Chapter 4: AEROSOL PROPERTIES James W. Fleming, Ph.D. and 

Ronald S. Sheinson, Ph.D., 

U.S. Naval Research Laboratory 

TABLE OF CONTENTS

4.1 Introduction......................................................................................................................340 

4.2 Previous Knowledge ........................................................................................................340 

4.2.1 Water .....................................................................................................................340 

4.2.2 Powders.................................................................................................................341 

4.2.3 Combined Agents - Water Additives ....................................................................342 

4.2.4 Burning Surfaces...................................................................................................344 

4.2.5 Large-scale Studies................................................................................................345 

4.2.6 Modeling ...............................................................................................................348 

4.3 The Nature of Aerosols....................................................................................................348 

4.3.1 Quantitative Descriptors........................................................................................348 

4.3.2 Generation Techniques..........................................................................................350 

4.3.3 Quantification of Aerosols ....................................................................................351 

4.3.4 Aerosol Size and Velocity Characterization .........................................................351 

4.3.5 Aerosol Properties Affecting Behavior.................................................................355 

4.4 The Nature of Flames.......................................................................................................358 

4.4.1 Flame Characteristics............................................................................................358 

4.4.2 Premixed Bunsen Type Flames.............................................................................359 

4.4.3 Non-premixed Flames...........................................................................................360 

4.5 Aerosol Flame Suppression Studies.................................................................................363 

4.5.1 Water, a Physical Suppression Agent....................................................................363 

4.5.2 Alkali Metal Bicarbonates, Chemically Active Fire Suppressants .......................397 

4.5.3 Aqueous Metal Salt Solutions, Combined Chemical and Physical Suppression ..412 

4.5.4 Dendrimers, Combined Chemical and Physical Flame Suppression ....................430 

4.6 Surface Cooling ...............................................................................................................438 

4.6.1 Introduction ...........................................................................................................438 

4.6.2 Evaporation of Liquid Droplets.............................................................................440 

4.6.3 Spray Cooling........................................................................................................443 

4.6.4 Convective Regime ...............................................................................................444 

4.6.5 Nucleate Boiling Regime ......................................................................................445 

4.6.6 Critical Heat Flux ..................................................................................................446 

4.6.7 Transition and Film Boiling Regimes...................................................................446 

4.6.8 Summary ...............................................................................................................448 

4.7 Aerosols and Burning Surfaces........................................................................................448 

4.7.1 Flame Spread Studies............................................................................................449 

4.7.2 Droplet Interaction with a Hot Surface .................................................................458 
340 Aerosol Properties 

4.7.3 Summary and Conclusions....................................................................................466 

4.8 The New State of Knowledge of the Mechanism of Flame Extinguishment 

by Aerosols......................................................................................................................467 

4.9 Relation of Findings to Practical Fires.............................................................................468 

4.10 References........................................................................................................................469 

4.1 INTRODUCTION 

From the beginning of the Next Generation Fire Suppression Technology Program (NGP), it was 

expected there would be condensed phase compounds that might offer sufficiently efficient fire protection 

that could be considered as replacements for halon 1301. These compounds would be delivered to the fire 

as aerosols. An aerosol is a suspension of small diameter solid particles and/or liquid droplets in a gas. 

Fog and dust clouds are examples of liquid and solid aerosols, respectively. The term particle is often 

used to refer to the individual solid components or the individual liquid drops that make up the aerosol. 

The realm of fire suppressant aerosols includes two-phase fluid combinations of air with water, 

halogenated organic chemicals, or solids such as alkali metal compounds or transition metal salts. 

Because of the influence of gravity, the size of the individual liquid droplets or solid particles must be 

sufficiently small in order to be carried or suspended by a finite flow of air. Unlike gaseous fire 

suppression agents, aerosol agents must rely on their environment and delivery method in order to remain 

suspended and reach the fire location. 

Complementing the understanding of the chemical action of fire suppression compounds as described in 

Chapter 3 with an understanding of the role of the colligative properties of aerosols was recognized as a 

central factor to the search for suitable replacements. However, the published data describing the fire 

suppression behavior of aerosols were often difficult to interpret, especially for larger-scale fire tests. It 

was difficult to separate possible chemical suppression effects from the physical effects which are almost 

always present. It was known that the aerosol particle size was an important parameter in suppression 

effectiveness although a quantitative understanding as to the exact role particle size played was hard to 

determine. In short, the phenomenon of fire suppression using aerosols needed further exploration. 

The NGP research described in this chapter was undertaken to address the lack of quantitative 

information on aerosol suppression behavior. Studies targeted the effects of aerosol size, aerosol droplet 

or particle velocity behavior in various flow fields, the physical and chemical properties of liquid and 

powder aerosols, and the predictive capabilities of aerosol suppression behavior. Most of the research 

predates the NGP focus on aircraft fire protection (Chapter 1) and is directed at a wider range of halon 

replacement applications. 

4.2 PREVIOUS KNOWLEDGE 

4.2.1 Water 

Certainly the fire fighting agent that has been used the longest is water (Chapter 1). The interactions of 

water droplets of various diameters with flames have been studied for almost four decades.1, , , , 2 3 4 5

 Many of 

the early studies examined the effectiveness of water sprays on liquid fires. The suppression of liquid fuel 
Previous Knowledge 341

counterflow flames by water was experimentally investigated by Seshadri6 using very small droplets that 

vaporized completely. McCaffrey7

 in the 1980s examined the effect of water introduced in the form of a 

mist of small droplets on natural gas diffusion flames of various sizes. He introduced the mist outside the 

base of the flame, where it was entrained along with the fuel. Mass flows of water mist approximately 

equal to the fuel flows extinguished the flame. In this instance, fine water mist was as effective as halon 

1301. In 1977, the U.S. Naval Research Laboratory (NRL) examined some of the advantages of water 

mist for shipboard use.8

 In addition to pointing out the physical roles (oxygen dilution and high molar 

heat absorption), the report also noted that aqueous solution of salts would form fine solid particles upon 

evaporation of the water. 

Although there was consensus in the community on the fundamental fire suppression mechanism of 

water, little detailed quantitative information on the various physical effects operating or any possible 

chemical effects of water mist was available. Such information underpins the ability to identify when a 

water-based suppression should be considered, how best to deliver the water, and how one might enhance 

the fire suppression ability of water mist. Aerosol fire suppression understanding was in its infancy at the 

beginning of the NGP. 

4.2.2 Powders 

In addition to the liquid aerosol properties that affect suppression (heat capacity, vaporization and 

possible additional chemical suppression pathways) powder aerosols possess additional properties that 

impact fire extinguishment including crystal decrepitation, decomposition, and surface mediated 

reaction.9

 In interpreting fire suppression results for early studies with various dry powders,10, , 11 12

researchers were able to utilize the surface area of the particles to explain thermal vs. chemical 

effectiveness of finely divided particles. 

Rosser et al.13 examined the effect of finely divided metal salts on the burning velocity of premixed 

hydrocarbon-air flames in the early 1960s. While several of the powders they studied were not effective, 

several compounds of sodium, potassium, and copper were effective. They observed that the measured 

inhibition correlated with the degree of particle evaporation. Successive steps were identified for 

successful suppression: particle heating, particle evaporation, particle decomposition, and chemical flame 

inhibition by the added metal compounds. 

Fischer et al.14 at NRL in the early 1980s studied methane flame extinction in a laboratory scale Bunsen 

burner flame apparatus for a number of powders including monoammonium phosphate, potassium 

bicarbonate, potassium chloride, sodium bicarbonate, and MonnexTM (primarily K2C2N2H3O3). 

Extinguishment efficiency increased rapidly with decreasing powder particle diameter with transition 

diameters ranging from 40 μm to 60 μm depending on the specific inorganic salt compound. Of the 

powders studied, MonnexTM displayed the largest size, ~ 60 μm, for onset of higher efficiency (probably 

related to the tendency of MonnexTM to undergo decrepitation). 

In the early 1980s, Ewing, Hughes, and Carhart15 at NRL developed an empirical predictor for the flame 

suppression effectiveness of liquid, gaseous, and solid agents. They observed a large increase in 

suppression efficiency with decreasing particle size. They concluded that suppression of a methane fire 

was achieved predominantly by heat extraction from the flame by means of heat capacity sinks and 

endothermic reaction sinks, such as vaporization, dissociation, and decomposition for the powders they 
342 Aerosol Properties 

investigated. They generalized that the extinguishing potential of a powder agent could be predicted from 

its thermodynamic properties. In later tests, Ewing et al.16 extinguished heptane pan fires in a 420 L test 

chamber. Both NaHCO3 and KHCO3 agents were applied to the fires via a pressurized discharge nozzle. 

