend on the fill volume, vessel pressure and pipe

diameter. If it was purely homogenous nucleation, it should only depend on the initial pressure. This

degassing phenomenon has implications for the discharge time and is explored in the model calculations to follow.

8.6.6.3 Straight Pipe Configuration. The straight pipe configuration was the baseline for the

tests at different fill conditions and vessel pressures. Figure 101(a) shows the configuration. Typical

results are shown in Figures 105 and 106. Two fill conditions (nominally 1/2 and 2/3 liquid filled) at 4.12 MPa and two vessel pressures (4.12 MPa and 2.75 MPa) at 1/2 liquid filled were examined. The

liquid discharge time was always longer for the 2/3 liquid filled condition compared to the 1/2 liquid
584 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 38. Liquid discharge times for HFC-227ea

Pipe geometry Fill conditions Figure

'd I d number (exp.) (calc.)

(s) (s)

Piping Diameter Pressure Liquid Equil.

config. (mm) (,Mra) fill urtl /C7 \ nil vol. ( /c)

Straight 9.5 2.75 56.4 B-54* 4.2 - 4.4 3.44 3.81

15.9 2.75 54.1 B-66 1.2 - 1.4 1.19 1 01

9.5 4.12 53.7 B-55* 2.9 - 3.1 2.43

15.9 4.12 49.4 B-68 0.8 - 0.9 0.80

9.5 4.12 70.0 B-57' 4.1 - 4.4 3.90 4 80

15.9 4.12 69.4 B-70* 1.7 - 1.9 1.40 1.64

3.5 m,90° 9.5 4.12 49.2 B-58 2.5 - 2.7 2.19 9 31 Bend

15.9 4.12 49.4 B-71 0.8 - 1.0 0.82 0.86

4.5 m, 90° 9.5 4.12 49.2 B-59 2.6 - 2.9 2.47 2.61 Bend

15.9 4.12 48.8 B-72 0.9 - 1.0 0.90 0.94

Expansion 9.5/15.9 4.12 49.0 B-61 1.8 - 2.0 1.46 1.55

Contraction 15.9/9.5 4.12 49.0 B-60 2.1 - 2.3 1.81 1.90

tee 15.9 4.12 48.4 B-73 0.8 - 0.9 0.79 0.82

straight 15.9 3.00^ 43.5 B-74 0.9 - 1.0 0.75 0.76

15.9 5.95' 59.8 B-76 0.8 - 0.9 0.90 1.03

Two-phase flow appears effected by recovery tanks pressure ^Cold temperature discharge *Hot temperature discharge

filled condition. The relative increase depends on the particular agent. The liquid discharge times for

the 2.75 MPa charge pressures were 30 % to 50 % greater than the 4.12 MPa charge pressures

8.6.6.4 Piping with a 90° Angle Bend. Tests with 3.5 m of piping with a smooth-bend-radius

90° angle bend located 2.5 m downstream, and 4.5 m of piping with a smooth-bend-radius 90° angle

bend located 3.5 m downstream (see Figures 101(b) and (c)) were performed. Comparing the straight

piping resuhs to the 3.5 m piping with a 90° angle bend, the liquid discharge times and the pressure

drops are longer and lower, respectively, for the configurations with the elbow compared to the

straight piping results. The 4.5 m piping with a 90° angle bend displayed still longer liquid discharge

times, due to the effects of the bend and the extra 1 .00 m of piping.

8.6.6.5 Straight Pipe with a Sudden Expansion. The sudden expansion configuration consisted

of 1.00 m of straight 9.5 mm I.D. piping connected to 2.5 m of straight 15.9 mm I.D. piping (see

Figure 101(d)). Figure 107 (Figure B-13 in Appendix B) is a typical pressure trace. Following the

I
8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 39. Liquid discharge times for CF3I

585

Pipe geometry Fill conditions Figure

'd number (exp.)

(s)

(calc.)

(s)

Piping Diameter Pressure Liquid Equil. Non-Eq.

config. (mm) yivlrd) fill vnl

Straight 9.5 2.75 49.7 B-80* 4.1 - 4.3 3.35 3.57

15.9 2.80 49.6 B-92 1.3 - 1.4 1.26 1.32

9.5 4.13 50.2 6-81* 3.4 - 3.6 2.61 2.76

15.9 4.12 48.7 B-93 1.0 - 1.1 0.94 0.98

9.5 4.12 67.9 B-83' 5.1 - 5.3 4.38 5.10

15.9 4.12 68.2 B-94* 1.9 - 2.1 1.62 1.84

3.5 m,90° 9.5 4.12 52.0 B-84' 3.6 - 3.8 2.84 3.02 Bend

15.9 4.12 49.4 B-95 1.0 - 1.1 0.98 1.02

4.5 m, 90° 9.5 4.12 51.2 B-85* 3.8 - 4.0 3.12 3.29 Bend

15.9 4.12 55.1 B-96 1.2 -1.4 1.29 1.35

Expansion 9.5/15.9 4.12 53.8 B-87 2.5 - 2.6 2.01 2.15

Contraction 15.9/9.5 4.12 52.4 B-86 2,9 - 3.1 2.40 2.53

tee 15.9 4.12 49.9 B-97 0.8 - 1.0 0.98 1.02

straight 15.9 3.05^ 44.2 B-98 1.1 - 1.2 0.95 0.96

15.9 5.70* 56.7 B-99 0.9 - 1.0 1.00 1.09

*Two-phase flow appears effected by recovery tanks pressure ^Cold temperature discharge *Hot temperature discharge

first pipe tap's pressure trace, after it initially rises, it tracks the vessel pressure trace very closely, with

a small pressure difference between the two. This is common to all tests where a section of small

piping follows the valve reducer. The reason these two pressure traces track so closely is because the

first pressure tap is located in the valve reducer where the inner diameter is 15.9 mm and most of the

flow resistance occurs in the small diameter piping located 0.05 m downstream. There is a large

pressure drop from the valve reducer to the next pressure tap 1 .00 m downstream. The next two

pressure records 2.00 m and 3.00 m downstream are only slightly different, and after some time they

follow the recovery tanks pressure trace. These observations suggest that the flow does not choke at the recovery tanks entrance, but at the expansion point. This implies that liquid discharge time to the

end of the first 1 .00 m of piping would be equivalent to the liquid discharge time through a 1 .00 m

section of piping discharging to atmosphere.

8.6.6,6 Straight Pipe with a Sudden Contraction. The sudden contraction configuration

consisted of 1.00 m of straight 15.9 mm I.D. piping connected to 2.00 m of straight 9.5 mm I.D.
586 8. OPTIMIZATION OF SYSTEM DISCHARGE

' ' End of liquid

' I

I I 1 I I I I

' ' ' I ' I I I I I I I I L

1000 2000 3000 4000 5000 6000

Time (ms)

Figure 106. Small diameter piping (9.5 mm I.D.) transient discharge test. The configuration was

straight piping. The vessel was filled to 1/2 liquid fill volume with HFC-227ea and

pressurized to 2.75 MPa.
OPTIMIZATION OF SYSTEM DISCHARGE 587

^ 3

0)

o

T—I—I—I—I—I—I—I—I—!—I—I—I—

T

T—1—r-1\—r—

r

Discharge

valve open

Storage vessel

tap #1

tap #2

tap #3

tap #4

Expansion after

second tap

location

J I L

Recovery tanks End of liquid

pressure

J I I I I I I I I I I I I I I I I I I r L

0 1000 2000 3000 4000 5000 6000

Time (ms)

Figure 107. Transient discharge test; straight pipe with a sudden expansion. The vessel was

filled to 1/2 liquid fill volume with halon 1301 and pressurized to 4.12 MPa.
588 8. OPTIMIZATION OF SYSTEM DISCHARGE

piping, then connected back to 0.5 m of 15.9 mm I.D. piping leading to the recovery tanks (see

Figure 101(e)). Figure 108 (Figure B-35 in Appendix B) is a typical trace. The first two pipe tap

pressure traces rapidly rise to just below the bottle pressure; they then track the bottle pressure for the

rest of the test. The first two traces follow the bottle pressure because the taps are located in the

15.9 mm I.D. valve reducer and piping which are before the contraction. The next two traces show

significant pressure drop from the bottle pressure. These traces are for the taps 2.00 m and 3.00 m

downstream located in the 9.5 mm I.D. piping. The liquid discharge time for this configuration should

be approximately equivalent to the discharge time for a 2.00 m section of 9.5 mm I.D. piping leading

from the vessel to the atmosphere.

8.6.6.7 Tee Configuration. The tee configuration consisted of 2.10 m of 15.9 mm I.D. piping

leading to a tee fitting, connected to two legs, one 1.80 m long and the other 2.05 m long, both with 2

smooth radius 90° angle bends leading to the recovery tanks (see Figure 101(f)). Figure 109

(Figure B-73 in Appendix B) is a typical example. The first pipe tap's pressure trace rises quickly

then follows the vessel pressure with some pressure difference. The second pipe tap's pressure trace

(the tap is located just before the tee) rises quickly, then falls much slower than the vessel or the first tap's pressure trace until the hump signifying the liquid emptying from the pipe. The pressure traces for the taps located in the tee legs 0.50 m downstream of the tee essentially follow the second pipe

tap's trace with some pressure drop. The slight difference in the two tee leg traces is most likely due

to the slight difference in length to the recovery tanks.

8.6.6.8 Cooled Agent Discharge. Cooled agent discharges were performed with 3.5 m of

straight 15.9 mm I.D. piping (Figure 101(a)). Table B-1 details the initial fill conditions, final test pressures, and the equilibrium conditions. Thermal equilibrium was never achieved between the liquid

and gas phases due to conduction of heat from the gas phase. Liquid phase temperatures ranged from - 44 °C to - 50 °C, while gas phase temperatures were around 4 °C. Though thermal equilibrium was

not achieved, the final pressures were only approximately 10 % higher than predicted equilibrium

pressures based on the liquid-phase temperature for all of the agents. Figure 110 (Figure B-98 in Appendix B) is a typical trace. The pressure trace results appear similar to ambient temperature

discharges at low storage vessel pressure. The longer liquid discharge times for the cooled agent

discharge appear to be due to the lower initial vessel pressure. The first pipe tap's pressure trace shows a temperature effect. The pressure tap is located in the

valve reducer which is in contact with the vessel via the discharge valve. Cold ambient temperature

affects the zero and the slope of the pressure transducer calibration. (Calibration of the gage recording

the recovery tanks pressure was done at the cold ambient temperature). The other pressure taps are located far downstream and do not appear to be cooled below ambient conditions appreciably during

the agent chilling period.

8.6.6.9 Heated Agent Discharge. High temperature agent discharges were performed with

3.5 m of straight piping 15.9 mm I.D. (Figure 101(a)). Table B-1 details the initial fill conditions, the

final temperatures and pressures, and calculated equilibrium conditions based on the final temperature.

The final temperature achieved in all of the tests was approximately 70 °C. PROFISSY calculations

indicate that two of the agents (halon 1301 and HFC- 125) are single-phase fluids at this elevated

temperature: the bottle conditions are given in Table B-1. The temperatures inside the bottle as measured with the two installed thermocouples were within 5 °C. Temperature differences could be

due to non-equilibrated gas and liquid phases or a temperature gradient in the fluid. The results are shown in Figures 111-114 (Figures B-25, B-50, B-75 and B-99 in Appendix B) for halon 1301,

HFC- 125, HFC-227ea and CF3I respectively. The pipe pressure traces for halon 1301 and HFC- 125
OPTIMIZATION OF SYSTEM DISCHARGE 589

Discharge

0 1000 2000 3000 4000 5000 6000

Time (ms)

Figure 108. Transient discharge test; straight pipe with a sudden contraction. The vessel was

filled to 1/2 liquid fill volume with HFC- 125 and pressurized to 4.12 MPa.
590 8. OPTIMIZATION OF SYSTEM DISCHARGE

T—T—I—I—

]—I—I—I—I—I—I—I

m—

I—I

I—I—I—

I—I—r—1—I—

I—I—I—I—

r

0 500 1000 1500 2000 2500 3000

Time (ms)

Figure 109. Transient discharge test; tee configuration. The vessel was filled to 1/2 liquid fill volume with HFC-227ea and pressurized to 4.12 MPa.
8. OPTIMIZATION OF SYSTEM DISCHARGE 591

T—I—I—I—

I—I—I—I—I—

I—I—I—I—I—

I—I—I—I i I—I—I—I—I—I I I

I I I I I ' L_l I I I I 1 1 1 1 1 1 1 1 1 1 L_J.

0 500 1000 1500 2000 2500 3000

Time (ms)

Figure 110. Transient cooled agent discharge test; straight piping. The vessel was filled to 1/2

liquid fill volume with CF3I, pressurized to 4.12 MPa, then cooled to - 45 °C or less

prior to discharge.
592 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 111. Transient heated agent discharge test; straight piping. The vessel was filled to 1/2

liquid fill volume with halon 1301, pressurized to 4.12 MPa, then heated to 70 °C

prior to discharge.
OPTIMIZATION OF SYSTEM DISCHARGE 593

0 500 1000 1500 2000 2500 3000

Time (ms)

Figure 112. Transient heated agent discharge test; straight piping. The vessel was filled to 1/2

liquid fill volume with HFC- 125, pressurized to 4.12 MPa, then heated to 70 °C

prior to discharge.
8. OPTIMIZATION OF SYSTEM DISCHARGE

T—I—I—r-|—I—I—I—r—

I—I—I—I—I—

I—I—I—1—1—I—I—I—I—I—I—I—I—I—

r

Discharge

valve open

0 500 1000 1500 2000 2500 3000

Time (ms)

Figure 113. Transient heated agent discharge test; straight piping. The vessel was filled to 1/2

liquid fill volume with HFC-227ea, pressurized to 4.12 MPa, then heated to 70 °C

prior to discharge.
OPTIMIZATION OF SYSTEM DISCHARGE 595

T—I—I—I—f-r—I—I—I—

I—I—I—I I j—r—1—I—I—|—r

i i i

| s i i t ~ Discharge

valve open

0 500 1000 1500 2000 2500 3000

Time (ms)

Figure 114. Transient heated agent discharge test; straight piping. The vessel was filled to 1/2

liquid fill volume with CF3I, pressurized to 4.12 MPa, then heated to 70 °C prior to discharge.
596 8. OPTIMIZATION OF SYSTEM DISCHARGE

do not show characteristic signs of liquid run out. This observation can be explained by the prediction

of single-phase halon 1 301/nitrogen and HFC-125/nitrogen mixtures at the recorded elevated temperatures. A single-phase compressible fluid is discharging in those cases. HFC-227ea and CF3I pipe

pressure traces do show Hquid run out. These liquid discharge times are not significantly less than the

discharge times if the agents had been discharged at 23 °C at the lower initial bottle pressures. One

effect that compensates for the higher discharge pressure is the increase in the initial liquid-phase

volume in the storage vessel and consequently a decrease in the ullage volume.

8.6.6.10 Visual Observations. The transparent test section was installed and high speed movies

taken of a limited number of experiments. CF3I attacked the plastic pipe strongly and would probably

shatter the tube consistently, and therefore, it was the only agent not filmed. Although the test section was hydrostatically tested to failure at 1 1 MPa, during testing a number of tube failures occurred at around 1 MPa. These failures were attributed to temperature induced failure and/or chemical attack on

the plastic tube. All failures occurred at the end of the test after the pressures had equilibrated.

The location of the sight tube was at the exit of the vessel just after the valve reducer section.

The initial vessel conditions were ambient temperature, 4.12 MPa charge pressure, and 1/2 liquid fill volume. In addition, one HFC-227ea test with the same fill conditions had the sight tube located

2.8 m downstream from the valve reducer. One constant-head test of HFC-227ea at 2.5 MPa was

filmed with the sight tube located just downstream from the valve reducer. The most important

observation for all of these tests was that a cloudy two-phase fluid was observed immediately after the

vessel valve was opened, and it persisted until essentially all of the liquid contents of the vessel had

emptied. This implies that bubbles nucleate somewhere before this location, and two-phase flow is always present during the liquid discharge period. Figure 115 shows four frames of the movie taken

of the HFC- 125 discharge. The first picture is of the frame just before the arrival of the agent. The

next picture is of the very next frame 2 ms later. The cloudy section indicates the two-phase flow. The third picture is of the flow during the middle of the two-phase flow time and is indicative of the frames throughout the two-phase flow period. The last picture is of a frame near the end of the twophase flow time. The tube is noticeably clearing up. The time interval between the initial cloudy flow and until it starts to clear up is indicative of the

liquid discharge time and it roughly corresponds to the time from the initial pipe pressure rise to the

end of the hump in the pipe pressure traces. Table 40 gives the time interval from the first cloudy

frame to the point where the flow is clearing up.