From the tests, a critical size below which particles fully decomposed in their fire scenario was 

determined for each agent. In a companion effort, Fischer and Leonard17 performed extinction studies on 

21 agents in a premixed flame and also observed a distinct transition region from effective to ineffective 

particle sizes. Similar experiments were performed by Chattaway et al.18 in evaluating the efficiency of 

very small aerosol particulates. 

Knuth, Ni, and Seeger19 studied the flame inhibition effectiveness of several dry chemicals, including 

KHCO3 and NaHCO3, solely in the 38μm to 43 μm size range, by adding powders to a counterflow 

diffusion flame and measuring species and temperature profiles. Their conclusions were primarily based 

on the decrease in flame temperature and production of the major combustion product species, H2O and 

CO2. The powders tested were then ranked in increasing reaction-inhibition effectiveness: Al2O3 (least 

effective), NaHCO3, KHCO3, NH4H2PO4, and KCl (most effective). Hamins et al.20 used a counterflow 

burner to study extinction of liquid pool fires of n-heptane and JP-8 by 13 gaseous agents and sodium 

bicarbonate. Powder sample in three nominal size ranges were tested: 0 μm to 10 μm, 10 μm to 20μm, 

and 20 μm to 30 μm. Results were not reported for powders above 30 μm. They concluded that the 0 μm 

to 10 μm size range was the most effective, followed by the 20 μm to 30 μm and then the 10 μm to 20 μm 

range. No explanation for the non-monotonic behavior of effectiveness with size was proposed. Using a 

cup burner, Hamins showed that, on a mass basis, small NaHCO3 particles (sub-10-μm diameter particles) 

which act both as a physical and a chemical suppressant, are three times more effective than halon 1301 

(CF3Br) and five times more effective than nitrogen.21

4.2.3 Combined Agents - Water Additives 

A brief, general discussion on the uses of water additives for fire fighting can be found in Wahl.22 The 

general intent of adding salts to water was to improve the thermophysical properties of neat water. An 

example is the lowering of the normal freezing point of water so that it can be used at sub-zero 

temperatures. For fire suppression, it had long been recognized that additives could significantly reduce 

the mass of water needed to quench flames. 

Monson et al.23,24 investigated various water additives for fire fighting. Several formulations were 

identified, which included the following aqueous solutions: 24 % by mass of LiCl, 5 % LiCl, 26 % CaCl2, 

10 % Ca(NO3)2, 27 % CaCl2, and 5 % FeCl3, 29 % CaCl2. Only the first of these formulations was 

actually tested for fire suppression effectiveness. Using a spray nozzle directed at a wood crib fire, the 

solution extinguished the fire 1.4 times faster than neat water applied at the same rate. 

In a series of reports, Grove and co-workers25, , , , , , , , 26 27 28 29 30 31 32 33 evaluated the effect of the addition of a 

viscosity modifier, a surfactant, and/or an opacifier to water on improving the fire suppression 

effectiveness of water (applied in the form of a spray). The purpose of using a viscosity modifier was to 

improve the blanketing and runoff properties of water. The opacifier additives were used to mitigate the 

radiative feedback from the fire to the burning fuel surface. The surfactant was used to facilitate the 

dispersion of opacifiers in water. Based on the experimental results of their scaled model Class A (solid 

fuel) fires, it was concluded that: 
Previous Knowledge 343

• The major contribution to the improvement of the fire suppression effectiveness was 

attributed to the viscosity modifier. 

• Viscous water produced more rapid initial control of the fires than neat water. 

• The rate of extinguishment of the fires was more rapid with viscous water; the 

extinguishment time decreased as the viscosity of water was increased. 

• The danger of re-ignition was reduced when the fires were extinguished with viscous water, 

as a result of less water runoff. 

• Reduced runoff of viscous water provided better water utilization for fire control and 

minimized collateral water damage. 

• The addition of an opacifier to viscous water further improved the fire fighting properties of 

water. 

• The surfactants used in the study had very little effect on the improvement of the 

extinguishing properties of water. 

In an attempt to improve the effectiveness of water to control and suppress forest fires, Davis et al.34,35

added several viscosity modifiers to plain water. The results from their controlled field and laboratory 

fire tests also indicated that fire suppression effectiveness was related to the viscosity of water. 

Fine spray solutions of sodium and potassium carbonates and bicarbonates were studied by Friedrich 36 to 

evaluate the effectiveness of these aqueous solutions in suppressing liquid pool fires. Concentrations 

ranging from a mass fraction of 0.01 to 0.1 were used in the study, which involved the application of the 

solution from a spray nozzle located at a fixed distance from the liquid pool. The salt solutions with the 

highest concentration exhibited the highest effectiveness; however, no improvement in the effectiveness 

was observed when the concentration was less than a mass fraction of 0.05. No comparison of the fire 

suppression effectiveness among various salt solutions was given in the cited reference. The author 

attributed the increase in fire suppression effectiveness of these aqueous solutions to the chemical 

inhibition ability of the salts and suggested that other compounds such as the oxalates would probably be 

more effective than the carbonates (albeit low solubility in water and toxicity are potential issues). 

Investigation of very high mass fraction additives to water was motivated by the need for fire suppression 

at low theater temperatures. At the Army Research Laboratory, Elkins studied the use of ethylene glycol 

anti-freeze as an additive to water.37 At a volume fraction of 0.3, the ethylene glycol/water solution was 

found to be approximately three to four times less effective (in terms of extinguishment time) than pure 

water. The author recommended that the concentration of ethylene glycol should not exceed a volume 

fraction of 0.3, lest the solution lose its fire fighting capability. 

The fire suppression effectiveness of aqueous solutions of KBr, KCl, NH4Cl, Na2CO3, NaHCO3, K2CO3, 

KHCO3, and HCOOK of various concentrations were examined by Kida.38 The solutions were sprayed 

vertically downward to a small hexane pool fire, and the extinction times were measured. Despite large 

scatter in the data, a 0.20 volume fraction KHCO3 solution was found, on the average, to be at least twice 

as effective (in terms of the extinction times) as pure water. The performance of the other aqueous 

solutions was not reported in the cited reference. 
344 Aerosol Properties 

In a United Kingdom Patent by Ball et al.,39 potential water additives such as phosphates, carbonates, 

amino compounds, citrates, anti-freeze agents, and surfactants were suggested to increase the solubility of 

CO2 in water for fire suppression applications. 

Finnerty et al.40 evaluated the fire suppression effectiveness of 13 water additives using an airless paint 

sprayer and a small JP-8 pan fire. Solutions of potassium lactate (0.60 mass fraction) and potassium 

acetate (also at 0.60 mass fraction) were found to be superior to the other additives and at least four times 

more effective than pure water in terms of extinguishment time. Although these two solutions had the 

same overall fire suppression efficiency, the acetate has a sufficiently higher normal freezing point as to 

render it undesirable for low temperature applications. The superior performance of the potassium lactate 

and acetate solutions was further demonstrated using the dispersed liquid agent fire screen apparatus 

developed by NIST41 under the NGP (Chapter 6). 

The significant enhancement in flame suppression ability of aqueous NaOH solutions (almost a factor of 

five for 17.5 % of NaOH by mass in water)40 was found to be consistent with previous experiments 

reported by Zheng et al.42 for premixed counterflow flames using aqueous NaCl solutions. 

Experiments by Mitani and Niioka 43 considered inhibition of premixed flames with ultra fine water 

droplets (less than 2.4 μm mean diameter) containing NaOH and NaHCO3, indicating a flame inhibition/ 

extinction with addition of chemical agents. In particular, their slower C2H4/O2/N2 premixed flame 

experiments (as opposed to faster H2/O2/N2 flames) showed distinct chemical inhibition effects, but the 

saturation effects due to condensation of the inhibitor species were not explored. Zheng et al.42

documented a significant enhancement in counterflow premixed flame suppression from fine droplets of 

solutions of water/ NaCl over that of pure water. 

Fire suppression studies relevant for protecting shipboard machinery spaces conducted by the British 

Royal Navy in Great Britain examined the suppression effect of adding proprietary wetting agents, 

Aqueous Film Forming Foam (AFFF), and Film Forming Fluoro Proteins (FFFP) to water sprays.44 They 

found that the additives did shorten the fire extinguishing times over neat water, yet the surface tension 

lowering actually resulted in a slight increase in the mean droplet size of the resulting water spray. In this 

case, the foam forming attributes and possible chemical suppression properties of the additives likely 

played a more significant role in the suppression effectiveness than the influence of the additive on the 

surface tension and subsequent droplet formation of the liquid. 

4.2.4 Burning Surfaces 

There have been numerous studies on flame spread velocities in opposed and counter flow configurations 

during the last three decades.45, , , , , , 46 47 48 49 50 51 These studies have shown that the flame spread velocity 

depends on the fuel thickness, flow configuration, oxygen concentration, and the size of the wind tunnel. 

Therefore, for flame suppression studies, it is best to obtain data with and without the suppressant in the 

same facility. 

Studies describing spray-target interactions have primarily focused on application of paint coatings to 

surfaces.52,53 These studies have shown that high gas phase strain rates where the spray impinges on the 

target dramatically affect turbulent transport of small droplets (diameter less than 25 μm), and hence the 

deposition of the droplets on the intended surface. In addition, the deposited droplet size distribution can 

be significantly different from that of the initial spray. Finally, droplet transport is dependent on initial 
Previous Knowledge 345

droplet momentum, aerodynamic drag, and the gravitational force, with aerodynamic drag being 

dominant. 

Many studies have been performed that quantify the vaporization process of both single and multiple 

droplet arrays impacting on hot surfaces.54,55 These studies have identified the need for spatially and 

temporally resolved data describing the droplet trajectory as it approaches a hot surface (including 

buoyancy effects), as well as the temperature distribution of the surface material. The NGP studies 

addressing the interaction of droplets with surfaces are described in Section 4.6 and 4.7. 

4.2.5 Large-scale Studies 

Need for Testing at Realistic Scale 

A candidate halon replacement (in-kind, i.e., vaporizing liquid) or alternative (not in-kind, e.g., aerosol, 

such as water mist or fine powder) may demonstrate great promise based on laboratory or small-scale 

evaluations. Implementation effectiveness will hinge on a number of additional considerations including 

applicability, compatibility, stability, shelf-life, toxicity, and dispersal. Performance against a full-scale, 

realistic threat and long-term certification is often needed to provide confidence that what looked good in 

the laboratory will work reliably in practice. 

In practice, gaseous fire suppressants deployed as total flooding agents are generally expected and 

required to disperse throughout a space. Even so, there will remain distribution inhomogeneities due to 

finite agent release locations and the effects of obstacles. NRL studies in a real-size simulated ship 

machinery space (840 m3

) found that, even for a very carefully designed gaseous agent discharge system, 

there were agent concentration inhomogeneities of at least ± 20 % that impacted the fire extinguishment 

time.56 For fire suppression agents that are solids or liquids at the application temperature, dissemination 

and distribution challenges become much more difficult than for gaseous agents. 

Liquid Aerosols - Water Mist 

Full-scale total flooding water mist extinguishment tests, including mists with additives such as AFFF, 

were conducted by NRL in the early 1980s in 83 m3 and 324 m3 highly instrumented compartments.57 

The testing demonstrated the effectiveness of a fresh water mist with a high fraction of droplet diameters 

under 50 μm in diameter. Fixed arrays of commercial nozzles, mainly Bete Fog Nozzles, were evaluated 

at discharge pressures up to 2 MPa (300 psi). Suppression challenges were included in a variety of Class 

A (solid) and Class B (liquid) fuels in compartments configured as very obstructed and cluttered 

machinery spaces or torpedo rooms. Except for occasional malfunctions, all Class B fires were 

extinguished in 15 s to 40 s. Some Class A fires in the 324 m3

 compartment were fully extinguished with 

one or two bursts of water mist of 1 min to 2 min duration. The flames from other Class A fires were 

extinguished, but smoldering continued. Flaming combustion reappeared after the water mist system was 

turned off. Even with slow extinguishment, significant cooling was achieved; the fires were “slowed 

down” and re-entry by protected firefighters would have been possible. Even for the largest tests, the 

total quantities of water used for extinguishment were usually less than 190 L to 265 L (50 gal to 70 gal). 

For comparison, a conventional 38 mm (11⁄2 in) fire hose delivers about 380 L/min (100 gal/min), and a 

single hose would hardly begin to quench these fires, especially in the obstructed geometries. With the 
346 Aerosol Properties 

water mist, cleanup was minimal. The experience showed the great merit of water mist for total flooding 

applications and the potential for optimization. 

Solid Aerosols - Powders 

Early NRL studies in the 1950s provided quantitative data on the effectiveness advantage of potassium 

bicarbonate (KHCO3) over sodium bicarbonate (NaHCO3), and addressed practical techniques for 

allowing better implementation of powders. In large-scale powder studies in the 1980s, Ewing et al. at 

NRL found that the increased efficiency with decreasing powder particle size observed in small, bench 

top experiments was also valid when extinguishing large heptane pool fires (0.29 m2

 and 2.32 m2).58,59 

Momentum factors for getting the particles into the flame sheet were considered. 

Pyrogenic Generated Aerosols 

Pyrogenic generation of fire suppression agents is one method of generating and dispersing aerosol agents 

in large-scale. The “Halon 90 - Second Conference on the Fire Protecting Halons and the Environment” 

in Geneva on 1-3 October 1990 presented Russian fine solid aerosol fire suppressant efforts via videos of 

pyrotechnic devices in operation. These combustion devices released a very hot product stream that 

produced very small solid aerosol particles by condensation. Also presented were a number of Russian 

pyrotechnic aerosol efforts, including work by Russian émigrés and further studies in Israel and the USA. 

Initial U.S. Navy evaluation at the Spectronix Israeli site in August 1992 confirmed that the buoyancy of 

the exothermically produced aerosol is a significant distribution issue in employing a heterogeneous 

fuel/oxidant propellant product as a total flooding agent in large spaces. 

Follow-up experiments at NRL measured aerosol generation temperatures and pressures and evaluated 

two approaches to removing heat: passive water bath and water spray generators. Fire threats included 

n-heptane pool fires of various sizes. In the 56 m3

 (2000 ft3), 4.3 m (14 ft) tall instrumented compartment 

tests,60 aerosol concentrations were determined from optical density measurements at three heights: 0.5 m, 

2.0 m, and 4.1 m (1.7 ft, 6.7 ft, and 13.5 ft). 

In tests of uncooled aerosol generation, 1/4 to 1/3 of the solid agent containing pyrotechnic material was 

aerosolized with over 200 L of CO2 per kg of original material. There was also considerable energy 

emitted. Visible flame plumes 30 cm long exited the generator horizontally, but then quickly rose almost 

vertically due to buoyancy effects. The highest and middle optical probes were quickly engulfed using 

1 kg of material. Only after almost 90 s did the aerosol cloud descend to within 0.5 m of the floor. The 

aerosol densities took over 4 min to equilibrate and maintain these levels at the three heights. There was 

little visible residue.61

Rapidly bubbling the hot aerosol through water for cooling produced both a rising buoyant plume and a 

dropping cooled cloud. The two layers did not coalesce until after 2 min, and the aerosol densities were 

lower with water removal of the aerosol. 

The generators utilizing water spray cooling dispensed agent horizontally for over 1 m. However, the 

1 kg aerosol charge was not cooled sufficiently by 1.6 L water injection and exited the generator at close 

to 1000 oC. The horizontal momentum thrust dissipated and buoyancy forces still dominated dispersion. 

Thus, the elevated temperatures of pyrotechnic-produced aerosols remained a concern. 
Previous Knowledge 347

Large-scale fire extinguishment tests evaluated both liquid and solid aerosols. Hot gas generators were 

found to extinguish 0.23 m2

 pan fires in 45 s with 4 kg of pyrotechnic material, equivalent to 71 g of solid 

fire extinguishing agent per m3

. Water bath cooled generators could not extinguish the fires with 3 kg of 

material; extinguishment required 5 kg and more than 115 s (equivalent to 89 g/m3

). In contrast, the 

water spray cooled generators extinguished a 0.23 m2

 pool fire in 65 s using just 3 kg of agent (equivalent 

to 54 g/m3

). 

Larger, 1.1 m2

 pan fires were extinguished more quickly and required powder concentrations as low as 

15 g/m3

. However, the larger pan fires were partially self-inhibited by oxygen depletion and were thus 

easier to extinguish. Even then, aerosol generation and fire extinguishment took significant lengths of 

time. Much faster generator mixture burning rates are possible with other compositions. However, the 

increased energy release rates can exceed pressure and temperature safety level requirements, requiring 

greatly improved aerosol cooling mechanisms to obtain uniform distributions. 