8.6.7 Transient Discharge Model. The experimental study yielded a great deal of quantitative

information on the flow characteristics of the alternative agents and halon. The effects of piping

configuration and initial fill conditions were shown to have significant impact on the flow characteristics. Even though a number of tests were performed, a very limited range of conditions and piping

configurations were explored. Modeling the pipe flow would allow other configurations and

conditions to be examined. With this in mind, a computer model was developed that predicts the

transient discharge of an alternative agent superpressurized with nitrogen from a storage bottle through

piping. The model assumptions and equations are presented below.

The model assumes homogeneous, equilibrium two-phase compressible flow. The homogeneous

assumption specifies that the gas and liquid phase are well mixed and traveling at the same velocity.

This is a good assumption for bubbly flow at low void fractions. Bubbles tend to coalesce at higher

void fractions; the transition from bubbly flow to plug flow occurs around a void fraction of 0.3

(Whalley, 1990). At high liquid flow rates though, large bubbles tend to be broken up. The fluid is assumed to be in thermodynamic equilibrium at all locations. Adiabatic, isenthalpic flow is assumed,

greatly simplifying the analysis. The change in entropy is attributed to lost work due to frictional
OPTIMIZATION OF SYSTEM DISCHARGE 597

f= 1200 ms r= 1400 ms

Figure 115. A sequence from a high-speed movie (500 frames/s) of the flow of HFC- 125 in 15.9 mm I.D. piping just after the valve reducer.
598 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 40. Two-phase time interval from high speed movies

Agent Figure number Pressure Cloudy flow interval tj (exp.) (MPa) (± 0.05 s) (s)

Halon 1301 b-18 4.12 1.2 1.0 - 1.2

HFC- 125 B-41 4.12 1.2 0.9 - 1.0

HFC-227ea B-67 4.12 1.1 0.8 - 0.9

HFC-227ea B-68 4.13 0.9 0.8 - 0.9

HFC-227ea B-63 2.5 1.1 0.6 - 0.7

losses. The agent flow in the piping is always assumed to be two-phase flow for initially nitrogensaturated bottle conditions. If the liquid agent is not saturated with nitrogen, the flowing agent is a single-phase liquid until the pipe pressure drops to the equilibrium nitrogen saturation pressure, and at that point degassing occurs. The liquid in the bottle can either follow the isenthalpic pressure/density

path, or be frozen (constant density, single-phase liquid) until a critical pressure is reached. There is no mass transfer across the initial liquid-gas boundary in the bottle; bubbles that form in the liquid

phase are assumed to stay in the liquid phase. The ullage gas expands adiabatically and isentropically,

and is treated as an ideal gas. The ratio of specific heats C^C^ = y is a constant value.

Figure 116 shows schematically the bottle and pipe fluid contents for the modeled conditions.

Pipe filling and emptying, and ullage discharge are calculated differently than the pipe full case.

8.6.7.1 Model Equations. The foundation for the flow prediction is an application of the steadystate mechanical energy balance and continuity equation for pipe flow. An important assumption is that the process is quasi-steady, i.e., the upstream stagnation properties (the bottle conditions) are fixed over a time increment, and a steady mass flow rate is calculated. The small out-flow of mass changes

the bottle conditions for the next time increment. The steady-state differential mechanical energy

balance is

udu + — + 6LW = 0 (38)

P

where bLW, the lost work term for pipe flow is given by

6LW = ^dL (39)

D

The equation of continuity for steady flow is
I

OPTIMIZATION OF SYSTEM DISCHARGE 599

E

before

discharge

pipe

filling

pipe full, single-phase liquid

inside vessel

pipe full, two-phase fluid

inside vessel

pipe

emptying

ullage

discharge

Figure 116. Schematic showing stages of agent discharge. The model calculation scheme is different for the various stages.

!

1
600 8. OPTIMIZATION OF SYSTEM DISCHARGE

pu = G - const. (40)

where G is the mass velocity.

Equation (38) can be re-arranged, substituting for u with the continuity equation, and integrated over the pipe length to yield

States / and 2 are the initial (upstream) and final (exit) states respectively. This equation is applicable

to steady compressible flow in pipes with diameter D. If the pipe length is known, G can be

calculated, though the calculation is not explicit since the pipe exit pressure is not specified and the

flow may be choked. The solution is the pipe exit pressure that maximizes the value of G in Equation (41). Re-arranging Equation (41) so that it is explicit in G, a one-dimensional search is performed to obtain the value of the exit pressure that maximizes G.

If the discharge pressure (the pressure of the receiving vessel, or atmospheric pressure if dis- charged to the open) is above some critical value, the pipe exit pressure is at the discharge pressure for G to be maximized. If the discharge pressure is below some critical value, the flow in the pipe is choked, and the pipe exit pressure is higher than the discharge pressure.

In the above formulation, it is necessary to have the density of the two-phase mixture expressed

explicitly as a function of pressure. Recall that the two-phase flow is assumed to expand adiabatically

at constant enthalpy. The Gibbs phase rule for non-reacting systems is

where F is the degrees of freedom, c is the number of components, and (j) is the number of phases. In a two-phase, two-component mixture, there are two degrees of freedom. Therefore, with enthalpy

fixed, density can be expressed as a function of pressure. An appropriate equation of state is all that is needed to generate the functional relationship between density and pressure. For any upstream condition, the mass velocity and pipe exit pressure can be calculated. This

alone is not enough though for a prediction of the transient bottle discharge. Over each time

increment, some mass exits the bottle and the pipe (after the pipe has filled). The difference between

the mass leaving the bottle and that leaving the pipe is the change in the mass of the pipe contents. A

mass balance must be applied to ensure that the mass exiting the pipe and the change in the mass of

the pipe contents is equal to the mass leaving the bottle. This allows for a time increment to be

calculated. The calculation scheme is described below.

Assuming the pipe is already filled at a given bottle pressure, a mass velocity and pipe exit

pressure are calculated. From the mass velocity a mass flow rate is obtained; this is the mass flow

rate leaving the bottle over a short time period. The mass in the pipe at that time is obtained by using

a variation of Equation (41) and solving for pressure as a function of distance along the pipe. Since

density is a known function of pressure, the density along the pipe is known. Therefore, the mass in the pipe is given by

(41)

F = c-(})+2 (42)
8. OPTIMIZATION OF SYSTEM DISCHARGE 601

= ^D'i:pa.,(^)^L (43)

1=0 4

^ ' avg

where Mp is the mass in the pipe, the sum of AL is the total length L, and p^vgix) is the mean density

at the pipe location (x). The mass lost from the pipe is given by the following "liquid" mass balance,

(a mass balance applied to the original liquid phase contents)

Mlp = {Mvl^^Mp) - (Mvl^.^,+Mp^^^) (44)

where Mlp is the mass lost from the pipe over the time increment At. Mvl is the liquid mass in the

vessel. The liquid mass is obtained from the liquid volume multiplied by the liquid density at the

vessel pressure. The liquid volume is the difference between the total volume and the ullage volume.

The ullage expands isentropically and is assumed to behave as an ideal gas so the following relationship holds

V P(—y = Const. (45)

where P is the bottle pressure, is the ullage volume and is the mass of the ullage contents. The time increment (At) over which the pressure in the bottle drops from P^ to is calculated

from

At = ^

(46)

where and W,^^ are the mass flow rates at times (t) and (t+At) respectively. The bottle pressure is then stepped down and the process repeated. A pressure increment of 5 kPa was found to be

sufficient for calculation accuracy. The filling and emptying of the pipe with the two-phase mixture is approximated using a different

method. In each case, the effective pipe length is segmented, and the calculation proceeds by the

filling or emptying of the segments. Eight segments provided sufficient accuracy for the calculations

described here. In the filling case, the addition of each segment represents an increase in the pipe length for the

calculation of pipe end pressure and mass velocity. The flow is assumed to be two-phase and choked

at the "liquid front" which is located at the end of the particular segment of interest. For the first time

increment, the pipe length is equal to the length of a segment. At this point, the bottle pressure, pipe

end pressure, mass velocity and mass in the pipe are all unknown. The solution is obtained by

guessing a bottle pressure, solving for the remaining unknowns and then comparing the "liquid" mass

balance equation to that solution. The "liquid" mass balance reduces to
602 8. OPTIMIZATION OF SYSTEM DISCHARGE

(47)

during the pipe filling stage. MvIq is the initial liquid-phase mass. Once the exit pressure and mass

velocity are obtained for the guessed bottle pressure, the mass in the pipe can be calculated. The

value of A/v/j is used to calculate a new liquid volume in the bottle, and an ullage volume is obtained

by difference. This ullage volume is compared to the ullage volume calculated from the isentropic

expansion relationship using the initial bottle pressure guess. Successive guesses of bottle pressure are required until the ullage volumes converge. The time to fill the first section of pipe is estimated as the mass in the pipe divided by the mass flow rate. The time to fill the pipe up to the yth segment is estimated by

When the bottle is emptied of its originally liquid contents, a calculation similar to the one above

is performed. To start, the first segment is empty of the two-phase mixture, and the bottle pressure is calculated from the isentropic expansion relationship, with the ullage volume equal to the bottle volume plus the empty pipe segment volume. Isentropic expansion of the ullage gas into the pipe is justified due to the relatively slow velocity of the two-phase mixture/gas front which acts as a slowmoving piston allowing the gas to expand essentially reversibly. The pipe length is equal to the

effective length minus the segment. The pipe exit pressure is solved using Equation (41) and

maximizing the mass velocity. The mass in the pipe is then calculated. The time increment for the

liquid two-phase/gas front to travel to the end of the jth segment is estimated by : , ,

,

^ Mp. - Af/7. (48)

'fill, j

t.empty, j (49)

Emptying of the ullage contents (assumed to behave as an ideal gas) is calculated in a similar

quasi-steady fashion. Assuming adiabatic flow of an ideal gas, the following relation holds

(Balzhiser et al., 1972) .

'
8. OPTIMIZATION OF SYSTEM DISCHARGE 603

After the two-phase slug leaves the pipe, the gas in the pipe accelerates to its maximum mass velocity

given by the vessel pressure, assuming no ullage mass has left the piping. This creates a discontinu- ous step down to the exit pressure for the gas flow as a result of the approximation that the gas

isentropically expands into the piping while the two-phase mixture is leaving the piping. In reality,

there is a smooth, steep decrease in the exit pressure as the two-phase mixture leaves and the ullage

begins to discharge.

The calculation proceeds in the following manner. The mass flow rate of the gas exiting the

piping is multiplied by a small time increment. This mass is subtracted from the mass balance for the next time step. Successive guesses of the vessel pressure are made until the guessed pressure and the

calculated pressure from isentropic expansion of the vessel contents converge. A new mass flow rate

is obtained, and the process is repeated.

The pressure recovery in the bottle is thought to be due to nitrogen coming out of solution from

the supersaturated liquid. In these tests there was no direct visual confirmation (i.e., visualization of

the formation of bubbles in the vessel liquid) proving that to be the case, but it is assumed to be the

correct interpretation. For some agents and fill conditions the pressure recovery is more pronounced,

whereas in others it is not evident. Pipe flow visualization with high speed movies does confirm the

two-phase nature of the fluid exiting the valve. To simulate the retardation of the nitrogen degassing

in the bottle, the calculation initially proceeds with the bottle liquid phase density fixed at the initial density. When the bottle pressure drops below a critical degassing pressure, nitrogen comes out of

solution. The critical degassing pressure used is an experimentally inferred value for a particular agent

and fill condition, and it is the lowest pressure achieved before the observed pressure rise. There is a

lot of scatter in these data for a given agent; a mean value was selected. For HFC-227ea and CF3I the

calculation assumes no degassing in the bottle, since no characteristic pressure recovery was noticed

for these agents (except one HFC-227ea and CF3I test at 2/3 fill volume). When the bottle pressure

drops below the degassing pressure, the liquid density decreases due to bubble formation. The liquid

(two-phase) mass is fixed, thus its volume must increase. The increase in the liquid (two-phase)

volume compresses the ullage gas, and the pressure rises. The mass velocity at the pipe exit does not

instantaneously increase to a value given by an elevated bottle pressure. There must be a pressure wave moving down the length of the pipe (similar to the pipe filling case), since after the bottle pressure rises, more mass can be stored in the pipe. This transient is approximated in the following manner. Over the transient time Ar^^^^^, the mass lost from the pipe is

Time (r) is at the beginning of the transient. The transient time is equal to the sum of the mass lost from the pipe and the change in the pipe mass divided by an average mass flow rate as shown below.

The average mass flow rate is the mean value of the mass flow rate after the transient and the mass

flow rate before the transient.

Mlp = PF, Af,

trans (51)

(52)

trans iW,^^t-W)|2

or substituting for Mlp
604 8. OPTIMIZATION OF SYSTEM DISCHARGE

trans ^Pt^At-^Pt

(53)

To calculate Af,^^^^, a bottle pressure at the end of the transient is guessed, W^^^, and Mp^^^^ are calculated, then Ar,^^^^ is calculated from above. The bottle liquid content mass at the end of the

transient is

From this bottle two-phase mass and the two-phase density based on the guessed pressure, the liquid

volume and then the ullage volume is calculated. Bottle pressure is calculated from the isentropic

relationship of the ullage gas, compared to the initial guessed pressure and, subsequently, new guesses

are made until the two pressures converge.

8.6.7.2 FiH Conditions and Thermodynamic Inputs. Specific fill conditions, i.e., mass in the

liquid and gas phase and fill volume are calculated by the program PROFISSY. The thermodynamic

relationships are obtained from a variation of the program PROFISSY that performs isentropic or

isenthalpic expansion calculations. While the process for obtaining the thermodynamic inputs is not automatic, the inputs are valid for any fill condition at a given charge pressure and temperature. First,

PROFISSY is used to calculate the mass of agent and nitrogen in the liquid and gas phases for a given

fill condition and temperature. An isenthalpic expansion of the liquid phase and an isentropic

expansion of the gas phase calculation is performed. For the isenthalpic expansion, at the initial storage pressure a single phase exists. During the expansion a two-phase gas/liquid mixture is formed.

The average density, temperature, and mass fraction in the liquid phase are obtained at each pressure

step. These results are fitted to third or fourth order polynomial functions of pressure. Appendix C

shows some plotted results. The coefficients for the pressure/density for each of the agents at various

pressures are given in Table C-1. The isentropic expansion of the gas phase begins with initially

gaseous agent and nitrogen; upon expanding, some agent condenses. The mean density is plotted as a

function of pressure, then fitted by a power law equation. The inverse of the exponent is defined as the ideal y coefficient. Results are shown in Appendix C for agent/nitrogen mixtures initially at 4.12 MPa. The power-law exponent was insensitive to initial pressure, so y was fixed for each

agent/nitrogen mixture.

8.6.7.3 Friction Factor. Since the flow is described as a homogeneous two-phase flow, the

friction factor analysis will follow the method for established for single phase flows. The Colebrook

equation (see Perry et al., 1984), shown below, relates the Fanning friction factor to pipe diameter,

roughness, and Reynolds number (Re) for turbulent flows as

Mlp (54)

1 e 1.256

4 log

10 + (55)

if \ 3.7i> Re{f
8. OPTIMIZATION OF SYSTEM DISCHARGE

which requires an iterative scheme to solve for /. The Reynolds number is defined as

605

Re uDp

(56)

At high Re, Equation (55) reduces to von Karman's equation (see Perry et al., 1984) — = -41og.o(-^) ^ 3.7D

(57)

Haaland (see Welty et al., 1984) has shown that for 4 x

10"^ < < 10^ and e/D < .05 the friction factor can be expressed (within ± 1 .5 %) as

— = -3.6 1ogio

if

10

Re 3.7D

(58)

which allows explicit calculation of/. Moody (see Perry et al., 1984) reports the value of surface

roughness for drawn tubing as 0.00152 mm, and it is the value used here. For the flows encountered in this study, the density and viscosity of the mixture are functions of

pressure and temperature. Therefore the Reynolds number, and thus the friction factor, is a function of

pipe position, and in addition, a function of time for the transient discharge. In the model, the friction factor is assumed to be constant. The validity of this assumption is demonstrated by the following

analysis.

The NIST program REFPROP (Gallagher et al., 1993) was used to obtain the viscosity of the

liquid and gas phase for halon 1301, HFC- 125, and HFC-227ea as a function of temperature; nitrogen

viscosity as a function of temperature was also obtained. The viscosity of the homogeneous two-phase

mixture was estimated from the functional form suggested by Dukler et al. (1964).

yPh (l-y)P;. .CON

where y is the mass fraction of the gas phase, and the subscripts /, g and h stand for liquid, gas, and

homogeneous respectively. The gas phase mixture viscosity at a given temperature was estimated as

the sum of the agent and nitrogen gas viscosities multiplied by their respective volume fractions in the

gas phase. Table C-2 in Appendix C gives the calculated results, expressed as polynomial fits, at

initial pressures of 2.75 MPa and 4.12 MPa at 20 °C. The first step in the analysis was to calculate a
606 8. OPTIMIZATION OF SYSTEM DISCHARGE

friction factor from the von Karman equation, then simulate the flow. From the calculations, the

velocity, density, and viscosity are known for any location and time in the pipe, thus local Reynolds

numbers are calculated, and Equation (58) is used to obtain the local friction factor. The mean friction factor at a specified time is given by

f - - T (60)

where is the mean friction factor, and x is a particular pipe location. Plots of the Reynolds number

and mean friction factor versus time for a 4.1 m straight pipe with two different internal piping

diameters (15.9 mm and 9.5 mm) are shown in Figures 117 to 122. In these figures, the thick dashed

lines are the high Reynolds number limits for the friction factors given by the von Karman equation.

The mean Reynolds number is shown along with two curves representing ± 2a. The standard

deviation of the mean friction factor is on the order of 10~^ and is not shown on the plots. Although

the Reynolds number decreases over the flow time, the mean friction factor rises only slightly. The mean friction factor and the friction factor calculated from the von Karman equation differ by 3 % for

the larger pipe and 6 % for the small pipe. The main conclusion is that for the conditions here, the

Reynolds number is high enough that for all practical purposes viscous effects can be ignored,

allowing the von Karman equation to be used. Pressure drops across bends, valves, and other piping elements can be included by treating those

items in terms of an effective pipe length. This technique is well established in hydraulic calculations.

Here the flow elements of interest are the discharge valve and the reducer, 90° angle bends, expansions, contractions, and tees. The effective lengths for the discharge valve and valve reducer section were determined from the experimental pressure drops. The effective length was varied until the

calculated pressure drop from the bottle pressure to the first pressure tap reproduced the experimental

pressure drop. For the large diameter piping, an effective length equivalent to the open volume of the

valve and the reducer (1.25 x 10~^ m'^ ± 2.5 cm^) divided by the pipe cross-sectional area gives

0.63 m and approximates the experimental pressure drop. For the small pipe, the pressure drop across

the valve and reducer was very small (most of the pressure drop occurs in the piping), and the

effective length was estimated to be 0.1 m for the constant head tests. For the transient tests the

pressure drop across the valve and reducer was ignored, and the open volume of the valve and reducer was added to the vessel volume. When the valve opens, it is assumed the fluid fills that volume and

the ullage isentropically expands. The effective bottle volume becomes the bottle volume plus the

valve and reducer volume.

The effective length for smooth bends in the piping was estimated from the following equation

(SFPE, 1988)

^= [0.13.1.85(-^)^^] (-^)0-^ (61)

D R 180°

where a° is the bend angle, R is the bend radius, and r is the pipe inner radius. With the friction factor given by the von Karman equation, the equivalent length for the small pipe with a 50 mm

radius bend was 0.25 m and for the large pipe with a 150 mm radius bend was 0.47 m.
OPTIMIZATION OF SYSTEM DISCHARGE 607

0.004

0.0035

o

CO LLc

o

T—I—I—I—-1—p—I—T—I—I—f

I I

I

I I—

r

-e—e—e—e—e—e—e—e e e— —e— Friction Factor

-High Re Friction Factor —B— Mean Re

^ I I I I I I I I I I I I I I I L i_ I I I J 0.0025 1 0 0.2 0.4 0.6 0.8

Time (s)

4

o

Q> JQ

E

0) ]D

O

c

, >»o

Figure 117. Mean friction factor and Reynolds number as a function of time for a simulated

transient discharge of halon 1301: 15.9 mm I.D. piping was specified, and the vessel was at 1/2 liquid fill volume and pressurized to 4.12 MPa.
608 8. OPTIMIZATION OF SYSTEM DISCHARGE

0.004 T—I—

r

T—I—I—

I—I—I—I—I—I—mr T—

r

0.0035

o

(0

o

O

u. 0.003

-e— Friction Factor — - High Re Friction Factor

-B— iviean Re

-e—e—e e e e e e e e-=

- -Q B B B a B- -B B B B—

:

0) n

E

z

(0o

c

>»

a>

0.0025

J I I I I I I I L J I I I I L

1 0.5 1 1.5 2 2.5

Time (s)

Figure 118. Mean friction factor and Reynolds number as a function of time for a simulated

transient discharge of halon 1301: 9.5 mm I.D. piping was specified, and the vessel was at 1/2 liquid fill volume and pressurized to 4.12 MPa.
OPTIMIZATION OF SYSTEM DISCHARGE 609

0.004

_ 0.0035

o

o

(0c

mmm o

u.

0.0025

T—I—1—I—I—I—I—

T

T 1 T

-e—e—e—e—e—e—e—e—e e

0.003 P- —e— Friction Factor — — - High Re Friction Factor —B— Mean Re

J I I I I I I I I I I I I I I I J ' '

0.2 0.4 0.6 0.8

Time (s)

Figure 119. Mean friction factor and Reynolds number as a function of time for a simulated

transient discharge of HFC-125: 15.9 mm I.D. piping was specified, and the vessel was at 1/2 liquid fill volume and pressurized to 4.12 MPa.
610 8. OPTIMIZATION OF SYSTEM DISCHARGE

0.004 1—I—

r

"I I I

I—I—I—I—I—

I—I—I—I—I—

I—I—I—

r

0.0035

o

(0

o

—e— Friction Factor

-High Re Friction Factor —B— Mean Re

-e—e—e—e e e e e e e0.003

0.0025

-I—I—I—I—I—I—I I I I I I I ' I J \ L

0.5 1 1.5

Time (s)

2.5

Figure 120. Mean friction factor and Reynolds number as a function of time for a simulated

transient discharge of HFC- 125: 9.5 mm I.D. piping was specified, and the vessel was at 1/2 liquid fill volume and pressurized to 4.12 MPa.
OPTIMIZATION OF SYSTEM DISCHARGE 611

0.004 T—I—I—I—

r

"1—I—I—I—I—I—I—I—

r

-e— Friction Factor

0.0035

-High Re Friction Factor —B— iVIean Re

-e—e—e—e—e e e e e e- u

(0c

LL 0.003

-B—G B B B &- -e B B &-

0.0025

J I I i_i I I I I I J J I L

0.5 1 1.5

Time (s)

Figure 121. Mean friction factor and Reynolds number as a function of time for a simulated

transient discharge of HFC-227ea: 15.9 mm I.D. piping was specified, and the vessel was at 1/2 liquid fill volume and pressurized to 4.12 MPa.
612 8. OPTIMIZATION OF SYSTEM DISCHARGE

0.0035 1

' ' ' I ' ' ' I ' '

' T—I—I—I—I—

r

o o Q Q o Q—e—e—e—e0.003

o

(0 LLC

o

0.0025

0.002

-e— Friction Factor — -High Re Friction Factor

-B— Mean Re

J I _l I I I

I '

I

' I I L I L

0 0.2 0.4 0.6 0.8

Time (s)

Figure 122. Mean friction factor and Reynolds number as a function of time for a simulated

transient discharge of HFC-227ea: 9.5 mm I.D. piping was specified, and the vessel was at 1/2 liquid fill volume and pressurized to 4.12 MPa.
8. OPTIMIZATION OF SYSTEM DISCHARGE 613

For the contraction run, the pressure drop in the first 1.00 m of large piping was very small,

therefore, the pipe volume was added to the valve and reducer volume which increased the effective

bottle volume. For the expansion run, the flow was assumed to be choked at the expansion, therefore,

the downstream pressure does not affect the mass flow rate. The calculations assume a 1 .00 m length

of small piping.

The effective length for the tee configuration was the sum of the effective length of the valve and

reducer (0.63 m) plus the length of piping to the tee (2.1 m), plus the effective length of the tee element (0.9 m), plus the length of one leg to the recovery tanks (1.2 m).

The input requirements for a piping element are the effective length and the location of the piping

element. The location of the piping element is important because the mass in the pipe is based on the sum of the mass in a number of small portions of the physical pipe. At a piping element with an

effective pipe length, a step change in the density is required for the pipe mass calculation.

8.6.8 Model Calculations

8.6.8.1 Constant-Head Conditions. The constant-head tests were simulated by fixing the

upstream conditions and calculating the steady mass flow rate and the pipe pressures at the tap

locations. The effective length of the 15.9 mm diameter pipe was specified as 4.13 m, 3.50 meters of

pipe from the valve reducer to the recovery tanks plus a 0.63 m effective length for the valve and

valve reducer section (which essentially reproduces the experimental pressure drop from the bottle to the first pressure tap location). For the 9.5 mm pipe an effective length of 3.10 m was specified.

There are 3.00 m of physical pipe length until the expansion to the larger pipe diameter which leads to the recovery tanks, and 0.10 m was chosen as an effective pipe length for the valve and reducer.

Table 35 compares the calculated steady mass flow rate with the experimentally determined

average mass flow rate. Figure 123 is a plot of experimentally determined flow rates versus the

calculated flow rates for all of the tests. The calculated results for the small diameter tests are within

10 % of the mean experimental value. For the large diameter tests, the experimental values are larger

than the calculated values. Given the uncertainty in the experimental values and the probability that the calculated steady values should underestimate the large diameter values, the results compare

favorably.

Calculated pipe pressures are shown in Figures 124-131. The pressure predictions show good

agreement with the experimental pressures and illustrate the validity of the flow assumptions.

Deviations from the experimental values may be due to metastable flow, heat transfer, and/or

departures from the assumed homogeneous flow condition.

8.6.8.2 Transient Conditions. Each transient test was simulated by the model. Experimental

fill conditions were used. It was assumed that the mass of dissolved nitrogen was the equilibrium

amount. Calculations were performed for the cases with the bottle liquid contents in thermodynamic

equilibrium and repeated with a specified degassing pressure for each agent at the given storage

pressure. Figures 132 and 133 are the simulated pressure traces for one test (halon 1301 at 1/2 liquid fill volume and 4.12 MPa storage pressure flowing through 3.5 m of straight piping). The results from

Figure 132 are for equilibrium bottle conditions, while the results for Figure 133 are for "frozen" nonequilibrium bottle conditions until the critical degassing pressure (in this case 2.5 MPa) is reached.

Figures 134 and 135 compare selected calculated pressure traces to experimental pressure traces. The
614 8. OPTIMIZATION OF SYSTEM DISCHARGE

^2

0)

o

(0

(0

(0

(0

c

o

e

Q.

X

HI

6

5 :

£ 4

I I I I I I I I i I

I I I I I

I I 1 I II I I I I I I I I

^ .4

• Halon 1301, 9.5 mm

O Halon 1301 15.9 mm

HFC-125, 9.5 mm

HFC-125, 15.9 mm

HFC-227ea, 9.5 mm

O HFC-227ea, 15.9 mm

CF3I, 9.5 mm

A CF3I, 15.9 mm

0 1 2 3 4 5 6 7

Calculated Steady Mass Flow Rate (kg/s)

Figure 123. Constant-head mass flow rate. The dashed lines represent ± 10 % deviation of

predictions. Estimated measurement errors are indicated by error bars.
8. OPTIMIZATION OF SYSTEM DISCHARGE 615

3.5

G

^

2.5 fi - - Q.

O

V)

o

1.5

0.5

I I I I II I I I II I I I II I I I

I

I I I I

I I I I II

I I I I

I TTTT

I

Vessel pressure

- -0 - - 3.0 MPa

----2.5 MPa

-A- -2.0 MPa

"A

I

J--L I I I I I I I I

' ' '

I

'

I I

' ' I '

0 0.5 1 1.5 2 2.5 3 3.5 4

Distance (m)

Figure 124. Pipe pressures for constant-head tests: halon 1301 and 3.1 m effective length of

9.5 mm I.D. piping. Symbols: calculations, and vertical lines with error bars: mean

experimental pressures ± 2o uncertainty.
616 8. OPTIMIZATION OF SYSTEM DISCHARGE

3.5

<? 2.5

£ 2

3

(0

(0

0)

1.5

0.5

Vessel pressure .

- -0 - - 3.0 MPa

- - - - 2.5 MPa

- - A - - 2.0 MPa

V o

A

1

2 3

Distance (m)

Figure 125. Pipe pressures for constant-head tests: halon 1301 and 4.13 m effective length of

15.9 mm I.D. piping. Symbols: calculations, and vertical lines with error bars: mean

experimental pressures ± 2a uncertainty.
8. OPTIMIZATION OF SYSTEM DISCHARGE 617

3.5

-I

2 2-5

o 2

3

(0

o

Q- 1.5

0.5

I I I I I

I I I I

I

I I I I I

I I I I

I

I I I I II I I I

I

I I I I I I I I I

I

I

Vessel pressure

-0- 3.0 MPa

- - - 2.5 MPa

- - A - 2.0 MPa

1

1

I I

' '

I ' '

I

' ' I I'

I

'

0 0.5 1 1.5 2 2.5 3 3.5 4

Distance (m)

Figure 126. Pipe pressures for constant-head tests: HFC- 125 and 3.1 m effective length of

9.5 mm I.D. piping. Symbols: calculations, and vertical lines with error bars: mean

experimental pressures ± 2a uncertainty.
618 8. OPTIMIZATION OF SYSTEM DISCHARGE

3.5 T T

<o 2.5

CL

A - - 3

0)

G

I

•A

1.5

0.5

I

t

t

1

Vessel pressure -

- - 0 - - 3.0 MPa

-B - - 2.5 MPa

- A -2.0 MPa

A .

i

O

J I L

2 3

Distance (m)

Figure 127. Pipe pressures for constant-head tests: HFC-125 and 4.13 m effective length of

15.9 mm I.D. piping. Symbols: calculations, and vertical lines with error bars: mean

experimental pressures ± 2a uncertainty.
OPTIMIZATION OF SYSTEM DISCHARGE 619

3.5

i3X

<^ 2 5

o 2

3

(0

1.5

0.5

I I I I I

I I I I I

I I ! I I I I < I II

I

I

A.

I

Vessel pressure

- -0 - - 3.0 MPa

- -B- - 2.5 MPa

-A - 2.0 MPa

I

S

N

1 I I I

I ' 'I '

I '

I

'

I ' 'I

' I' '

0 0.5 1 1.5 2 2.5 3 3.5 4

Distance (m)

Figure 128. Pipe pressures for constant-head tests: HFC-227ea and 3.1 m effective length of

9.5 mm I.D. piping. Symbols: calculations, and vertical lines with error bars: mean

experimental pressures ± 2c uncertainty.
620 8. OPTIMIZATION OF SYSTEM DISCHARGE

2.5

[3 ^

(0 a.