Researchers did note that dry chemical fire extinguishing powders exhibited suppression vs. size behavior 

similar to water aerosols. For a fixed mass of extinguishant, larger particles were found to be less 

effective and were observed to fall out faster than smaller ones. Very small pyrotechnically generated 

particles, with estimated diameters ranging from 0.5 μm to 4 μm, propelled at a fire, should be easily 

swept away by fire-generated convection currents. The small particle size and high concentration aerosol 

cloud should also cause concern in occupied enclosures. Particles less than 5 μm in diameter can 

penetrate deep into the lung and cause pulmonary complications, physically clogging the small pulmonary 

air paths. Some aerosol formulations that have been pursued, while possibly more efficient as fire 

extinguishants, may generate more toxic and abrasive products. A U.S. Navy toxicological evaluation of 

a pyrotechnic fire extinguishing formulation with rats showed short-term breathing difficulties although 

no residual health effects were observed. For the protected area, the hygroscopic nature of some aerosol 

residues may lead to corrosion and electrical problems. 

Addional pyrotechnic approaches were pursued in England, the U.S., and the former Soviet Union. 

Detailed results were often not available due to the proprietary nature of the materials tested. Solid 

energy absorbers, evaluated by various producers, increased weight penalties. "VNIIPO" (Moscow) 

studied nine different types of pyrotechnic-generated aerosol units from "GABAR," "GRANIT," 

"PERM," "SOYUS," and "Technology," in October 1993. Tests were conducted on 5A and 13B 

challenges in a 64 m3

 volume. Only the two GABAR single units succeeded in extinguishing both test 

fires and preventing reignition.62 Dr. Baratov claimed his GABAR company units addressed the concerns 

of ignition of combustibles by hot propellant exhaust, cooled aerosol generation for better total volume 

distribution, and effectiveness on smoldering combustion. One U.S. Army test series on tank engine 

compartment fires found a pyrotechnic agent successfully extinguished the fire only to later reignite it. 

The weights of the heat abstracting units can be considerably more than the pyrotechnic charge weights. 

Still, total system weight was low compared to the pressure vessels and piping of halon-like systems. 

Achieving homogeneous distribution can be a significant implementation problem for pyrogenic 

approaches in large spaces. However, such fire suppression strategies may be quite satisfactory for 

smaller spaces with lower ceiling heights, such as small rooms, cabinets or storage closets. 
348 Aerosol Properties 

4.2.6 Modeling 

For most of the 20th century, there were few basic models that described details of aerosol/flame 

interactions for non-combustible aerosols. Prior to the NGP, there were no detailed models describing the 

size reduction and evaporation of droplets in flames, or their effect on surface reaction kinetics. Even 

models that predict what size droplets could reach burning surfaces were not generally available. 

Understandably, there was interest in understanding the overall performance of larger-scale water mist 

systems.42,63 In 1955, Braidech et al.1

 concluded that, in most cases, the efficacy of water mist is 

predominantly due to the oxygen displacement effect of the water evaporating, rather than to thermal 

cooling. In the late 1990s, modeling research began to examine the interactions of droplets with flames. 

Chen et al.64 considered the motion of droplets in a non-premixed counterflow field. They considered 

both n-heptane and water droplets. For the droplet sizes considered, Chen et al. showed that a droplet can 

penetrate through the stagnation plane established by the gaseous streams and reverse its trajectory at 

some point in the opposing side of the flow. Prasad et al.65 modeled water mist suppression of diffusion 

flames, including the dependence on the droplet trajectories into the flame. 

4.3 THE NATURE OF AEROSOLS 

4.3.1 Quantitative Descriptors 

Particle Diameter 

Aerosols are almost never monodisperse (all drops/particles of exactly the same size). Real aerosols can 

be nearly monodisperse, but more typically exhibit a range of particle sizes. A mean or average diameter 

is useful for providing specific information relating to a particular phenomenon controlling the process 

under investigation.66 (See Table 4–1) These definitions are strictly true only for spherical aerosol 

particles or liquid drops. 

Table 4–1. Mean Diameter Definitions, Symbols, and Notations used to Describe Aerosol 

Size Distributions. 

Mean Diameter Basis Symbol Notation 

Arithmetic or length D10 ΣNiDi /Σ Ni

Surface area D20 (ΣNiD2

i /Σ Ni Di)

1/2

Volume D30 (ΣNiD3

i /Σ Ni)

1/3

Sauter (SMD) D32 ΣNiD3

i /Σ Ni D2i

The arithmetic or length mean, D10, is simply the mean diameter of the spherical particles. The Sauter 

Mean Diameter (SMD) is that diameter for a single particle whose ratio of volume to surface area is the 

same as that of the entire aerosol. While D10 and the SMD values can be used to compare the mean 

diameters or surface area ratios of various aerosols, they do not convey information regarding the 

distribution of particle sizes. 

Aerosols can be also be described by defining a mass accumulation diameter such that all droplets below 

this diameter make up a certain percentage of the total mass of the aerosol. For instance, D0.5 is that 

droplet size such that half (0.5) of the liquid volume is in droplets of smaller diameter. The D0.5 diameter 
The Nature of Aerosols 349

is also referred to as the mass median diameter (MMD). Other common mass accumulation diameters are

for 10 %, D0.1, and for 90 %, D0.9. 

Particle Size Distribution 

Aerosols with similar D10 values, D0.5 values, D30 values, or SMDs can have very different size

distributions. Since aerosols of different size can exhibit different degrees of fire suppression 

effectiveness, characterization of the range of sizes or size distribution is necessary to compare the

suppression performance of different aerosols and to make valid comparisons between experimental

observations and modeling predictions. Several mathematical expressions for size distributions have been

proposed, although each of them has some shortcomings. The expression to use is the one that best

characterizes the relevant properties of the mist under study. 

A simple and practical method for displaying droplet size information is to plot a histogram where the 

number of droplets in a defined size bin is plotted for that bin. In practice, the minimum size plotted is

dictated by the physical limitation of the instrument settings. A water mist histogram (obtained with a

phase Doppler interferometer as discussed below) is presented in Figure 4-1. Values for the mean 

diameters defined in Table 4–1 for this mist are indicated in the table. In the limit of small bin size, the 

histogram becomes a frequency or size distribution. 

Figure 4-1. Water Mist 

Size Histogram and 

Values of Several Mean

Diameters. 

Two commonly used functional expressions describe the droplets as (a) a Gaussian distribution or

(b) a mass accumulation.67 These simple expressions cannot be used to describe bi- or multi-modal 

distributions. For these, linear combinations of the different single expressions are sometimes used. 

The Gaussian or normal distribution with respect to the natural logarithm (lognormal) of the diameter is 

given in Equation (4−1). This expression is often sufficient to fit observed experimental droplet size

distributions, where dmn is the number geometric mean droplet diameter D10, and s is the corresponding

standard deviation describing the spread in the distribution. 
350 Aerosol Properties 

(4-1) ( ) ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ −

∗ = 2

2

2

{log( ) log( )} exp

0.399

s

d d

s d

N d mn

The Rosin-Rammler equation (mass accumulation expression) describes the aerosol by relating the mass

of aerosol material contained in particles up to a certain diameter. The Rosin-Rammler expression is

given in Equation (4−2). In this expression, a and b are constants related to a mean diameter and the

spread, respectively.

( ) ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ ⎟

⎠

⎞ ⎜

⎝

⎛ = −

b

b

b a

d d

a

b d exp 4 N (4-2)

Note that each expression includes two parameters to describe the distribution. One advantage of using 

the Rosin-Rammler expression is that all of the representative diameters (e.g. D0.1, D0.5, etc.) can be 

related to each other through the spread.

Both of these two mathematical descriptions of the histogram in Figure 4-1 are presented in Figure 4-2. 

The values of the two parameters are: D10 = 93 μm, s=0.15 (lognormal) and a=101, b=9.4 (RosinRammler). Both expressions qualitatively describe the mist, although the lognormal expression is a better

match for this mist in both the peak position and spread in the distribution. 

Figure 4-2. Lognormal (dashed line) 

and Rosin-Rammler (solid line) 

Expressions Fit to the Water Mist 

Histogram Presented in Figure 4-1. 