0)

1.5

G

•Q

I

A

I

1

I

1

Vessel pressure

- -0 - - 3.0 MPa

----2.5 MPa

- - A - - 2.0 MPa

Q

4^

0.5

J I J L_JL j_l i_L

2 3

Distance (m)

4

Figure 129. Pipe pressures for constant-head tests: HFC-227ea and 4.13 m effective length of

15.9 mm I.D. piping. Symbols: calculations, and vertical lines with error bars: mean

experimental pressures ± 2a uncertainty.
8. OPTIMIZATION OF SYSTEM DISCHARGE 621

3.5

TO 2.5

^

O

3

(0

(0

o

1.5

0.5

I I I

I I I I I

I I I I II

'

"T

I I I I I I I I

I I I I

Vessel pressure

- -e- - 3.0 MPa

- -S- -2.5 MPa

- - A - - 2.0 MPa

•Q

I I '

I I I I I

0.5 1 1.5 2 2.5 3 3.5 4

Distance (m)

Figure 130. Pipe pressures for constant-head tests: CF3I and 3.1 m effective length of 9.5 mm

I.D. piping. Symbols: calculations, and vertical Unes with error bars: mean experimental pressures ± 2a uncertainty.
622 8. OPTIMIZATION OF SYSTEM DISCHARGE

2.5

(0

Q.

o

3

0) 1.5

0.5

G

I

I

A

I

I

A

I

"I—'—I

»

I

* Vessel pressure

--0--3.O MPa

- - - 2.5 MPa

- A - - 2.0 MPa

Q

V

\ :a

I I I I \ I I I L

2 3

Distance (m)

Figure 131. Pipe pressures for constant-head tests: CF3I and 4.13 m effective length of 15.9 mm

I.D. piping. Symbols: calculations, and vertical lines with error bars; mean experimental pressures ± 2o uncertainty.
OPTIMIZATION OF SYSTEM DISCHARGE 623

0 500 1000 1500 2000

Time (s)

Figure 132. Simulated pressure traces for transient halon 1301 discharge: 1/2 liquid fill volume,

pressurized to 4.12 MPa, and 3.5 m of 15.9 mm I.D. piping. Vessel liquid contents

are assumed to be in equilibrium at all times.
624 8. OPTIMIZATION OF SYSTEM DISCHARGE

0 500 1000 1500 2000

Time (s)

Figure 133. Simulated pressure traces for transient halon 1301 discharge: 1/2 liquid fill volume,

pressurized to 4.12 MPa, and 3.5 m of 15.9 mm I.D. piping. Vessel liquid contents

are in a frozen, non-equilibrium state until degassing occurs.
8. OPTIMIZATION OF SYSTEM DISCHARGE 625

T r—T 11 1 1 1 11

1 1 1 1——

j

1 r—I r

0 500 1000 1500 2000

Time (s)

Figure 134. Selected simulated pressure traces from Figure 132 compared to the experimental

data presented in Figure B-19.
626 8. OPTIMIZATION OF SYSTEM DISCHARGE

Time (s)

Figure 135. Selected simulated pressure traces from Figure 133 compared to the experimental

data presented in Figure B-19.
8. OPTIMIZATION OF SYSTEM DISCHARGE 627

non-equilibrium bottle condition calculations yield vessel and pipe pressure results that compare more

favorably to experiments than do the equilibrium bottle condition calculations.

The model describes flow conditions which may be of interest. Figure 136 is a plot of the mean

density of the fluid at the pipe exit (or at the liquid two-phase front before the pipe is filled). Figure 137 is the mass flow rate. Figure 138 is the exit velocity for the two-phase mixture.

Figure 139 is the exit temperature of the two-phase mixture. Figure 140 is the exit quality of the twophase mixture. Figure 141 is the mass fraction discharged from the pipe.

The results for the calculated and the experimental liquid discharge times are presented in Tables 36-38. Figures 142-145 show the experimental liquid discharge times versus the calculated

discharge times. In these figures, the dashed lines represent ± 10 % deviation from the experimental

discharge times, and the error bars indicate 2a uncertainty in the experimental values. The liquid flow

time calculations based on equilibrium bottle conditions tend to be shorter than the calculations based

on a fixed degassing pressure or no bottle degassing. For HFC-227ea and CF3I, there appears to be a systematic trend for the cases with the small diameter piping. For most cases, the calculated liquid

discharge time is less than 90 % of the experimental discharge time. These deviations may be due to the recovery tanks pressure build-up during those tests which decreases the flow rate below the choked

condition for a period of time. Another systematic trend appears to be an overprediction in the liquid

discharge time for the 2/3 fill, small piping cases. One possible explanation may be that significant

heat transfer to the ullage contents may raise the pressure above the isentropic expansion condition.

Since the ullage can expand by a factor of 3 before the liquid runs out, the adiabatic temperature of

the ullage will be lower than the 1/2 fill case, which favors heat transfer from the vessel walls. The

two-phase flow time is the longest for these tests which also favors heat transfer. The cold temperature discharge simulation of halon 1301 significantly underpredicts the liquid discharge time, which

suggests that heat transfer from the pipe walls may be increasing the fluid expansion which slows the

flow down.

There is no obvious reason why the model cannot be extended to higher bottle pressure and larger

pipe diameter predictions. Testing at higher bottle pressures and larger pipe diameters would confirm

the model's ability to predict liquid discharge times under those conditions and add confidence to extending the predictive capabilities.

8.6.9 Hardware Design Scheme. Since the model predicts the experimental discharge times within

± 10 %, it is assumed that the model can be used to calculate the liquid discharge times at higher

pressures and larger pipe diameters, though it should be validated for those cases. A design exercise

for hardware and fill specifications for each of the three alternative agents is detailed below. The

exercise illustrates the model's potential usefulness in selecting hardware for a replacement agent.

Assume that a given halon fire suppression system is designed to protect an engine nacelle and

meets the following specifications: 3.37 kg of halon 1301 superpressurized with nitrogen to 4.12 MPa

at 23 °C in a 0.0042 m^ bottle, and 4 m of 0.016 m diameter piping from the bottle to the injection

port. The calculated liquid discharge time is 1.0 s which meets the duration of discharge criterion in MIL-E-22285 section 3.9 which states "The period of time required to discharge the calculated amount

of agent shall be one second or less, measured from the time the agent starts to leave the tubing ends

until the required amount of agent has been discharged." The "calculated amount" means the mass

stored in the liquid phase which is 90 % of the total agent mass in this case; the discharge time of the

agent vapor in the ullage would take another second for this configuration. For a given alternative

agent, the mass of agent, bottle size, storage pressure, fill condition, and pipe diameter that fulfills the

discharge requirement of a one second liquid discharge time needs to be determined.

Realizing that a number of conditions may meet the discharge requirement, an "optimized" system

requires knowledge of the economics of specific configurations and the space limitations for a given
628 8. OPTIMIZATION OF SYSTEM DISCHARGE

X

UJ

o

Q.

il

o

c

o

o

^ </) CO

Q. U)

I o

H

•«-• (0

(0

c

o

a

c

(0

o

1000 I I 1 1

1 T 1 1

1

I

800

600

400

200

-e— Eq. bottle conditions

-B— Non-Eq. bottle conditions

Pipe full

500 1000 1500 2000

Time (ms)

Figure 136. Calculated mean density of the two-phase fluid at the pipe exit (or at the two-phase

front if it has not reached the pipe exit): halon 1301, 1/2 liquid fill volume, pressurized to 4.12 MPa, and 3.5 m of straight 15.9 mm I.D. piping.
OPTIMIZATION OF SYSTEM DISCHARGE 629

Figure 137. Calculated mass flow rate of the two-phase fluid: halon 1301, 1/2 liquid fill volume,

pressurized to 4.12 MPa, and 3.5 m of straight 15.9 mm I.D. piping.
630 8. OPTIMIZATION OF SYSTEM DISCHARGE

100 T—T r I I 1 1 1 r

X

UJ

O

Q.

K

o

c

o

80

60

m E

Q.

6 40

(0

^ 20

o

o

o

>

Pipe full

—e

—e

—e—

J 1 L

-e— Eq. bottle conditions

-B— Non -Eq. bottle conditions

J L t . J L

200 400 600 800

Time (ms)

1000

Figure 138. Calculated exit velocity of the two-phase fluid: halon 1301, 1/2 liquid fill volume,

pressurized to 4.12 MPa, and 3.5 m of straight 15.9 mm I.D. piping.
8. OPTIMIZATION OF SYSTEM DISCHARGE 631

*- 300 -! r X

UJ

o

Q.

o

c

o

280

260

o

240

B 220

(0

o

3*-< (0

k.o

a.E

o

I- 200 J L

T r I I ^ r

T 1—

r

Pipe full

-e— Eq. bottle conditions

-B— Non-Eq. bottle conditions

J L J I I I I L. 200 400 600 800

Time (ms)

000

Figure 139. Calculated exit temperature of the two-phase fluid: halon 1301, 1/2 liquid fill volume, pressurized to 4.12 MPa, and 3.5 m of straight 15.9 mm I.D. piping.
632 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure 140. Calculated exit flow quality of the two-phase fluid: halon 1301, 1/2 liquid fill volume, pressurized to 4.12 MPa, and 3.5 m of straight 15.9 mm I.D. piping.
OPTIMIZATION OF SYSTEM DISCHARGE

0 500 1000 1500 2000

Time (ms)

Figure 141. Calculated mass fraction discharged from the vessel: halon 1301, 1/2 liquid fill volume, pressurized to 4.12 MPa, and 3.5 m of straight 15.9 mm I.D. piping.
634 8. OPTIMIZATION OF SYSTEM DISCHARGE

^

5

o

E

o

u>

COu

(0

3

or

(0

c

o

E

Q.

X

UJ

o 7

I

' ' '

I

'

' '

I ' '

'

I

' ' yi g

Eq. bottle conditions /

Non-Eq. bottle conditions / ^ / /

/ /

/ 4> i

/

If $4

/

/

J I I I I I I I I I I I I I I I L-l ! I I I L

0 1 2 3 4 5

Calculated Liquid Discharge Time (s)

Figure 142. Experimental and calculated liquid discharge times for halon 1301. Calculations

assuming equilibrium and non-equilibrium bottle conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 635

^

5

o

E

o 4

k.TOO

Q 3

3

(0 *^c

o

E

o

Q.

X

LU

T—T—

r

T T—

r

T—I—I—

r

T—I—I—

r

O Eq. bottle conditions

Non-Eq. bottle conditions

/

/

/

/

/

/

/

/

/

o []

^^^^ s

J__J I I I I I I I I 1—I—I—I—I—

L

I I I I

0 1 2 3 4 5

Calculated Liquid Discharge Time (s)

Figure 143. Experimental and calculated liquid discharge times for HFC-125. Calculations

assuming equilibrium and non-equilibrium bottle conditions.
636 8. OPTIMIZATION OF SYSTEM DISCHARGE

^

5

T—

r

' ' I

' ' ' ' I' ' ' ' I' " ' ' I

' ' ' '/I

O Eq. bottle conditions

Non-Eq= bottle conditions /

J L

/

/

/

/

/

/

/

J I L 1 1 1 I L J I L 1

0 1 2 3 4 5

Calculated Liquid Discharge Time (s)

Figure 144. Experimental and calculated liquid discharge times for HFC-227ea. Calculations

assuming equilibrium and non-equilibrium bottle conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 637

^

5

o

E

o 4

(0

o

(0

2

-

(0

*-< c

o

£

Q.

X

LU

T—

r

O

T—I—1—T—[—I—I—I—I—

I—I—r—I—I—I—I—I—I

r-p\—I—

r

Eq. bottle conditions

Non-Eq. bottle conditions

!/

$5

$5 / y

/ /

/

/ /

/

J L 1 J I I L—J I I I I 1 L

0 1 2 3 4 5

Calculated Liquid Discharge Time (s)

Figure 145. Experimental and calculated liquid discharge times for CF3I. Calculations assuming

equilibrium and non-equilibrium bottle conditions.
638 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 41. Agent mass requirements for a retro-fit system

Agent Mcsn volume factor I inuid (iensitv at 25 °C

(kg/m^)

PniiivJiIfnf mncc ct/^rfS'H

in liquid phase

(kg)

Halon 1301 1 1538 3.02

HFC- 125 3.0 1190 7.02

HFC-227ea 2.5 1395 6.86

CF3I 0.8 2016 3.31

bottle location. Such a design optimization is well beyond the scope of this study, but for the sake of

this exercise, the following considerations are probably close to an actual optimal design scheme.

First, assume that the preferred case would be if the original bottle size and pressure could be used

and that the piping remains the same for an alternative agent. This would represent a "drop-in"

replacement for halon 1301. The next best case would be if the only change was an increase in the

storage pressure. This implies that the bottle can withstand the elevated pressure. If the discharge

time cannot be achieved with the original bottle size, a larger bottle size and increased storage pressure

then is checked. Finally, the piping can be replaced with larger diameter piping to achieve

the desired discharge time.

The pressure range searched was from 4.12 MPa to 8.24 MPa in increments of 1.37 MPa. Three

liquid fill volumes where included; 1/3, 1/2, and 2/3 filled, as were piping diameters of 15.9 mm and

25.4 mm. The calculations assume equilibrium bottle conditions. The specific amount of each

alternative agent equivalent in suppression capability to halon 1301 for engine nacelle protection was

estimated from the results reported in NIST SP 861 (Grosshandler et al., 1994) on the volume factor. The volume factor (VF) is a measure of the liquid volume of the agent in the storage vessel normalized by the liquid volume of halon 1301 required to suppress a flame under identical conditions. The

average value of the volume factors for the opposed flow diffusion burner, cup burner, spray burner,

and deflagration mbe was used. Table 41 shows the average volume factor and the required amount

of agent for equivalent protection. A major assumption here is that this average volume factor

prescribes the proper amount of alternative agent for the engine nacelle case. This must be validated

by actual engine nacelle testing, or a more appropriate relationship should be used (see Section 9). From the replacement agent liquid densities and the mass in the liquid phase required, only CF3I

has a chance to fit in the original bottle and therefore is a candidate for "drop-in" replacement. From

PROFISSY, 3.5 kg of CF3I in a 0.0042 m^ liter vessel pressurized to 4.12 MPa will yield 3.34 kg of

agent in the liquid phase with a liquid fill volume of 40.1 % at 23 °C. The liquid discharge time

through 4 m of 15.9 mm diameter piping is calculated to be 0.71 s. So, CF3I can be used as a direct

replacement of halon 1301. No more calculations with CF3I were performed.

HFC- 125 and HFC-227ea will require a larger bottle size than the 0.0042 m-^ halon bottle. Since

the search includes 3 liquid fill volumes and 4 pressures, a total of 12 bottle conditions were

examined. The exact bottle size is fixed by the liquid fill volume and pressure for each of the agents.

If bottle sizes are limited to off-the-shelf items, then a search of bottle size and pressure could be

performed. The bottle size and the total agent mass tends to increase as the pressure increases for a

fixed liquid fill volume. Tables 42 and 43 give the fill conditions for the various pressures and liquid

fill volumes and the calculated liquid discharge times for the simulations with the two pipe diameters.
8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 42. Fill conditions and liquid discharge times for HFC- 125

639

Storage

pressure (MPa)

Total mass

of agent

(kg)

Mass of

agent in liquid phase

(kg)

Fill volume

(%)

Storage

vessel volume

v'" )

Liquid discharge time

(s)

pipe I.D. 15.9 mm

pipe I.D. 25.4 mm

4.12 8.60 7.10 32.5 (1/3) 0.0200 2.00 0.75

4.12 7.80 7.08 49.9 (1/2) 0.0130 2.37 0.90

4.12 7.30 6.95 67.1 (2/3) 0.0095 2.86 1.10

5.49 8.75 6.97 33.0 (1/3) 0.0205 1.66 0.62

5.49 7.85 6.98 50.3 (1/2) 0.0135 1.98 0.74

5.49 7.40 6.98 67.9 (2/3) 0.0100 2.46 0.94

6.87 9.30 7.05 33.3 (1/3) 0.0220 1.52 0.56

6.87 8.20 7.02 48.7 (1/2) 0.0150 1.75 0.66

6.87 7.60 7.04 66.5 (2/3) 0.0110 2.18 0.82

8.24 9.95 7.01 32.9 (1/3) 0.0240 1.40 0.52

8.24 8.50 7.02 49.4 (1/2) 0.0160 1.65 0.61

8.24 7.80 7.06 66.3 (2/3) 0.0120 2.03 0.77

The liquid discharge times are also plotted against bottle pressure for the different fill conditions in Figures 146 and 147. A lower liquid fill volume and higher bottle pressure decrease the liquid

discharge time for a given pipe diameter. The rate of decrease in the liquid discharge time slows as

the bottle pressure is increased at any one fill condition.

The liquid discharge times for HFC- 125 through the 15.9 mm diameter piping are not under one

second for any fill condition and pressure, which necessitates an increase in piping diameter. At a

piping diameter of 25.4 mm, all liquid discharges are under 1 s except for the highest liquid fill volume and lowest bottle pressure (2/3 fill and 4.12 MPa). For HFC-227ea, the liquid discharge times

through the 15.9 mm diameter piping also are not under 1 s, but at the lowest liquid fill volume and

highest pressure, (1/3 filled and 8.24 MPa) the discharge time is nearly 1 s, which might warrant a

closer look. The penalty for the high pressure and low liquid fill volume is a relatively large bottle volume. Discharge times through the 25.4 mm diameter piping are all below 1 s for HFC-227ea.

The above exercise illustrates the vastly different hardware requirements for HFC- 125 and

HFC-227ea versus halon 1301 and CF3I. The results are indicative only of this particular case; other

cases require separate study, hence the importance of a general calculation scheme. Once the

particular alternative agent is selected and the design agent mass specified (using the volume factor

approach, or other suitable approach), this design scheme can be repeated for any number of retro-fits or new systems.

8.6.10 Conclusions and Recommendations. The experimental study and modeling effort of the

transient discharge of agent through piping presented here should assist in the design of engine nacelle
640 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table 43. Fill conditions and liquid discharge times for HFC-227ea

Storage

pressure (MPa)

Total mass

of agent

(kg)

Mass of

agent in liquid phase

(kg)

Fill volume

(%)

Storage

vessel volume (m )

Liquid discharge time

(s)

pipe I.D. 15.9 mm

pipe I.D. 25.4 mm

4.12 7.40 6.88 33.9 (1/3) 0.0150 1.46 0.56

4.12 7.15 6.88 50.0 (1/2) 0.0105 1.73 0.67

4.12 7.00 6.86 66.3 (2/3) 0.0079 2.13 0.84

5.49 7.50 6.87 33.6 (1/3) 0.0160 1.27 0.48

5.49 7.25 6.92 49.2 (1/2) 0.0110 1.48 0.57

5.49 7.05 6.89 66.6 (2/3) 0.0081 1.84 0.72

6.87 7.70 6.90 32.7 (1/3) 0.0170 1.14 0.43

6.87 7.30 6.90 49.7 (1/2) 0.0112 1.33 0.51

6.87 7.10 6.91 67.0 (2/3) 0.0083 1.66 0.65

8.24 7.85 6.88 32.9 (1/3) 0.0175 1.07 0.40

8.24 7.35 6.88 50.0 (1/2) 0.0115 1.25 0.48

8.24 7.10 6.87 67.6 (2/3) 0.0085 1.55 0.60

fire protection hardware systems which specify a replacement candidate. The observed flow characteristics for all of the replacement candidates are similar to halon 1301. Non-linear pressure drop versus pipe length increment indicates the high compressibility of the flows, while high speed movies of

transparent piping sections validate the two-phase nature of the flow after it leaves the storage bottle. Discharge times for the initially liquid phase in the bottle for fixed fill volume, pressure, and piping

condition did not vary widely for the three alternative agents and halon 1301. The amount of mass

discharged did vary from agent to agent due mostly to the density differences. Liquid flow times from

the small pipe diameters were approximately a factor of 2.5 greater than those for the large pipe

diameter. There was significant scatter in the observed bottle pressure where the pressure recovery

occurred (the lowest bottle pressure achieved before it began to rise) for any given agent. In most

cases, model calculations of the steady-state mass flow rates and transient discharge times are within

± 10 % of the experimental values, suggesting that the model can be used as a design tool for

hardware selection of an alternative agent fire suppression system. A hardware design scheme for

retro-fitting a halon 1301 system illustrated the model's utility.

Some areas of further research for transient discharge would include the behavior of a squib

activated discharge, its interaction with nitrogen degassing, and the degassing phenomena in general.

Discharges at elevated bottle pressure and larger piping diameters should be performed to validate the

model for those conditions. Unbalanced flow networks should be studied if they are important for

engine nacelle application.
OPTIMIZATION OF SYSTEM DISCHARGE 641

0)

E

IO

(0

o

Q

73

I I I I

I

I I I I

I

I I I I

I

I I I I

I

I I > I

I

I I I I

I

I I I I

I

I I I I

- - 1/3 fill 15.9 mm I.

- - 1/2 fill 15.9 mm I.D.

2/3 fill 15.9 mm I.D.

s., •o....

•o--:::::::g:: ^ o

,...[7}..

1/3 fill 25.4 mm I.D.

1/2 fill 25.4 mm I.D.

2/3 fill 25.4 mm I.D.

I I I I ' ' ' ' '

I ' ' ' ' 'I ' ' ' '

I ' 'I ' ' ''

5 6 7 8 9 10

Storage Pressure (MPa)

11 12

Figure 146. Calculated liquid discharge times for HFC- 125 for a various liquid fill volumes and

storage pressures. The mass in the liquid-phase was held constant. Calculations assume equilibrium bottle conditions.
642 8. OPTIMIZATION OF SYSTEM DISCHARGE

0)

Ui

JZ

o

tii 5

'5

1 1 11 1 1 1f |-r"^~^-^ 1 1 1 1

1

1 1 1 1

-

- -• - - 1/3 fill

\

15.9 mm I.D.

"

- - - - 1/2 fill 15.9 mm I.D.

2/3 fill 15.9 mm I.D. -

- X

- -o - 1/3 fill '• ^ 25.4 mm I.D.

- " - ^ _

----D--- 1/2 fill

- - - . ^

25.4 mm I.D.

"

---<£>--- 2/3 fill 25.4 mm I.D.

'

D

1 1 1 1 1 1 1 1 1 L.i_i-i-i J .1 1 .1L 1 \ .l.J 1 1 1 1 1 ,1.1 1 1 1 1 I 1 1

4 5 6 7 8 9 10 11 12

Storage Pressure (MPa)

Figure 147. Calculated liquid discharge times for HFC-227ea for a various liquid fill volumes

and storage pressures. The mass in the liquid-phase was held constant. Calculations assume equilibrium bottle conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 643

8.7 Concluding Remarks

Several important aspects related to agent storage and its subsequent discharge have been studied in

detail. They are summarized in the following.

1 . A computer program (PROFISSY) for calculating binary vapor-liquid equilibrium has been

developed to aid the bottle designers by providing useful and reliable data on the pressure- temperature characteristics of the selected agent/nitrogen mixtures.

2. The dissolved nitrogen in the liquid agent was found to play a very important role in the

discharge whether in a dry bay or in a piping system.

3. In simulated dry bay discharges, CF3I has the fastest volumetric discharge rate but the slowest

dispersion velocities.

4. A discharge model which incorporates nitrogen degassing has also been developed. The

model predicts agent discharge times within a factor of four (usually much better) without any

a priori information other than the initial conditions of the vessel.

5. The exploratory study using complex computer codes {e.g., KIVA-II) has identified the

specification of the initial spray condition and the rapid evaporation of the droplets as problem

areas. Methods need to be developed further to handle the highly transient flashing process in a computationally efficient manner before the detailed velocity, temperature, and concentration

field predictions which have been made can be properly interpreted.

6. A computer model based on a homogeneous two-phase flow has also been developed to

facilitate transient pipe-flow calculations for engine nacelle applications. The model has been

validated by the experimental results. A design exercise was given to illustrate the usefulness

of the model in system hardware selection for a halon alterative.

Figure 148 typifies a generic flow chart for discharge system optimization. To optimize a new

discharge system, system operational requirements {e.g., amount of agent required to suppress a fire) and constraints {e.g., vessel growth potential, operating temperature) have to be identified or provided

to the system designers. Once these constraints are known, the designers can then proceed to size the

bottle using thermodynamic data of the agent {e.g., PROFISSY). After the bottle size and the

allowable nitrogen pressurization have been determined, the discharge time of the agent can be

estimated by using a discharge model {e.g., the one developed above) for dry bay applications. For

engine nacelle fire protection, the pipe-flow model discussed above can be used to size the pipe and to predict agent discharge time. These calculation processes may have to be iterated several times until

all the calculated parameters meet the system constraints and operational requirements.
644 8. OPTIMIZATION OF SYSTEM DISCHARGE

System constraints and operational requirements:

% system growth potential

required discharge time (/^g^) amount of agent required for suppression

operating temperature range 1

Determine vessel volume ( Vp and nitrogen

pressurization level ( Pf) using thermodynamic data

(e.g., PROnSSY)

1

^"^ W bay

Engine

nacelle Calculate discharge time ( i^)

{e.g., using the discharge model in Section 8.4)

Yes

Select piping size I

Calculate discharge time ( t^)

{e.g., using the pipe- flow model in Section 8.6)

Adjust V, andP;

Yes

Perform full-scale validations

Adjust pipe size

Figure 148. A generic flow chart for discharge system optimization.
8. OPTIMIZATION OF SYSTEM DISCHARGE 645

8.8 Acknowledgments

This work was funded by the United States Air Force Wright Laboratory, Naval Air Systems

Command, and Army Aviation and Troop Command and the Federal Aviation Administration

Technical Center. Their financial support was greatly appreciated. The authors would like to thank

the following BFRL, NIST staff: Dr. William M. Pitts for his constant input to the project, Mr. Roy

McLane for fabricating some of the experimental hardware, and Ms. Paula Garrett for her effort in editing this section. Discussions regarding the degassing of dissolved nitrogen with Professor Charles

A. Ward of the University of Toronto were very helpful. The authors would also like to acknowledge

the following companies for providing the information on current halon 1301 system specifications:

Boeing, Booz-Allen & Hamilton, Lockheed Martin, McDonnell Douglas, Northrop Grumman, Pacific

Scientific, Systron-Donner, and Walter Kidde.

8.9 References

Allied Signal Inc., Selected Physical Properties: HCFC-123, HCFC-123a, HCFC-124, HFC-125,

HFC-134, HFC-134a, and HFC-143a, Allied Signal Inc., Buffalo Research Laboratory, Buffalo, NY,

1990.

Amsden, A. A., O'Rourke, P.J., and Butler, T.D., KIVA-II: A Computer Program for Chemically

Reactive Flows with Sprays, Los Alamos National Laboratory Report No. LA-11560-MS, 1989.

Amsden, A. A., Butler, T.D., O'Rourke, PJ., KIVA-III: A KIVA Program with Block-structured

Mesh for Complex Geometries, Los Alamos National Laboratory Report No. LA-12503-MS, 1993.

ASHRAE Inc., ASHRAE Thermodynamic Properties of Refrigerants, ASHRAE Inc., New York,

1969.

Balzhiser, R.E., Samuels, M.R., and Eliassen, J.D., Chemical Engineering Thermodynamics, The

Study of Energy, Entropy, and Equilibrium, Prentice-Hall, Englewood Cliffs, NJ, 1972.

Braker, W., and Mossman, A.L., Matheson Gas Data Book, 6th ed., Matheson Gas Products, New

Jersey, 1980.

Carnahan, B., Luther, H.A., and Wilkes, J.O., Applied Numerical Methods, John Wiley & Sons,

New York, 1969.

Carey, V. P., Liquid-Vapor Phase-Change Phenomena: An Introduction to the Thermophysics of

Vaporization and Condensation Processes in Heat Transfer Equipment, Hemisphere Publishing Co.,

Washington, DC, 1992.

Celata, G.P., Cumo, M., Farello, G.E., and Incalcaterra, PC, "On the Critical flows of Subcooled

Liquids," in Heat Transfer 1982, Proceedings of the Seventh International Heat Transfer Conference,

Grigull, U., Hahne, E., Stephan, K., and Straub, J. (eds.). Hemisphere, New York, 1982.

Celata, G.P, Cumo, M., Farello, G.E., Incalcaterra, PC, and Naviglio, A., "Thermodynamic

Disequilibrium in the Critical Flow of Subcooled Liquids," Nuclear Technology 60, 137 (1983).

Chueh, PL. and Prausnitz, J.M., "Vapor-Liquid Equilibria at High Pressurs: Calculation of Partial Molar Volumes in Nonpolar Liquid Mixtures," AIChE J. 13, 1099 (1967).

Cloutman, L.D., Dukowicz, J.K., Ramshaw, J.D., and Amsden, A.A., Conchas-Spray: A Computer

Code for Reactive Flows with Fuel Sprays, Los Alamos National Laboratory Report No. LA-9294-MS,

1982.

Cooper, L.Y., "The Dispersion of Fire Suppression Agents Discharged from High Pressure

Vessels: Establishing Initial/Boundary Conditions for the Flow Outside the Vessel," Proc. lCLASS-94,

Rouen, France, 1994.
646 8. OPTIMIZATION OF SYSTEM DISCHARGE

Dinneno, P.J., and Budnick, E.K., Halon 1301 Discharge Testing: A Technical Analysis, National

Fire Protection Research Foundation, Quincy, MA, 1988.

Dukler, A.E., Wicks, M., and Cleveland, R.G., "Frictional Pressure Drops in Two-Phase Flow,"

AlChEJ. 10, 44 (1964).

Dukowicz, J.K., Quasi-Steady Droplet Phase Change in the Presence of Convection, Los Alamos

Scientific Laboratory Report No. LA-7997-MS, 1979.

Dymond, J.H. and Smith, E.B., The Virial Coefficients of Gases: A Critical Compilation,

Clarendon Press, Oxford, 1969.

Elliot, D.G., Garrison, P.W., Klein, G.A., Moran, K.M., and Zydowicz, M.P.. Flow of NitrogenPressurized Halon 1301 in Fire Extinguishing Systems, Jet Propulsion Laboratory, JPL Publication 84-

62, 1984.

Ely, J.F. and Hanley, H.J.M., A Computer Program for the Prediction of Viscosity and Thermal

Conductivity in Hydrocarbon Mixtures, NBS Technical Note 1039, U.S. Department of Commerce,

Washington, DC 1981a.

Ely, J.F. and Hanley, H.J.M., "Prediction of Transport Properties. 1. Viscosity of Fluids and

Mixtures," Ind. Eng. Chem. Fundam. 20, 323 (1981b).

Ely, J.F., "Prediction of Dense Fluid Viscosities in Hydrocarbon Mixtures," Proc. 61st Gas

Processors Association, pp. 9-17 (1982).

Ely, J.F. and Hanley, H.J.M., "Prediction of Transport Properties. 2. Thermal Conductivity of Pure

Fluids and Mixtures," Ind. Eng. Chem. Fundam. 22, 90 (1983).

Ely, J.F., "A Predictive, Exact Shape Factor Extended Corresponding States Model for Mixtures,"

Advances in Cryogenic Engineering 35, 1511 (1990).

Forest, TW. and Ward, C.A., "Effect of a Dissolved Gas on the Homogeneous Nucleation

Pressure of a Liquid," J. Chem. Phys. 66, 2322 (1977).

Friend, D.G., N1ST Standard Reference Database 14: N1ST Mixture Property Database, U.S.

Department of Commerce, Washington, DC, Version 9.08, 1992.

Gallagher, J., McLinden, M., Huber, M., and Ely, J., NIST Standard Reference Database 23:

Thermodynamic Properties of Refrigerants and Refrigerant Mixtures Database (REFPROP), Version

3.0, U.S. Department of Commerce, Washington, DC 1991.

Gallagher, J., Mclinden, M., Huber, M., and Ely, J., NIST Standard Reference Database 23:

Thermodynamic Properties of Refrigerants and Refrigerant Mixtures Database (REFPROP), Version

4.0, U.S. Department of Commerce, Washington, DC 1993.

Gallagher, J., Huber, M.L., Morrison, G., and McLinden, M.O., NIST Standard Reference

Database 23: Thermodynamic Properties of Refrigerants and Refrigerant Mixtures, Version 5.0, U.S.

Department of Commerce, Washington, DC, 1995 (in preparation).

Great Lakes Chemical Corporation, FM-200TM: Environmentally Superior Fire Extinguishant for

Total Flooding Applications, Great Lakes Chemical Corporation, West Lafayette, IN, May 1993.