4.3.2 Generation Techniques 

Generating, controlling, and measuring aerosols are challenging but necessary tasks for studying fire

suppression as well as evaluating fire suppression systems. Although concentration standards are 

straightforward for gases, there is no aerosol standard. An aerosol generation system that could 

reproducibly produce a uniform distribution of a monodisperse (or polydisperse) aerosol would be

extremely useful. 

Aerosols of liquid droplets can be produced using ultrasonic agitation of the liquid, either disrupting the

liquid surface or modulation of a liquid jet stream. NGP studies employed both methods; the specifics 

are outlined in Chapters 7 and 8. Assisted breakup of liquid jets can produce aerosol droplet size 

distributions that are narrow, approaching a single, monodisperse droplet size for all of the drops. This
The Nature of Aerosols 351

generation method is ideal for small, bench-top studies only requiring a small number of drops. The 

number of droplets produced is in general not sufficient for large-scale suppression applications. 

Powder aerosol samples are typically produced by physically crushing solid samples, followed by size 

separation using various sieving protocols. Sample crushing often results in a range of particle shapes in 

a given size bin, where individual shape differences will depend on the nature of the powder and the 

crushing/sieving protocol. Sample drying is often required to improve the flow properties, especially for 

very small particles. Small concentrations of flow enhancers such as silica can improve flow properties. 

Large masses of powder particles can be generated using pyrogenic methods discussed in Section 4.2.5. 

Particle size and composition (including the addition of chemical suppressant moieties) can be varied 

depending on the type of generator and the particular propellant formula. Because of the high particle 

momentum and often high gas temperatures exiting the generators, pyrogenic powder aerosol generation 

techniques are rarely practical for bench-scale studies. However, in large-scale applications, the high 

particle momentum can be very useful in aiding agent dispersal. 

4.3.3 Quantification of Aerosols 

Gaseous fire suppression agents can generally be monitored and controlled by regulating the volumetric 

flow of the agent. Gaseous agents generally distribute throughout the measurement volume so that agent 

spatial distribution issues are generally dictated by the physical constraints of the application geometry. 

Aerosol agents will not uniformly fill the application space, especially very large, obstructed spaces. In 

order to determine the amount of aerosol agent in the application area, measurement techniques must be 

employed that are applicable and can provide relevant, valid information. Relevant information includes 

data on how the aerosol is distributed spatially in the application area, especially in and around any fires 

that may be present. 

4.3.4 Aerosol Size and Velocity Characterization 

One of the most challenging aspects of working with aerosols is characterizing the aerosol that has been 

generated in order to obtain relevant, quantitative information. Such quantitative information, even on a 

relative basis, is required in order to make valid suppression effectiveness comparisons. A general 

discussion of the characterization methods used in NGP aerosol fire suppression studies is outlined below. 

Specific details are given in the individual studies. 

Diffraction Method 

Diffraction-based systems are capable of providing size information (either mass or volume) of an 

aerosol.68 The experimental layout is fairly simple, as shown in Figure 4-3. A coherent light source (i.e., 

a laser) and a two-dimensional detector (i.e., a linear array of detectors or a charge coupled device, CCD, 

camera) are needed. The laser light, directed at the detector, passes through the sample volume 

containing the aerosol. Laser light diffracts around individual droplets in the aerosol (Fraunhofer 

diffraction). The light intensity pattern falling on the camera is a superposition of all of the single droplet 

diffraction patterns. By deconvoluting the diffracted intensity from the ensemble of drops, the mass or 

volume distribution can be derived. Diffraction methods do not rely on a spherical droplet or particle, 
352 Aerosol Properties 

although the determination of volume or mass must assume some shape. The ease of set-up and 

straightforward calibration requirement make diffraction methods attractive. 

Figure 4-3. 

Schematic Diagram 

for Diffraction-based 

Particle Sizing. 

Diffraction methods lack the spatial resolution of other optical approaches. They measure everything that 

lies in the beam path. Because they rely on diffraction, higher power laser sources are often needed. 

Diffraction methods also do not provide velocity information, and thus cannot provide flux information

directly. Although the lack of velocity information is an issue, diffraction methods can provide droplet

size distribution and density information. An effective mass flux can be estimated for well-characterized

flows.

Phase Doppler Interferometry Method 

Phase Doppler Interferometry (PDI, also referred to as phase Dopler anemometry, PDA) is an established

technique that can provide velocity information or both velocity and size for individual, spherical 

particles.69 PDI systems which only provide velocity information are called Laser Doppler Velocimetry 

(LDV) systems. PDI measurement times can be as short as 1 μs, so that statistically significant size

histograms can be generated in less than a second depending on the particle number density. Droplet

sizes from ~ 1 μm to several mm can be measured, although not with a single system configuration. 

PDI droplet sizing relies on the phase difference between two signals originating from the same point and 

arriving at two spatially separated detectors. Each signal is generated by a droplet passing through a set

of parallel grating lines produced from the overlap of coherent laser beams. The technique is depicted 

schematically in Figure 4-4 with a representation for the ellipsoidal "grating" measurement volume. PDI

determined diameters are only strictly valid for spherical drops. A third detector, providing a second 

phase difference, is used to check the sphericity assumption. 
The Nature of Aerosols 353

Figure 4-4. Diagram for 

1-D Phase Doppler 

Interferometry System for 

Obtaining Size and 

Velocity of “Spherical” 

Aerosol Particles. 

The size of the ellipsoidal measurement volume depends on the wavelength of the laser, the beam 

diameter at focus, and the crossing angle. Typical measurement volumes are ≈ 1 mm long by 0.2 mm 

diameter for droplet sizes relevant to water mist suppression. Only a single droplet can be present in this

volume during the measurement period for the measurement to be valid. This requirement can be 

problematic for very large number densities (e.g., > ≈ 104

 drops/cm3 for uniformly dispersed 50 μm 

diameter drops).

The laser wavelength and beam-crossing angle control the grating line spacing. Passage of the droplet 

through the grating generates a Doppler burst pattern containing the droplet 1-D velocity information

perpendicular to the fringes. Additional hardware is required for 2-D and 3-D velocity information. The

grating fringe spacing must be reasonable for the droplet size in order to determine a valid burst pattern 

and provide the corresponding velocity. For closely spaced lines and large drops, there may not be

sufficient modulation in the signal to qualify as a valid burst. Large spacings and small droplets may not 

provide enough modulations for the phase shift validation. PDI signals can originate from light reflected

from the droplet surface or refracted through the drop. The angle of the refracted light exiting the droplet

is dependent on the liquid index of refraction that must be known. Reflected light does not depend on the

index of refraction but signals are generally much weaker. The choice of scattering angle will determine 

which scattering mode is dominant. When a single scattering mode is dominant there will be a linear

correlation of the droplet diameter with the phase difference. If more than one scattering mode

significantly contributes to the signal, the phase angle/diameter relationship will not be linear. For water 

droplets in air, first order refraction dominates light scattered at 30o in the forward direction. A scattering

angle of 30o was used for measuring water droplets in the NGP studies. 

Uncertainties in the PDI size determinations associated with the scattering angle and refractive index are

generally less than a few percent. Because of the finite range for the detector and the non-linear

dependence of the scattering intensity on the droplet size, the usable dynamic range dictates a droplet size

dependent precision in the determination of the sizes. Additional potential limiting factors are the laser

power, detector gain setting, and S/N ratio. These combined effects typically limit the uncertainty for a 

PDI size determination to at least 0.5 % of the full-scale reading. Velocity uncertainties depend on the

processor settings, primarily the need to divide the velocity space into a finite number of bins. Binning 

depends on the size range to be measured, which depends on the optical configuration used. In addition 

to the dependence on processor settings, uncertainties in the velocity determination depend also on the
354 Aerosol Properties 

instrument A/D resolution, fringe spacing uncertainty, and sampling frequency. Velocity uncertainties 

can be as high as ± 25 %.

The mass flux determination relies on the droplet size, velocity, and the probe volume. The droplet 

number density also relies on the probe volume. However, the probe volume is droplet size dependent, 

requiring a correction for both number density and mass flux. The dynamic range for a single PDI system 

configuration is typically 40:1. This value depends on several factors, including the optical system, laser

power, detector gain and dynamic range, and system S/N ratio. Obtaining accurate size information

becomes a compromise between the size resolution desired and the range of droplet sizes that can be

determined. 