Grosshandler, W.L., Gann, R.G., and Pitts, W.M. (ed.). Evaluation of Alternative In-Flight Fire

Suppressants for Full-Scale Testing in Simulated Aircraft Engine Nacelles and Dry Bays, NIST

SP 861, U.S. Department of Commerce, Washington, DC, April 1994.

Henry, R.E., "Calculational Techniques for Two-Phase Critical Flow," in Two-Phase Flow

Dynamics, Japan-U.S. Seminar 1979, Bergles, A.E. and Ishigai, S. (eds.), pp. 415-436, Hemisphere

Publishing Corp., Washington, 1979.

Huber, M.L. and Ely, J.F., NIST Standard Reference Database 4: NIST Thermophysical Properties of Hydrocarbon Mixtures, U.S. Department of Commerce, Washington, DC, 1990.
8. OPTIMIZATION OF SYSTEM DISCHARGE 647

Huber, M.L. and McLinden, M.O., "Thermodynamic Properties of R134a (1,1,1,2-

tetrafluoroethane)," Proc 1992 International Refrigeration Conference, Purdue University, July 14-17,

453 (1992).

Huber, M.L. and Ely, J.F., "Prediction of Viscosity of Refrigerants and Refrigerant Mixtures,"

Fluid Phase Equilibria 80, 239 (1992).

Huber, M.L., Friend, D.G., and Ely, J.F., "Prediction of the Thermal Conductivity of Refrigerants

and Refrigerant Mixtures," Fluid Phase Equilibria 80, 249 (1992).

Huber, M.L. and Ely, J.F., "A Predictive Extended Corresponding States Model for Pure and

Mixed Refrigerants Including an Equation of State for R134a," Int. J. Refrig. 77, 18 (1994).

Jacobsen, R.T. and Stewart, R.B., "Thermodynamic Properties of Nitrogen Including Liquid and

Vapor Phases from 63 K to 2000 K with Pressures to 10,000 Bar," J. Phys. Chem. Ref Data 2, 757

(1973).

Kanury, A.M., Introduction to Combustion Phenomena, Gordon and Breach Science Publishers,

New York, 1977.

Kim-E, M.E. and Reid, R.C., "The Rapid Depressurization of Hot, High Pressure Liquids or

Supercritical Fluids," in Chemical Engineering at Supercritical Fluid Conditions, Paulaitia, M.E.,

Penniger, J.M.L., Gray, Jr., R.D., and Davidson, P. (eds.), pp. 81-100, Ann Arbor Science, Ann Arbor,

MI, 1983.

Kudchadker, A.P., Kudchadker, S.A., Shukla, R.P, and Patnaik, PR., "Vapor Pressures and

Boiling Points of Selected Halomethanes," J. Phys. Chem. Ref Data 8, 499 (1979).

Kurschat, Th., Chaves, H., and Meier, G.E.A., "Complete Adiabatic Evaporation of Highly

Superheated Liquid Jets," J. Fluid Mechanics 236, 43 (1992).

Leach, J.W., Chappelear, P.S., and Leland, T.W., "Use of Molecular Shape Factors in Vapor

Liquid Equilibrium Calculations with the Corresponding States Principle," AIChE Journal 14, 568

(1968).

McCarty, R.D., "Mathematical Models for the Prediction of Liquified-Natural-Gas Densities," J. Chem. Thermodynamics 14, 837 (1982).

National Fire Codes, NFPA I2A Standard on Halon 1301 Fire Extinguishing Systems 1992

Edition, National Fire Protection Association, Quincy, MA, 1995.

Olujic, Z., "Predicting Two-Phase-Flow Friction Loss in Horizontal Pipes," Chemical Engineering,

1985.

Perry, R.H., Green, D., and Maloney, J.O., Perry's Chemical Engineers' Handbook, 6th ed.,

McGraw-Hill, New York, NY, 1984.

Prausnitz, J.M., Anderson, T.F., Grens, E.A., Eckert, C.A., Hsieh,R., and O'Connell, J. P., Computer Calculations for Multicomponent Vapor-Liquid and Liquid-Liquid Equilibria, Prentice-Hall,

New Jersey, 1980.

Ranz, W.E., and Marshall, W.R., "Evaporation from Drops. Part I and II," Chemical Engineering

Progress 48, U\ &. \13 {\952).

Reid, R.C., "Superheated Liquids," American Scientists 64, 146 (1976).

Reid, R.C., Prausnitz, J.M., and Poling, B.E., The Properties of Gases and Liquids, 4th ed.,

McGraw-Hill, New York, 1987.

SFPE Handbook of Fire Protection Engineering, First ed.. Society of Fire Protection Engineers,

National Fire Protection Association, Quincy, MA, 1988.

Sherman, G.J., Magee, J.W., and Ely, J.F, "PVT Relationships in a Carbon Dioxide Rich Mixture

with Ethane," Int. J. Thermophysics 10, 47 (1989).

Sladkov, I.B., and Bogacheva, A.V., "Critical Parameters of Mixed Carbon Halides," Zhurnal

Prikladnoi Khimii 64, 2435 (1991); 64, 2276 (1992) (English translation).
648 8. OPTIMIZATION OF SYSTEM DISCHARGE

Smith, B.D. and Srivastava, R., Thermodynamic Data for Pure Compounds. Part B Halogenated

Hydrocarbons and Alcohols, Elsevier, Amsterdam, 1986.

Smith, J.M., and Van Ness, H.C., Introduction to Chemical Engineering Thermodynamics, 3rd ed.,

McGraw-Hill, New York, 1975.

Taylor, B.N., and Kuyatt, C.E., Guidelines for Evaluating and Expressing the Uncertainty of N1ST

Measurement Results, NIST Technical Note 1297, U.S. Department of Commerce, Washington, DC,

1994.

Thompson P.A., Carofano G.C., and Kim Y.G., "Shock Waves and Phase Changes in a LargeHeat-Capacity Fluid Emerging from a Tube," J. Fluid Mechanics 166, 57 (1986).

Volmer, M., Kenetik der Phasenbildung, Edwards Brothers, Inc., Ann Arbor, MI, 1945.

Walas, S.M., Phase Equilibrium in Chemical Engineering, Butterworth, Boston, 1985.

Welty, J.R., Wicks, C.E., and Wilson, R.E., Fundamentals of Momentum, Heat, and Mass

Transfer, 3rd ed., John Wiley & Sons, New York, NY, 1984.

Whalley, P.B., Boiling Condensation and Gas-Liquid Flow, Clarendon Press, Oxford, 1990.
8. OPTIMIZATION OF SYSTEM DISCHARGE 649

Appendix A. A User's Guide to PROFISSY

There are two files (PROFISSY.EXE and PROFLIB) in the disc. The current version of PROFISSY is Version 0.08. In the following, the bold type style represents any text one must type. Sentences in

italic represent the instructions given by the computer. To start a run, simply type PROFISSY,

followed by hitting the <enter> key, from the drive or directory where the program is installed. The

screen will appear as follows.

NISI Program PROFISSY - Version 0.08 PROperties of Fire Suppression SYstems

National Institute of Standards and Technology

Thermophysics Division, 838.02 M. L. Huber, 303-497-5252

Based on research sponsored by

United States Air Force, Army, and Navy, and

the Federal Aviation Administration

NOTE: NIST makes every effort to deliver a high

quality product; however — NIST shall not be

liable for any damage that may result from errors or omissions in this program.

The program output will be written to file OUTDAT and to the screen. The default units are F, lb., ft**3, Psia Do you want to change the units ? (Yes/No)

Type y and hit <enter> if you want to change the default units. The Property Unit Menu will appear on the screen.
650 8. OPTIMIZATION OF SYSTEM DISCHARGE

Property Unit Menu

Property current

Temperature (K ,R ,c ,F , )- F

Pressure (atm ,bar ,MPa ,kpf ,Psia ,mmHG)- Psia

Volume (liter .m**3 ,cm**3 ,in**3 ,ft**3 ft**3

Energy (cal •J ,btu ,kcal ,y btu

Mass (mol ,lb-mol ,kg -g ,1b lb

Velocity (m/s ,cm/s ,ft/s ,in/s ft/s

Viscosity (uP ,cP ,uPa.s ,lb/ft.s ,lb/ft.h Ib/ft.s

Thm. Cond. W/m.K ,cal/cm.s.K ,btu/ft.s.F ,btuyft.h.F ,mW/m.K btu/ft.h.F

default set options

(1) Scientific (K ,atm .liter ,cal ,mol ,cm/s ,uP ,cal/cm.s.K)

(2) S.I. (K ,MPa ,m**3 ,U ,kg ,m/s ,uPa.s , W/m.K )

(3) Engineering (F ,Psia ,ft**3 ,btu ,Ib ,ft/s ,lb/ft.s ,btuyft.h.F)

(4) Mixed (K ,bar ,liter ,kJ ,mol ,m/s ,uP ,mW/m.K )

Enter the new unit or default option (X to exit) ?

Type C,MPa,cin**3,g,x and hit <enter> if you want temperature in C, pressure in MPa, volume

in cm , and mass in g. The screen will appear as follows.

Please input the fire suppression agent. Choose from the following list: Name Synonym

CF3I trifluoroiodomethane : R218 perfluoropropane

R125 pentafluoroethane

R227ea heptafluoropropane

R13bl trifluorobromomethane

Type cf3i and hit <enter> if you want to run CF3I. Follow the instructions given by the

computer to input the data.
8. OPTIMIZATION OF SYSTEM DISCHARGE 651

Input the amount (in g) offire suppression agent.700 <enter>

Input the container volume (in cm **i^500 <enter>

Do you want to input system pressure (P) or N2 mass (M) ?p <enter>

Input the pressure (in MPa)4.l2 <enter>

Input the fill temperature (in C)22 <enter>

Do you want to specify the temperature limits for the table? (Y,N)

If NO, limits will be generated automatically y <enter>

If the input is n, the computer will generate the table from -60 to 150 °C with 10 °C increment.

Input the Tmin, Tmax, and T increment ( in Q20,40,10 <enter>

The computer will response: please wait, computations in progress ....

After the computation, the screen will appear as follows.

CF3I with nitrogen superpressure

Pressure - Temperature Characteristics

Agent mass 700.000(g)

Nitrogen Mass 15.522(g)

Container Volume 5OO.00O(cm**3)

Fill Temperature 22.000(C) Fill Pressure 4.120(MPa)

Bubble point temp. 92.006(C) Bubble point pres. 6.784(MPa)

Press Enter to Continue

After hit <enter>, the computer will list the calculated results on the screen. The calculated

bubble point pressure and temperature are obtained at the fill conditions.
652 8. OPTIMIZATION OF SYSTEM DISCHARGE

Temp. Vol. Pet. Mass Liquid Nitrogen Mass N2 (g) Pressure (C) Liq. Fill CF3I (g) xN2 wN2 liquid vapor MPa

20.0 68.99 0.6910E+03 0.085 0.013 0.9129E+0] 0.6393E+01 4.1

30.0 70.83 0.6890E+03 0.088 0.014 0.9558E+01 0.5964E+01 4.3

40.0 72.90 0.6869E+03 0.093 0.014 0.1005E+02 0.5473E+01 4.6 Do you want another run? (Yes,No)

If you do not want another run, type n and press <enter>.

Goodbye !

"Vol. Pet. Liq. Fill" represents the volume percent of liquid fill in the container. "xN2" and

"wN2" represent the mole and mass fractions of nitrogen in the liquid mixture respectively.

Compilation of Results. For quick references, three cases with initial pure (before nitrogen

pressurization) liquid agent volume fractions of 1/3, 1/2, and 2/3 (see Section 8.3.2.1) have been

run for all the agents which the computer code currently supports and are summarized in Figures

A-1 to A- 15. In the calculations, the initial fill temperature is 22 °C and the three initial fill pressures are 2.75 MPa, 4.12 MPa, and 5.50 MPa.
8. OPTIMIZATION OF SYSTEM DISCHARGE 653

i

!

HFC-227ea / nitrogen mixture ^ 1/3 fill • 1/2 fill Pj = 2.75 MPa @ 22°C v 2/3 fill

-100 -50 0 50 100 150 200

Temperature Cc)

Figure A-1. Calculated pressure as a function of temperature for HFC-227ea/N2 mixture

under three fill conditions and = 2.75 MPa.
654 8. OPTIMIZATION OF SYSTEM DISCHARGE

24

CD

20 -

16

12 -

0

-100

1 1 1

HFC-227ea / nitrogen mixture

P. = 4 12 MPa (5) 22 °C

1 1

o 1/3 fill • 1/2 fill V^ 9/'^ Z/j fill 1111 7

/

/

/

/

/

/

y r 0

1 1 1

1 1 -50 0 50 100 150

o Temperature ( C)

200

Figure A-2. Calculated pressure as a function of temperature for HFC-227ea/N2 mixture

under three initial fill conditions and P- = 4.12 MPa.
8. OPTIMIZATION OF SYSTEM DISCHARGE 655

Oh

Oh

28

24

20

16

12

- HFC-227ea / nitrogen mixture ^ 1/3 fill 1/2 fill P. = 5.50 MPa @ 22 C

100 -50

_L 0 50 100

o Temperature ( C)

150 200

Figure A-3. Calculated pressure as a function of temperature for HFC-227ea/N2 mixture

under three initial fill conditions and P- = 5.50 MPa.
656 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure A-4. Calculated pressure as a function of temperature for CF3I/N2 mixture under

three initial fill conditions and = 2.75 MPa.
OPTIMIZATION OF SYSTEM DISCHARGE 657

CFJ / nitrogen mixture ° 1/3 fill • 1/2 fill Pj = 4.12MPa@22''C - 2/3 fill

/ /

-100 -50 0 50 100 150 200

Temperature (^C)

Figure A-5. Calculated pressure as a function of temperature for CF3I/N2 mixture under

three initial fill conditions and P, = 4.12 MPa.
658 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure A-6. Calculated pressure as a function of temperature for CF3I/N2 mixture under

three initial fill conditions and P- = 5.50 MPa.
OPTIMIZATION OF SYSTEM DISCHARGE 659

20

16

FC-218 / nitrogen mixture

P. = 2.75 MPa m 22 C

1/3 fill 1/2 fill 2/3 fill

OS Oh

0^

12

/ , 0

0

100 -50 0 50 100 150 200

Temperature ( C)

Figure A-7. Calculated pressure as a function of temperature for FC-2I8/N2 mixture under

three initial fill conditions and P, = 2.75 MPa.
660 8. OPTIMIZATION OF SYSTEM DISCHARGE

i

Oh

CO

CD

24

20

16

12

FC-218 / nitrogen mixture

P: = 4.12MPa@22 C

1/3 fill 1/2 fill 2/3 fill

/ /

/

/

0

-100 -50 0 50 100 150 200

Temperature ( C)

Figure A-8. Calculated pressure as a function of temperature for FC-2I8/N2 mixture under

three initial fill conditions and P- = 4.12 MPa.
OPTIMIZATION OF SYSTEM DISCHARGE 661

28

Oh

24

20

16

12

FC-218 / nitrogen mixture

P. = 5.50 MPa 22 °C

1/3 fill 1/2 fill 2/3 fill

0

100 -50 0 50 100 150 200

Temperature ( C)

Figure A-9. Calculated pressure as a function of temperature for FC-2I8/N2 mixture under

three initial fill conditions and P, = 5.50 MPa.
662 8. OPTIMIZATION OF SYSTEM DISCHARGE

24

Oh

CD

20

16

12

0

HFC- 125 / nitrogen mixture ^ 1/3 fill 1/2 fill Pj = 2.75 MPa @ 22 "C ^ 2/3 fill

100 -50 0 50 100 150 200

Temperature ( C)

Figure A- 10. Calculated pressure as a function of temperature for HFC-I25/N2 mixture

under three initial fill conditions and = 2.75 MPa.
OPTIMIZATION OF SYSTEM DISCHARGE 663

28

Oh

24

20

16

12

HFC- 1 25 / nitrogen mixture ^ 1/3 fill /

P. = 4.12 MPa @ 22 C

1/2 fill V 2/3 fill 7

0

-100 -50 0 50 100 150 200

Temperature ( C)

Figure A- 11

.