Imaging Method 

Capturing an image of the aerosol particle can provide both size and velocity information. Unlike 

diffraction-based techniques, imaging techniques can be used to “size” non-spherical particles or

spherical objects such as drops. Imaging techniques have been used extensively to study and understand 

sprays and spray formation and development.70 In its basic form, an imaging configuration is fairly

simple, as shown in Figure 4-5. It consists of a light source and a recording medium, either photographic

film or more typically a CCD camera. The light source is normally a laser, which provides sufficient

spectral brightness in a short time period to "freeze" the motion of the drop. Droplets can either be back

illuminated in shadow method, as indicated in Figure 4-5, or illuminated from the front. The shadow

method is more often used, as the droplet size dependent back-reflected light can complicate data 

collection. In the shadow method, droplet images appear as dark spots on a light background. The 2-D 

velocity information in the image plane can be determined from timing droplet exposure "tracks" or by

double pulsing the light source, creating two images of the same droplet displaced according to its 

velocity and the illumination time delay. The two images can be in the same camera frame or in 

sequential frames. Software algorithms correlate the images, permitting both speed and trajectory

information. Algorithms permit false shadow rejection from images that remain in the same location 

frame after frame (e.g., a spot on a lens or window). 

Figure 4-5. Schematic 

Diagram for Imaging-based 

Particle Sizing. 

Although cameras with photographic film can be used, CCD cameras provide greater flexibility. Cameras 

can be digital or analog. Digital imaging requires a large data storage capacity if the raw images are

saved. This may limit the time for continuous data collection. Images can be processed either on-line or

off-line. On-line image processing generally saves time but unless the images are also stored, valuable 

information may be lost. Off-line processing requires that the raw images be stored. Analog cameras
The Nature of Aerosols 355

with tape storage (usually super VHS) offer the advantage of long data collection times (several hours). 

Acquiring the raw images and performing off-line data reduction is an added time expense, but can be an 

advantage. Repeated data reduction using different fitting parameters can sometimes identify issues with 

the data collection or point out systematic problems. 

Droplets from a few μm to several mm may be imaged but the total size range requires a change of optics. 

The measurement region that can be imaged depends on the magnification but is typically 1 mm2

 for 

detecting ~ 4 μm to 300 μm diameter drops. Similar to microscope optics, imaging magnification factors 

must be greater than one, requiring that the distance of the droplets from the focusing lens be much 

shorter than the distance of the image from the lens. In practice, this means that the sample volume must 

lie fairly close to the collection lens for the smallest droplets. 

Size and velocity measurement accuracy is controlled by the lens focusing system, the media resolution 

(photographic grain limit or camera pixel resolution), and the determination of degree of focus. Since the 

data obtained are a mixture of in-focus and out-of-focus images, image analysis schemes must be 

employed to decide which images are valid and then render sizes. The degree of focus impacts both the 

location of the droplet edges and also the determination of the measurement volume. 

The uncertainty in the measurement volume is the largest source of uncertainty for imaging detection. 

Both the droplet number density and mass flux determination are impacted by this uncertainty. The 

derived particle volume depends on the size of the drop but fortunately, computers can aid in eliminating 

the human bias in the decision-making. The key to achieving a reasonable accuracy lies in the calibration 

and validation processes. 

Particle Image Velocimetry (PIV) is the macroscopic analog of droplet size imaging.71 In PIV, the 

marker particle image is not resolved since size resolution is not desired. Marker particles are illuminated 

with two temporally delayed laser sheets. Small particles, typically < 1 μm, are used to insure that the 

marker particles validly represent the flow field. Large areas (> 1 m by 1 m, depending on the spectral 

brightness of the laser source and the collection optics) can be characterized. By recording two shots in 

the same frame or in sequential frames, both speed and direction of movement of the particle can be 

derived through image analysis. PIV is typically used to monitor bulk gas or liquid flows. It can also be 

used to monitor the flow behavior of larger aerosol particles that, because of their size, do not follow the 

flow field streamlines. As is shown in this chapter, suppression aerosol particle velocity (speed and 

trajectory) controls the overall aerosol effectiveness. Such velocity information becomes key in 

understanding the suppression mechanism for aerosol agents. 

4.3.5 Aerosol Properties Affecting Behavior 

Evaporation/decomposition behavior 

Rosser et al.13 and Mitani and Niioka,43 based on observations for particle suppression behavior in 

premixed flames, described the flame suppression mechanism in terms of particle size and residence 

times. Generalizing these findings to both liquids and solids, four steps with their characteristic times 

relevant to suppression behavior of an aerosol (liquid droplet or solid particle) can be assumed: 
356 Aerosol Properties 

Step I: aerosol heating, 

Step II: aerosol evaporation (solid or liquid) and/or decomposition (primarily solids), 

Step III: generation of the inhibitor radical species in the gas phase, and 

Step IV: inhibition of the gas-phase combustion process. 

Optimum inhibition efficiency requires completion of all four steps. The slower steps determine the

extent of inhibition and what aspects of the condensed phased suppression mechanism contribute to the 

inhibition process. The times for Steps III and IV involving gas-phase reactions are typically much 

shorter than the drop/particle heating times or the decomposition/evaporation times for Steps I and II. 

Heating of the aerosol in Step I needs to raise the temperature to the aerosol's boiling/decomposition 

temperature. The particle heating will depend on the size of the particle, the particle thermal conductivity

and that of the surrounding gas, and the gas temperature. Flames temperatures encountered are typically 

≈ 1800 K. Assuming the aerosol particles are small, the heating time for Step I is typically short,

<< 1 ms. 

The evaporation of the aerosol particle in Step II involves losing mass (vapor) to the surroundings.

Several models have been proposed for treating the evaporation of liquid drops, depending on what

regimes are relevant and what assumptions are appropriate. One model suitable for the experimental

conditions under consideration here is72

d (t) = d − Kt 2

0

2

(4-3)

where K, the evaporation constant, is given by 

( ) ⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ = + −

vap

p

boil

p h

c

T T

c

K ln 1 8

ρ

λ

(4-4)

The time for complete evaporation to occur can be estimated from:

K

d

tvap

2

0 = (4-5)

Equation (4−3) is referred to as the d-square law for droplet evaporation and predicts the droplet diameter, 

d, at any time, t, in terms of the original droplet diameter, d0. The mass transfer is dependent on the 

temperature, T, and thermal conductivity, λ , of the surrounding gas as well as on the properties of the 

drop: heat capacity at constant pressure, cp; density, ρ; enthalpy of evaporation, hvap; and boiling point,

Tboil. According to Equation (4−5), the time required for the droplet to evaporate completely is dependent 

on the original diameter of the drop. 

Evaporation times for water as a function of droplet size for temperatures ranging from 1100 k to 2200 K 

are shown in Figure 4-6. The evaporation time for a 30 μm diameter water droplet in an air stream at 

2100 K is predicted to be ≈ 2.5 ms. The thermal conductivity data for air are from Kays and Crawford.73

Whether the droplet will have sufficient time to evaporate partially or completely will depend on how 

long the droplet resides in the higher temperature region. This time depends on the specifics of the flow

field configuration. 
The Nature of Aerosols 357

Figure 4-6. Predicted Evaporation Times 

using Equations (4−3) through (4−5) for 

Water Droplets in Air at the Indicated 

Temperature as a Function of Initial 

Droplet Diameter. 

Flow Entrainment 

Generation of the aerosol is only a first step in producing an aerosol suitable for further study. For flame 

studies, the aerosols must be seeded or entrained into a flow that hopefully results in as uniform a 

distribution as practical. Both uniformity in the particle mass flow as well as uniformity of the 

distribution of particle sizes representative of the entire distribution are needed to be able to make valid

comparisons and quantitative evaluations. In order to entrain the aerosol particle in the carrier gas flow 

with a minimal difference in the particle and gas velocities, the gas flow velocity must equal or exceed the

particle’s terminal settling velocity74 in that gas. 

The terminal settling velocity assuming Stokes flow for a spherical particle can be expressed as 

( )

gas

P gas

T gd

η

ρ ρ

ν

18

2 − = (4-6)

The terminal velocity, vT, depends on the acceleration due to gravity, g, the density difference between the 

particle, ρp, and the carrier gas, ρgas, and the viscosity of the carrier gas, ηgas. Values of vT for water

droplets and sodium bicarbonate particles, two aerosols studied in the NGP program, are shown in 

Figure 4-7. Terminal settling velocities are centimeters per second depending on diameter and have a

quadratic dependence on particle diameter. 