Calculate pressure as a function of temperature for HFC-I25/N2 mixture under

three initial fill conditions and P- = 4.12 MPa.
664 8. OPTIMIZATION OF SYSTEM DISCHARGE

32

28

24 -

20

HFC- 125 / nitrogen mixture

P. = 5.50 MPa @ 22 C

1/3 fill 1/2 fill /

2/3 fill 7

/

/

/

r

16

12 -

8

-

4

-

0

/

/

0

100 -50 50 100 150 200

Temperature ( C)

Figure A- 12. Calculated pressure as a function of temperature for HFC-125/Nt mixture

under three initial fill conditions and P- = 5.50 MPa.
8. OPTIMIZATION OF SYSTEM DISCHARGE 665

24

20

16

12

0

-100

CF.Br / nitrogen mixture ^ 1/3 fill • 1/2 fill P. = 2.75 MPa@22 C - 2/3 fill

-50 0 50 100 150 200

Temperature ( C)

Figure A- 13. Calculated pressure as a function of temperature for CF3Br/N2 mixture under

three initial fill conditions and P, = 2.75 MPa.
666 8. OPTIMIZATION OF SYSTEM DISCHARGE

28

24

20

CF-Br / nitrogen mixture

P. = 4.12 MPa 1 22

1/3 fill 1/2 fill 2/3 fill

Oh

16

12 -

8

-

/

/

/.-.

7,*

,o

0

-100 -50 0 50 100 150 200

Temperature ( C)

Figure A- 14. Calculated pressure as a function of temperature for CF3Br/N2 mixture under

three initial fill conditions and = 4.12 MPa.
OPTIMIZATION OF SYSTEM DISCHARGE 667

32

28 -

24

1^ 20

Oh

16 -

12

CF-Br / nitrogen mixture

P. = 5.50 MPa @ 22 C

1/3 fill 1/2 fill J

2/3 fill

0

•100 -50 0 50 100 150 200

Temperature ( C)

Figure A- 15. Calculated pressure as a function of temperature for CF3Br/N2 mixture under

three initial fill conditions and = 5.50 MPa. ,
668 8. OPTIMIZATION OF SYSTEM DISCHARGE

Appendix B. Tabulated fill Conditions and Pressure Traces for All Pipe

Flow Discharge Tests

This Appendix lists the experimental conditions and results for all the pipe flow tests performed in this study. Table B-1 tabulates the temperatures and fill conditions for hot and cold agent discharge tests. Tables B-2 to B-5 tabulate the experimental conditions for halon 1301, HFC- 125, HFC-227ea, and

CF3I, respectively. The fill voulmes were estimated by the computer code, PROFISSY. The

experimental results for halon 1301, HFC-125, HFC-227ea, and CF3I are shown in Figures B-1 to B-25, Figures B-26 to B-50, Figures B-51 to B-76, and Figures B-77 to B-99, respectively. The

individual pressure traces are not labeled on each figure, but follow the logical sequence of the vessel

pressure being the highest, then followed in descending order by the pipe pressure traces for locations

successively farther downstream. Finally, the pressure trace of the recover tank is the lowest. In some

tests, not all pressures at each tap location were recorded. In those cases, the location of the tap where

data was not taken is noted.
8. OPTIMIZATION OF SYSTEM DISCHARGE

Table B- 1 . Temperature and fill conditions for hot and cold agent discharge tests

669

Agent Figure

number

Initial

fill volume

(%)

Liquid

temp.

(°C)

Gas

temp.

(°C)

Final pressure (± 0.05 MPa)

Equilibirum

fill volume

(%)

Equilibrium

pressure (MPa)

CFjBr B-24 48.9 -50 4 2.20 41.8 2.0

B-25 49.1 70 66 8.00 Single phase 7.4

HFC- 125 B-49 46.4 -44 4 2.55 39.1 2.3

B-50 49.3 68 73 7.40 Single phase 7.4

HFC-227ea B-74 49.2 -45 4 3.00 43.5 2.8

B-75 48.8 68 72 6.00 58.4 5.5

B-76 49.0 70 68 5.95 59.8 5.6

CF3I B-98 50.4 -46 4 3.05 44.2 2.7

B-99 49.7 70 70 5.70 56.7 5.7
670

Table B-2. Test fill conditions for halon 1301

8. OPTIMIZATION OF SYSTEM DISCHARGE

Storage

pressure

(+ 0 DOS MPnl

Agent mass

(± 0.01 kg)

Fill volume

(%)

Diameter (mm)

Configuration Figure

number

^J .jyj TO Q s y.j Straight pipe rS1

1

1

J . 1

1

M A Qy S.J rj 9 rS-Z

4n MIN .A.A 7.3 13 9

9 77 9^; J .zu tU.D y .J RD-4A

A 1 9 *f . 1

Z

9Z.ooi?8 Q S y.j D-3

4. 1 3

fin Qy S.J £>-D

A I 7 J. I J 7.3 n 7

H. I z _J . J ^ 50.4 y .J RD-o8

4 1 9 4. 1 Z 44.Z9 11 7.3 D-y R 0

44.1/1 9 J.J I 7.3 9 m Qn° h*anH 3 m, 7U DcnQ R 1 n D- lU

y| 1 9 4. 1 Z

1 19 0 '\ 7.3 A m Qn° K^irtrf 4 m, yu oeno R 1 1 t>- 1 1

A 1 9 4. 1

Z

T 99 4R 1 ^5 Q/Q S 1 3.7/7.3 Contraction R 1 9 D- 1 Z

/I 1 9 4. 1 Z

"5 9Q J.Zo 7.3/ 1 3.7 Expansion R I 9 tS- 1 3

9 nn z.UU /I no 4.uy IN. A. 11 3.7 ^ Q Straight pipe R 1 A D- 14

Z.5U i.oj IN. A. 1 ^ Q 1 3.V R 1 ^ D-13

T nn A A1 4.4 J M A IN. A. 1 3.7 R 1 A D- 10

9 9^ Z. / J9 ^9 ^ 1 1 31.1 13.

y

R 17 D- 1 /

y1 1 9 4.1Z 9 9^ j.Zj Al C 4 /.O 1 3.y R 1 Q D- 1 8

/I 1 9 9 9^ i.jj 13.

y

R Id D-iy

4.12 4.20 65.5 15.9 B-20

4.12 3.30 48.7 15.9 3 m, 90° bend B-21

4.12 3.30 48.7 15.9 4 m, 90° bend B-22

4.12 3.26 47.9 15,9 "T" flow split B-23

4.12 3.31 48.9 15.9 Cold B-24

4.12 3.32 49.1 15.9 Hot B-25
8. OPTIMIZATION OF SYSTEM DISCHARGE

Table B-3. Test fill conditions for halon HFC- 125

671

Storage

(± 0.005 MPa)

Agent mass

(+ 0 01 kg)

Fill volume

(%)

Diameter (mm)Configuration Figure m1 lUIimhf 1 lU^lr

2.00 2.51 N.A. 9.5 Straight pipe B-26

2.50 2.92 N.A. 9.5 B-27

3.00 2.53 N.A. 9.5 B-28

4.12 3.69 N.A. 9.5 B-29

2.76 2.62 50.4 9.5 B-30

4.13 2.78 56.4 9.5

" B-31

4.16 3.33 69.9 9.5

• B-32

4.12 2.40 47.1 9.5 3 m, 90° bend B-33

4.12 2.44 48.1 9.5 4 m, 90° bend B-34

4.12 2.50 49.6 15.9/9.5 Contraction B-35

4.12 2.51 49.8 9.5/15.9 Expansion B-36

2.00 2.72 N.A. 15.9 Straight pipe B-37

2.55 3.29 N.A. 15.9 B-38

3.00 2.81 N.A. 15.9 B-39

2.75 2.79 54.2 15.9 B-40

4.12 2.50 49.6 15.9 B-41

4.12 2.63 52.7 15.9 B-42

4.12 2.64 53.0 15.9 B-43

4.12 3.17 65.9 15.9 B-44

4.14 3.82 81.8 15.9 B-45

4.12 2.47 48.8 15.9 3 m, 90° bend B-46

4.12 2.38 46.6 15.9 4 m, 90° bend B-47

4.12 2.70 54.4 15.9 "T" flow split B-48

4.12 2.37 46.4 15.9 Cold B-49

4.12 2.49 49.3 15.9 Hot B-50
672 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table B-4. Test fill conditions for halon HFC-227ea

Storage

pressure

(± 0.005 MPa)

Agent mass

(± 0.01 kg)

Fill volume

(%)

Diameter (mm)

Configuration Figure

number

2.00 3.31 N.A. 9.5 Straight pipe B-51

2.50 3.18 N.A. 9.5 B-52

3.00 3.24 N.A. 9.5 B-53

2.75 3.26 56.5 9.5 B-54

4.12 3.05 53.7 9.5 B-55

4.12 3,36 60.2 9.5

„ B-56

4.12 3.91 70.0 9.5 B-57

4.12 2.81 49.2 9.5 3 m, 90° bend B-58

4.12 2.81 49.2 9.5 4 m, 90° bend B-59

4.12 2.80 49.0 15.9/9.5 Contraction B-60

4.12 2.80 49.0 9.5/15.9 Expansion B-61

2.00 3.74 N.A. 15.9 Straight pipe B-62

2.50 2.73 N.A. 15.9 B-63

2.50 3.65 N.A. 15.9 B-64

3.00 3.56 N.A. 15.9 B-65

2.75 3.13 54.1 15.9 B-66

4.12 2.77 48.4 15.9 B-67

2.82 49.4 15.9 B-68

4.12 3.23 57.1 15.9

„ B-69

4.12 3.88 69.4 15.9 B-70

4.12 2.82 49.4 15.9 3 m, 90° bend B-71

4.12 2.79 48.8 15.9 4 m, 90° bend B-72

4.12 2.77 48.4 15.9 "T" flow split B-73

4.12 2.81 49.2 15.9 Cold B-74

4.12 2.79 48.8 15.9 Hot B-75

4.12 2.80 49.0 15.9 Hot B-76
8. OPTIMIZATION OF SYSTEM DISCHARGE

Table B-5. Test fill conditions for CF3I

673

Storage

pressure

(± 0.005 MPa)

Agent mass

(± 0.01 kg)

Fill volume

(%)

Diameter (mm)

Configuration Figure

number

2.00 5.40 N.A. 9.5 Straight pipe B-77

2.50 5.58 N.A. 9.5 B-78

3.00 5.22 N.A. 9.5 B-79

2.75 4.34 49.7 9.5 B-80

4.13 4.31 50.2 9.5 B-81

4.12 5.57 65.8 9.5 B-82

4.12 5.74 67.9 9.5 B-83

4.12 4.46 52.0 9.5 3 m, 90° bend B-84

4.12 4.39 51.2 9.5 4 m, 90° bend B-85

4.12 4.49 52.4 15.9/9.5 Contraction B-86

4.12 4.60 53.8 9.5/15.9 Expansion B-87

2.00 4.03 N.A. 15.9 Straight pipe B-88

2.50 4.21 N.A. 15.9 B-89

2.50 6.27 N.A. 15.9 B-90

3.00 4.32 N.A. 15.9 B-91

2.80 4.33 49.6 15.9 B-92

4.13 4.19 48.7 15.9 B-93

4.20 5.76 68.2 15.9 B-94

4.17 4.24 49.4 15.9 3 m, 90° bend B-95

4.14 4.71 55.2 15.9 4 m, 90° bend B-96

4.12 4.29 49.9 15.9 "T" flow split B-97

4.19 4.32 50.4 15.9 Cold B-98

4.12 4.27 49.7 15.9 Hot B-99
674 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure B-1. Halon 1301 test. Refer to Table B-2 for experimental conditions.
Figure B-2. Halon 1301 test. Refer to Table B-2 for experimental conditions.
676 8. OPTIMIZATION OF SYSTEM DISCHARGE

0 500 1000 1500 2000 2500 3000 3500 4000

Time (ms)

Figure B-3. Halon 1301 test. Refer to Table B-2 for experimental conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 677

Figure B-4. Halon 1301 test. Refer to Table B-2 for experimental conditions.
678 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure B-5. Halon 1301 test. Refer to Table B-2 for experimental conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 679

Figure B-6. Haion 1301 test. Refer to Table B-2 for experimental conditions.
Figure B-7. Halon 1301 test. Refer to Table B-2 for experimental conditions.
Figure B-8. Halon 1301 test. Refer to Table B-2 for experimental conditions.
Figure B-9. Halon 1301 test. Refer to Table B-2 for experimental conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 683

Figure B-10. Halon 1301 test. Refer to Table B-2 for experimental conditions.
Figure B-11 Halon 1301 test. Refer to Table B-2 for experimental conditions.
Figure B-12. Halon 1301 test. Refer to Table B-2 for experimental conditions.
686 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure B-13. Halon 1301 test. Refer to Table B-2 for experimental conditions.
OPTIMIZATION OF SYSTEM DISCHARGE 687

Figure B-14. Halon 1301 test. Refer to Table B-2 for experimental conditions.
688 8. OPTIMIZATION OF SYSTEM DISCHARGE

1—I—I—I—

I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—

I—I—I—I—I—I—I—I—I—

r

I I I I I I I I I I I I I I I—I—I—I—I—I—I—I—I—I—I—I—I—

L

0 500 1000 1500 2000 2500 3000

Time (ms)

Figure B-15. Halon 1301 test. Refer to Table B-2 for experimental conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 689

Figure B-16. Halon 1301 test. Refer to Table B-2 for experimental conditions.
690 8. OPTIMIZATION OF SYSTEM DISCHARGE

1—I—I—I—

j—I—1—I—I—I—I—I—I—I—I—I—I—I—I—

I—I—I—I—I—

I—I—I—I—

r

_j I I I I I I I I I I I I I I I I I—I—I—I—I—I—I—I—I—I—I—I

0 500 1000 1500 2000 2500 3000

Time (ms)

Figure B-17. Halon 1301 test. Refer to Table B-2 for experimental conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 691

Figure B-18. Halon 1301 test. Refer to Table B-2 for experimental conditions.
Figure B-19. Halon 1301 test. Refer to Table B-2 for experimental conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 693

Figure B-20. Halon 1301 test. Refer to Table B-2 for experimental conditions.
Figure B-21. Halon 1301 test. Refer to Table B-2 for experimental conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 695

Figure B-22. Halon 1301 test. Refer to Table B-2 for experimental conditions.
Figure B-23. Halon 1301 test. Refer to Table B-2 for experimental conditions.
OPTIMIZATION OF SYSTEM DISCHARGE 697

Figure B-24. Halon 1301 test. Refer to Table B-2 for experimental conditions.
Figure B-25. Halon 1301 test. Refer to Table B-2 for experimental conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 699

Figure B-26. HFC- 125 test. Refer to Table B-3 for experimental conditions.
Figure B-27. HFC- 125 test. Refer to Table B-3 for experimental conditions.
Figure B-28. HFC- 125 test. Refer to Table B-3 for experimental conditions.
702

5

0 1000 2000 3000 4000 5000 6000

Time (ms)

Figure B-29. HFC- 125 test. Refer to Table B-3 for experimental conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 703
Figure B-3I HFC- 1 25 test. Refer to Table B-3 for experimental conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 705

Figure B-32. HFC- 125 test. Refer to Table B-3 for experimental conditions.
Figure B-33. HFC- 125 test. Refer to Table B-3 for experimental conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 707

Figure B-34. HFC- 125 test. Refer to Table B-3 for experimental conditions.
Figure B-35. HFC- 125 test. Refer to Table B-3 for experimental conditions.
OPTIMIZATION OF SYSTEM DISCHARGE 709

Figure B-36. HFC- 125 test. Refer to Table B-3 for experimental conditions.
710 8. OPTIMIZATION OF SYSTEM DISCHARGE

3

2.5

If 2

0)1.5

3

0)

o

q: 1

0.5

T—I—I—I—

I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—

I—I—I—I—I—

I—I—I—I—

r

/

///

//

i—I—I—I—I—I—I—I—I—I I—I—I I I I I I I I I I I I I I I I I

0 500 1000 1500 2000 2500 3000

Time (ms)

Figure B-37. HFC- 125 test. Refer to Table B-3 for experimental conditions.
Figure B-38. HFC- 125 test. Refer to Table B-3 for experimental conditions.
Figure B-39. HFC- 125 test. Refer to Table B-3 for experimental conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 713

Figure B-40. HFC- 125 test. Refer to Table B-3 for experimental conditions.
Figure B-41. HFC- 125 test. Refer to Table B-3 for experimental conditions.
Figure B-42. HFC- 125 test. Refer to Table B-3 for experimental conditions.
Figure B-43. HFC- 125 test. Refer to Table B-3 for experimental conditions.
Figure B-44. HFC- 125 test. Refer to Table B-3 for experimental conditions.
Figure B-45. HFC- 125 test. Refer to Table B-3 for experimental conditions.
5

0 500 1000 1500 2000 2500 3000

Time (ms)

Figure B-46. HFC- 125 test. Refer to Table B-3 for experimental conditions.
720 8. OPTIMIZATION OF SYSTEM DISCHARGE

(0

O

3

0)

V>

o

0 500 1000 1500 2000 2500 3000

Time (ms)

Figure B-47. HFC- 125 test. Refer to Table B-3 for experimental conditions.
OPTIMIZATION OF SYSTEM DISCHARGE 721

Figure B-48. HFC-125 test. Refer to Table B-3 for experimental conditions.
722 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure B-49. HFC- 125 test. Refer to Table B-3 for experimental conditions.
OPTIMIZATION OF SYSTEM DISCHARGE 723

Figure B-50. HFC- 125 test. Refer to Table B-3 for experimental conditions.
724 8. OPTIMIZATION OF SYSTEM DISCHARGE

0 1000 2000 3000 4000 5000 6000

Time (ms)

Figure B-51. HFC-227ea test. Refer to Table B-4 for experimental conditions.
Figure B-52. HFC-227ea test. Refer to Table B-4 for experimental conditions.
Figure B-53. HFC-227ea test. Refer to Table B-4 for experimental conditions.
OPTIMIZATION OF SYSTEM DISCHARGE 727

Figure B-54. HFC-227ea test. Refer to Table B-4 for experimental conditions.
Figure B-55. HFC-227ea test. Refer to Table B-4 for experimental conditions.
Figure B-56. HFC-227ea test. Refer to Table B-4 for experimental conditions.
730 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure B-57. HFC-227ea test. Refer to Table B-4 for experimental conditions.
OPTIMIZATION OF SYSTEM DISCHARGE 731

Figure B-58. HFC-227ea test. Refer to Table B-4 for experimental conditions.
Figure B-59. HFC-227ea test. Refer to Table B-4 for experimental conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 733

5 r

T—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—1—I—

(0

Q.