Figure 4-7 Predicted Terminal Settling 

Velocity for Water Droplets and Sodium 

Bicarbonate Particles in Air at 20 oC. 
358 Aerosol Properties 

Entrainment velocities adequate for studying water droplets may be insufficient for studying the same size 

sodium bicarbonate particles. Thus, the particle terminal settling velocity put limits on the size of aerosol 

particles that can be entrained and/or delivered to the experiment in a primarily vertical flow. For 

aerosols whose size distributions are non-monodisperse, low entrainment velocities can result in particle 

size discrimination. In some cases, velocity discrimination can be used to restrict the aerosol particle 

sizes introduced to the experiment. Thus, entrainment flow conditions must be considered when carrying 

out experiments as well as in the interpretation of data. 

Particles moving in a flow with a thermal gradient will also tend to experience a thermophorectic force 

pushing them in the direction of lower temperature. The effect is greater for smaller particles and large 

temperature gradients. High temperature environments are exactly the conditions that favor rapid 

disappearance of the condensed phase aerosol agents due to evaporation or decomposition. 

4.4 THE NATURE OF FLAMES 

4.4.1 Flame Characteristics 

Suppression of fires involves heat extraction, oxygen reduction, and, for chemical agents, kinetic effects 

at the molecular level that involve heat production and the concentration and distribution of key flame and 

inhibitor species. Despite the studies carried out prior to the NGP, considerable uncertainties existed in 

the assessment of aerosol suppression pathways, especially in the quantitative interpretation of 

suppression behavior. Early literature alluded to a dependence of aerosol suppression effectiveness on the 

fundamental aerosol size (liquid droplets or solid particles), and properties (composition, thermal 

stability, reactivity). In addition, the application environment (flow field, flow orientation, flame 

temperature) of the various fires investigated imposed characteristic residence times controlling the 

performance of the aerosols. Thus, characterizing these flame conditions is necessary to interpret and 

understand the aerosol suppression results. 

For a system that undergoes combustion, the fuel and oxidizer can either be premixed (combined) or nonpremixed (kept separate) upon entering the reaction region. Real fires often exhibit a combination of both 

mixing protocols. Figure 4-8 compares gaseous flame configurations to each other and the variation in 

the flow field complexity needed to describe and understand them. In general, the greater the complexity 

of the flow field, the more computationally challenging it is to include both the details of the flow and the 

details of the chemistry in the modeling treatment of the flame process. This is even more apparent in 

multiphase flow problems, such as the interaction of solids or liquids with gaseous flames. In order to 

understand the extinction mechanism of aerosols and approach a quantitative understanding of the 

aerosol/flame interaction, NGP researchers studied a number of flames depicted in Figure 4-8. This 

sometimes required a re-examination of key experiments in the literature with the focus on minimizing 

the number of parameters that might vary and exploring the dependence of several of these parameters on 

the overall suppression performance. Researchers paid close attention to the control of experimental 

parameters that could influence the results. In this way, the new data could be used more directly to 

validate the predictions of multi-phase flame suppression codes as well as aid in their development. The 

next sections address the different flame configurations explored under the NGP aerosol studies, discuss 

the metrics used to determine suppression effectiveness, and explore the nature and impact of the imposed 

residence times on the suppression properties of the aerosols investigated. 
The Nature of Flames 359

Figure 4-8. Relationship of Flame Dimensionality and Modeling Complexity. 

4.4.2 Premixed Bunsen Type Flames 

Premixed flames offer the least complicated flow field for gaseous flames since they can be configured to

give flat, essentially one-dimensional flames. Two such flames are shown schematically in Figure 4-9 

from Reference 75. The ideal one dimensional flat flame shown at the left can be experimentally

achieved with gaseous reactants and gaseous suppression agents. This configuration is difficult to 

achieve when aerosols must be introduced uniformly into the flame. Experimentally seeding aerosols into 

the premixed flow and maintaining the flat, one-dimensional geometry is very difficult. 

Figure 4-9. Representations of Aerosol 

Suppressed Premixed Flames: (left) Ideal Mist 

Inhibited Flat Flame as Simulated in 1-D 

Modeling; (right) Bunsen-type Flame Typically

Studied Experimentally. 
360 Aerosol Properties 

The air/fuel flow in a premixed Bunsen-type flame, shown on the right, can be fairly easily seeded with 

aerosols. The figure shows introduction of the aerosol which are seeded into the premixed gases. This is 

the experimental flame more typically studied with aerosols. 

The laminar burning velocity is a fundamental property of a flammable gas mixture and is frequently used 

as an indicator of the effectiveness of an inhibiting agent.76 There are various methods to measure 

laminar burning velocity.77 Each method involves recording some aspect related to the burning region. 

The visible flame emission can be converted into a total burning surface area as can schlieren images of 

density gradients in the flame region due to regions of varying temperature. Both have been used to 

quantify the burning velocity. Although each gives a slightly different absolute surface area, they provide 

the same information when comparing suppressed flames to the unsuppressed flame. Ratio comparison 

minimizes systematic differences between various methods for burning surface derivation. 

4.4.3 Non-premixed Flames 

Counterflow 

The counterflow or opposed flow flame configuration provides an excellent experimental and modeling 

arrangement to study the effects of aerosols on a flame and the effects of high temperature environments 

on the aerosol.64,78 79 This configuration has been used to study the effect of aerosols on premixed flames 

by using two opposing premixed fuel and oxidizer flows.64 Advantages of the counterflow arrangement 

include the ability to investigate variable flow residence times and potentially to separate chemical from 

physical effects. 

The counterflow flame configuration is depicted schematically in Figure 4-10. The delivery tubes can be 

converging nozzles or straight tubes. The counterflow burner can be used to study gaseous and aerosol 

agents. In the non-premixed mode, fuel is introduced through one inlet and oxidizer though the other. 

Non-premixed flames of hydrocarbon fuels in air burn near the fuel/air stagnation plane in a region that 

contains the stoichiometric composition for complete combustion. For hydrocarbon fuels, the flame 

resides on the air side of the stagnation plane, as depicted in Figure 4-10. The fuel can enter from the top 

or bottom for non-premixed flames, with air introduced from the opposite tube. The aerosol can be added 

to either the fuel or the air. Air side introduction of the aerosol is relevant to total flooding fire 

suppression scenarios and is the configuration most often studied. Transport and composition dependent 

suppression effects can be studied by comparing extinction results of air-introduced aerosols with results 

obtained when the aerosol is introduced into the fuel. 

Representative centerline temperature and axial velocity profiles are indicated, as is the strain rate, a, the 

maximum gradient in the axial velocity on the air side of the flame. In this flow field, stable flames can 

be maintained if sufficient amounts of heat and flame radicals generated from the combustion process can 

be conducted to the unreacted fuel and air as they flow into the reaction zone. Extinction occurs when the 

flow is too fast to maintain the flame. Under these conditions, the flame strain rate exceeds a critical 

value referred to as the extinction strain rate. Suppression agents lower the extinction strain rate and this 

reduction is a direct measure of the suppression effectiveness. The flame strain rate can be estimated 

from the global strain rate.80 Since the global strain rate can vary between burner designs, the local strain 

rate, determined by experimentally measuring the local velocity field, is used for absolute comparisons. 
The Nature of Flames 361

Figure 4-10. Experimental Schematic for 

Studying the Effects of Aerosols on the 

Extinction Strain Rate of Counterflow 

Non-Premixed Flames.81

The reciprocal of the strain rate can be used to estimate a characteristic residence time for an aerosol 

particle in this flow field. Depending on burner design, the local strain rate can range from ~ 10 s-1 to the 

value for extinction in the absence of any inhibitor (~ 600 s-1 for propane/air flames). Thus, the particle 

residence time can be varied by an order of magnitude. Chen et al.64 note that for certain conditions the

aerosol particles can often be trapped and oscillate in the opposing streams, thereby increasing their 

effective residence times. 

Another design for studying suppression of non-premixed flames in opposed flow is the Tsuji burner.41 

Gaseous fuel is introduced through a porous tube perpendicular to the air flow as seen in Figure 4-11. 

NIST incorporated a variant of this design to develop a suitable NGP screening test for condensed phase

agents.82 Details of that work can be found in Chapter 8. 

In the Tsuji configuration, flames can be established in the forward, stagnation region or downstream of 

the air flow in the fuel tube wake. In the forward direction, the two-dimensional flames are very stable

and can be accessed readily with various diagnostic probes. The Tsuji flame established in this flow field 

also lends itself to modeling. Since the fuel and oxidizer flows can be independently controlled, flame 

strain rate can also be easily varied and studied. The air stream velocity required to blow-off the forward,

stagnation flame is used to evaluate suppressant effectiveness. 