O

V.

31

V)

O

t

I

I I I I I I—I—I—I—I—I—I—I—I—I—I—I—I—I—

J—I—I—

0 1000 2000 3000 4000 5000 6000

Time (ms)

Figure B-60. HFC-227ea test. Refer to Table B-4 for experimental conditions.
Figure B-61 . HFC-227ea test. Refer to Table B-4 for experimental conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 735

Figure B-62. HFC-227ea test. Refer to Table B-4 for experimental conditions.

i

II

I
Figure B-63. HFC-227ea test. Refer to Table B-4 for experimental conditions.
OPTIMIZATION OF SYSTEM DISCHARGE 737

2.5

2

01.5

3

(0

o

£ 1

0.5

T—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—

r

III I I I I I—I—1—1—I—I—I—J—I—I—I—I—I—I—I—I—I—I—

0 500 1000 1500 2000 2500 3000

Time (ms)

Figure B-64. HFC-227ea test. Refer to Table B-4 for experimental conditions.
738 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure B-65. HFC-227ea test. Refer to Table B-4 for experimental conditions.
Figure B-66. HFC-227ea test. Refer to Table B-4 for experimental conditions.
Figure B-67. HFC-227ea test. Refer to Table B-4 for experimental conditions.
OPTIMIZATION OF SYSTEM DISCHARGE 741

Figure B-68. HFC-227ea test. Refer to Table B-4 for experimental conditions.
Figure B-69. HFC-227ea test. Refer to Table B-4 for experimental conditions.
OPTIMIZATION OF SYSTEM DISCHARGE 743

Figure B-70. HFC-227ea test. Refer to Table B-4 for experimental conditions.
744 8. OPTIMIZATION OF SYSTEM DISCHARGE

0 500 1000 1500 2000 2500 3000

Time (ms)

Figure B-71. HFC-227ea test. Refer to Table B-4 for experimental conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 745

Figure B-72. HFC-227ea test. Refer to Table B-4 for experimental conditions.
746 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure B-73. HFC-227ea test. Refer to Table B-4 for experimental conditions.
Figure B-74. HFC-227ea test. Refer to Table B-4 for experimental conditions.
Figure B-75. HFC-227ea test. Refer to Table B-4 for experimental conditions.
Figure B-76. HFC-227ea test. Refer to Table B-4 for experimental conditions.
750

3

2.5

Figure B-77. CF3I test. Refer to Table B-5 for experimental conditions.
OPTIMIZATION OF SYSTEM DISCHARGE 751

2.5

1? 2

Q.

0)1.5

eno

q: 1

0.5

T—I—I—I—

I—I—I—I—I—I—I—I—I—I—

I—I—I—I—I—I—I—I—I—I—

I—1—I—I—

r

I I I I I

I I I uJ I I—I—I_I—I—I—I—I—I—J—I—I—I—J—1—1—I—

L

1000 2000 3000 4000 5000

Time (ms)

6000

Figure B-78. CF3I test. Refer to Table B-5 for experimental conditions.
752 8. OPTIMIZATION OF SYSTEM DISCHARGE

4 pT—I—I—I—

I—I—I—I—'—

I—I—I—I—I—

!—I—I—I—I—I—I—I—I—I—

I—I—I—I—I

3.5

5:2.5 t

fl) 2 t

3

(0

Sl.5 t

1

0.5

r

i

r

J—I—I—I—I—I—I—I—I—I—I I I I I—I—I—I—I—I—I—I—I I—I I I I I

0 1000 2000 3000 4000 5000 6000

Time (ms)

Figure B-79. CF3I test. Refer to Table B-5 for experimental conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 753

Figure B-80. CF3I test. Refer to Table B-5 for experimental conditions.
Figure B-81. CF3I test. Refer to Table B-5 for experimental conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 755

Figure B-82. CF3I test. Refer to Table B-5 for experimental conditions.
Figure B-83. CF3I test. Refer to Table B-5 for experimental conditions.
Figure B-84. CF3I test. Refer to Table B-5 for experimental conditions.
Figure B-85. CF3I test. Refer to Table B-5 for experimental conditions.
Figure B-86. CF3I test. Refer to Table B-5 for experimental conditions.
Figure B-87. CF3I test. Refer to Table B-5 for experimental conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 761
762 8. OPTIMIZATION OF SYSTEM DISCHARGE

2 r—i—I—I—I—

j—I—I—i—I—

I—I—I—I—I—

j—I—1—I—

r

2.5

(0 2

3

0)

q: 1

0.5

T—

I—I—I—I—

r

j

ill

III

J—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—

L

0 500 1000 1500 2000 2500 3000

Time (ms)

Figure B-89. CF3I test. Refer to Table B-5 for experimental conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 763

Figure B-90. CF3I test. Refer to Table B-5 for experimental conditions.
764 8. OPTIMIZATION OF SYSTEM DISCHARGE

Figure B-91. CF3I test. Refer to Table B-5 for experimental conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 765

Figure B-92. CF3I test. Refer to Table B-5 for experimental conditions.
Figure B-93. CF3I test. Refer to Table B-5 for experimental conditions.
OPTIMIZATION OF SYSTEM DISCHARGE 767

Figure B-94. CF3I test. Refer to Table B-5 for experimental conditions.
Figure B-95. CF3I test. Refer to Table B-5 for experimental conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 769

Figure B-96. CF3I test. Refer to Table B-5 for experimental conditions.
Figure B-97. CF3I test. Refer to Table B-5 for experimental conditions.
OPTIMIZATION OF SYSTEM DISCHARGE

4 r '

I I' I

'

I I I j

I I I I

I

I I I I I

I I I I I

I I I I

3.5 :

3 :

- I I I I I I I I I —I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—I

0 500 1000 1500 2000 2500 30

Time (ms)

Figure B-98. CF3I test. Refer to Table B-5 for experimental conditions.
Figure B-99. CF3I test. Refer to Table B-5 for experimental conditions.
8. OPTIMIZATION OF SYSTEM DISCHARGE 773

Appendix C. Thermodynamic and Transport Properties of Agent/Nitrogen

Mixtures

Coefficients to polynomial fits of calculated thermodynamic and transport properties of agent/nitrogen

mixtures during isenthalpic expansion are tabulated in Tables C-1 and C-2. Figures C-1 to C-6 show

the polynomial fits to the density, temperature and mass fraction results at initial pressures of 2.75 and

4.12 MPa. Figure C-1 shows the pressure/density relationships for isentropic expansion of gas-phase

agent/nitrogen mixture initially at 20 °C. The curves through the data represent power-law fits to the

points. In general, the exponent was not a function of initial pressure, thus the pseudo-ideal y was

taken as a constant for each agent. The values obtained for halon 1301, HFC-125, HFC-227ea, and

CF3I were 1.07, 1.12, 1.01, and 0.99 respectively.
774 8. OPTIMIZATION OF SYSTEM DISCHARGE

Table C- 1 . Mean agent/nitrogen mixture densities

Agent

(MPa)

(kg/m^) = a^, + P in

a,*P + a2*P^ + a3*P^ kPa

^0

CFjBr 2.00

2.50

2.75

3.00

4.12

-7.815

14.77

18.49

16.36

3.334

231.5

79.10

66.39

68.25

118.4

-46.81

184.6

189.5

180.4

81.99

160.8

11.39

-5.087

-12.23

-5.742

HFC- 125 2.00

2.50

2.75

3.00

4.12

5.49

6.87

8.24

15.55

17.09

14.32

9.038

3.586

-0.9375

-5.628

-7.401

29.97

5.372

21.18

47.35

76.32

88.697

92.75

85.04

185.7

226.7

198.0

153.9

81.13

40.26

18.41

7.405

47.50

-17.28

-19.27

-14.11

-8.235

-3.845

-1.511

-0.5212

HFC-227ea 2.00

2.50

2.75

3.00

4.12

5.49

6.87

8.24

-46.28

-50.25

-49.05

-56.03

-50.43

-44.59

-35.94

-30.56

375.9

407.1

402.7

418.8

362.2

297.1

241.5

193.6

478.8

222.5

174.3

98.02

28.29

3.781

-1.419

-1.828

-159.2

-65.64

-50.13

-28.24

-8.473

-2.334

-0.8257

-0.3262

CF3I 2.00

2.50

2.75

3.00

4.12

-52.59

-73.51

-80.37

-98.79

-87.52

467.7

613.6

648.9

690.8

607.0

893.7

402.7

269.3

151.0

22.44

-303.6

-124.6

-82.36

-49.62

-11.49
8. OPTIMIZATION OF SYSTEM DISCHARGE

Table C-2. Mean agent/nitrogen mixture viscosities

775

Agent pinitial (MPa) @ 20 °C(g/cm s) = + a,*/' + dLj*P^ + aj*/*^ P in MPa

So ^2

Halon 1301 2.75 1.456 X 10"^ 7.500 X

10-5 1.461 X 10"" 6.043 X 10-^

4.12 1.294 X lO"'' 1.192 X

10"'* 5.682 X

10-5 6.407 X10"^

HFC- 125 2.75 1.238 X

10"^ 6.513 X

10-5 1.865 X

10-"* -1.171 X

10-5

4.12 1.111 X

10"^ 1.194 X 10"* 6.457 X

10-5 -3.316 X 10-^

HFC-227ea 2.75 2.783 X

10-5 7.386 X

10""* 2.947 X

10-" -8.376 X10-5

4.12 3.707 X

10"' 6.770 X 10~* 4.709 X

10-5 -1.312 X

10-5
776 8. OPTIMIZATION OF SYSTEM DISCHARGE

2000 I

' ' I '

I

' ' I

'

I '

' ' '

I

'

' ' '

I '

' ' ' M ' '

'

E 1500

'55c

oO

c

0)

1000

500

- o- -halon 1301

--B--HFC-125

o - - HFC-227ea

-CFJ

o' /

/

J I I I I I I L J L I I I I I I

0.5 1 1.5 2 2.5

Pressure (MPa)

Figure C- 1 . Mean density as a function of pressure for initially liquid-phase agent/nitrogen

mixture undergoing isenthalpic expansion. Initial pressure and temperature are 2.75 MPa and 20 °C.
8. OPTIMIZATION OF SYSTEM DISCHARGE 777

O

3

(0

0)

a

E

I300

280

£ 260

240

220

200

I I I

I I I I

I I I 1 II

I I I I

I 1 I I I

6

0

' /

/

<?>'

9

/

/9

/

I

, I P

' L_l L

- e- - halon 1301

----HFC-125

- o - -HFC-227ea

^--CF3l

I . . I L

0.5 1 1.5 2 2.5

Pressure (MPa)

Figure C-2. Temperature as a function of pressure for initially liquid-phase agent/nitrogen

mixture undergoing isenthalpic expansion. Initial pressure and temperature are 2.75 MPa and 20 °C.
778 8. OPTIMIZATION OF SYSTEM DISCHARGE

T—III!—I—I—TT

o

(0

c

o

o

(0

(0

(0

0.9

0.8

0.7

0.6

0.5

:/

I

.0

- e- - halon 1301

----HFC-125

o - HFC-227ea

0.4

j-J u-i I I I—J I—I i-J i_i I I L_j I uj L-i. X_J L

0.5 1 1.5 2 2.5

Pressure (MPa)

Figure C-3. Liquid-phase mass fraction as a function of pressure for initially liquid-phase

agent/nitrogen mixture undergoing isenthalpic expansion. Initial pressure and

temperature are 2.75 MPa and 20 °C.
8. OPTIMIZATION OF SYSTEM DISCHARGE 779

CO

c

Q

(0

2000 ' ' ' '

I

' ' ' ' I' ' " '

I ' ' ' ' ^l- '

E 1500

1000

500

- e- - halon 1301

----HFC-125

o - - HFC-227ea

A

^- - CF3I

/ O B

O 0

J L J I I I 1—1 L a I L

1 2 a

Pressure (MPa)

Figure C-4. Mean density as a function of pressure for initially liquid-phase agent/nitrogen

mixture undergoing isenthalpic expansion. Initial pressure and temperature are 4.12 MPa and 20 °C.
780 8. OPTIMIZATION OF SYSTEM DISCHARGE

300

280

J> 260

240

220

200

T—r—I—I—I—I—!—I—I—I—I—I—

r

1—

r

JD

/

. ' /

/o

0- -halon 1301

--HFC-125

o - - HFC-227ea

J I I I I I I I I I I I I I L 12 3 4

Pressure (MPa)

Figure C-5. Temperature as a function of pressure for initially liquid-phase agent/nitrogen

mixture undergoing isenthalpic expansion. Initial pressure and temperature are 4.12 MPa and 20 °C.
OPTIMIZATION OF SYSTEM DISCHARGE

o

3

(0

(0

0.9

0.8

c

o

o

S 0.6

0.5

T—I—I—

1—I—r-r

0.7

-

L /

-6.'

- ©- - halon 1301

----HFC-125

o -HFC-227ea

0.4

' 1 J I I I I I I L 1 J L 12 3 4

Pressure (MPa)

Figure C-6. Liquid-phase mass fraction as a function of pressure for initially liquid-phase

agentynitrogen mixture undergoing isenthalpic expansion. Initial pressure and

temperature are 4.12 MPa and 20 °C.
782 8. OPTIMIZATION OF SYSTEM DISCHARGE

200 T—

f—1—1—r—1—j—I I I I—

I—I—I—I—I—I I I—I—

r

CO£ 150

"55 o

CO

0- -halon 1301

--HFC-125

o - HFC-227ea

-A- - CFJ

9

0'

50

0

C3

AJ I I i I I I I i I I I L. J_J I L

1 2 3 4 5

Pressure (MPa)

Figure C-7. Mean density as a function of pressure for initially gas-phase agent/nitrogen mixture

undergoing isentropic expansion. Initial pressure and temperature are 4.12 MPa and

20 °C.

ft U.S. GOVERNMENT PRINTING OFFICE: 1995 - 400 - 019



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