In the Tsuji configuration, as in the premixed and opposed flow configurations introduced above, aerosols 

introduced with the air stream (or their decomposition products) are restricted to enter the flame. This 

flow arrangement assures the optimum effectiveness for the added agent. Aerosol particles that are too 

large to completely evaporate or decompose in the flame could potentially collide with the porous fuel 

tube, requiring periodic tube cleaning or replacement. 
362 Aerosol Properties 

Figure 4-11. Schematic of Tsuji Burner (left41) and Flame Configuration (right). 

Cup Burner Co-flow Flames 

Cup burner flame extinction results are commonly used in the fire protection industry for suppression 

agent effectiveness ranking as well as for suppression system design guidance. The cup burner flame is a 

non-premixed flame stabilized on a cylindrical cup that holds liquid fuel or delivers a low-velocity gas jet. 

A surrounding laminar co-flow of air carries the suppression agent of interest.83 The majority of the

agents studied have been gaseous. 

Cup burner flames have many similarities to real fires: initial separation of reactants, flame flickering, and 

regions of varying strain rates. These similarities make cup burner extinction results useful for

suppression agent ranking and for evaluating extinction requirements in real fire scenarios. These 

similarities also make the cup burner flames challenging for studying aerosol suppressants. Nonetheless, 

researchers had investigated the suppression behavior and effectiveness of condensed-phase particles in

cup burner flames prior to the NGP. (See the work of Hamins et al.20,21

) 

A cup burner apparatus modified to study the suppression properties of water mist is shown in Figure 

4-12.84 Water droplet entrainment and flow conditioning occur in the lower section, water quantification 

occurs in the middle section, and the flame is located in the uppermost section. 

Under the NGP, Linteris et al. studied the effects of various metallic compounds on cup burner flames. 

Several of the metallic compounds were found to be less effective than expected, surprisingly less 

effective compared to their performance in premixed or opposed flow flames. Even though the agents are 

introduced in the gas phase, their poor fire suppression effectiveness has been attributed to the formation

of condensed phase particles in the cup burner flow field. Details of that work are presented in Chapter 3. 
The Nature of Flames 363

Air In

Fuel In

Mist

generator

Auxiliary reservoir 

Mesh screens

Entrainment 

tube

1370-nm

DFB DL

Detector

mount

Chimney

Quartz cup

Figure 4-12. Cup Burner Apparatus Modified for 

Water Mist Suppression Studies. 84

4.5 AEROSOL FLAME SUPPRESSION STUDIES 

4.5.1 Water, a Physical Suppression Agent 

Water Mist Studies in Premixed Flames 

Experimental

NGP researchers at NRL studied the flame suppression properties of water mist as a function of droplet 

size, measuring the effectiveness of the mist on reducing the burning velocity of methane/air flames.

85 

Mists were generated using various methods. Droplets smaller than 1 μm in diameter were generated 

using a venturi-based nebulizer (TSI Model 3076). A schematic of the nebulizer is shown in Figure 4-13. 

Pressurized air, flowing through a small orifice, produced a high velocity air jet. The resulting pressure 

drop drew liquid from a reservoir through a small tube and entrained it in the air jet, breaking the liquid 

into drops. Larger droplets impacted the nebulizer wall and returned to the reservoir. Droplets small 

enough to remain entrained in the air flow were carried from the nebulizer. The amount of mist exiting 

the nebulizer was determined by measuring the change in mass of the liquid reservoir with time for a 

fixed air flow. Mists of polydisperse droplets in the range of 1 μm to 20 μm were produced using a 

second venturi-based nebulizer (Airlife Nebulizer with Air Entrainment and Heater Adapter), requiring a 

minimum air flow of ≈ 2.9 L/min to generate drops. 
364 Aerosol Properties 

Figure 4-13. Schematic of 

Water Aerosol Generation 

Methods. (left) VenturiBased Atomizer For 

Generation Of Sub-Micron 

Diameter Water Drops; 

(right) Vibrating Orifice 

Aerosol Generator for 

Generating Monodisperse 

Water Droplets > 15 μm 

Diameter. 

Monodisperse droplets of larger size (> 15 μm diameter) were generated using a vibrating orifice aerosol

generator (VOAG, TSI Inc. Model 3450). A schematic of the droplet generator is shown in Figure 4-13 

(left). To generate the mist, liquid was forced through a pinhole using either a syringe pump or a plastic 

bag under pressure. Limited control over the liquid flow was accomplished using either method. The 

size of the pinhole ultimately determined how much liquid could be delivered to the burner. The pinhole 

was acoustically excited by a piezoelectric ceramic driven at specific frequencies to break up the liquid jet 

into a stream of extremely monodisperse drops. The droplet size depended on the pinhole diameter, the 

liquid flow, the forcing frequency, and the liquid properties. A dispersion cap was used to distribute the 

droplet stream uniformly into a three-dimensional mist. This sacrificed some of the size monodispersity

as shown in Figure 4-14 for the three water mists that were studied. In the absence of the dispersion cap 

and dispersing air, all droplets occupied a single PDPA size bin. 

Figure 4-14. PDPA-determined Droplet Diameter Histogram 2 mm above the Air Tube Exit 

for Three Water Mists.85

Droplet size, velocity, and number density for droplets greater than ≈ 2 μm in diameter were measured

using a phase Doppler interferometry system (Phase Doppler Particle Anemometer – PDPA, Dantec 

Measurement Technology). For droplets smaller than ≈ 1 μm diameter, water concentration in the
Aerosol Flame Suppression Studies 365

reactant stream was monitored by 90o scattering of a Helium-Neon laser at a wavelength of 0.6328 μm as

shown in Figure 4-15. The scattered laser light was detected with a photomultiplier tube (PMT) using a 

narrow-line optical filter. The laser beam was modulated at 1 kHz by passing the beam through a rotating 

slotted wheel. The laser beam was positioned just above the burner exit. A reference beam was split off

just before the burner and monitored by a photodiode. The scattered and reference signals were processed

with a lock-in amplifier (EG&G Instruments, Model 7265 DSP) and integrated over a period of 0.5 s. The

scattering intensity was calibrated to a mist delivery rate by correlating the scattering signal with the 

change in water mass of the reservoir. 

The total flame surface area derived by visible flame imaging was used to quantify the reduction in

burning velocity as a function of added mist for suppressed premixed conical flames. A schematic of the 

arrangement is shown in Figure 4-15. 

Figure 4-15. Schematic for Liquid Aerosol Inhibited Burning Velocity Determination in 

Premixed Flames.75

The burner was a converging nozzle with an exit diameter of 1.0 cm. An entrainment device to straighten

the flow and uniformly seed droplets in the premixed gas flow utilized a perforated disk. The sub-micron

size mist droplets were small enough to pass through the ≈ 0.16 cm diameter perforation holes. Larger

droplets (> 5 μm diameter) were introduced through an opening in the middle of the perforated disk. The

burner was enclosed inside a 13 cm diameter inner diameter acrylic tube that allowed optical access to the
366 Aerosol Properties 

flame. The acrylic chamber was vented to an exhaust hood through a conical metal covering open at the 

top. A nitrogen flow of approximately 3 L/min was introduced through a sintered disk in the bottom of 

the chamber. This purged the chamber of excess oxygen and product gases. Flame images were recorded 

using a CCD camera and a value for the total flame surface was determined by fitting the shape and then 

mathematically generating a three dimensional surface. Burning velocities were determined by 

comparing the total flame surface to the volumetric flow of gases in the burner. Light from a modulated 

He-Ne laser, scattered from the aerosol in the premixed flow and detected with a lock-in detector, was 

used to determine the amount of water aerosol in the form of small droplets introduced into the flame. 

PDPA was used to quantify the water concentration for mists of larger drops. 

Burning velocities for the inhibited flames were reported relative to the uninhibited burning velocity, 

thereby minimizing systematic errors in determining an absolute value for the burning velocity. Relative 

values derived from schlieren images for some of the flames were in good agreement with values derived 

from the visible flame surface. The presence of the aerosol reduced the flame burning velocity, which 

resulted in a taller flame with a larger flame surface area for a fixed fuel-air flow. To minimize any 

systematic effects due to varying flame height, fuel and air flows were adjusted to maintain a 

stoichiometric mixture and a flame height of 1.0 cm ± 0.15 cm. The measured burning velocities of 

nitrogen-inhibited flames were found to be insensitive to flame height over this range. 

Modeling 

Both theoretical86,87 and computational88,89 models have been developed to predict premixed flame 

structure and extinction characteristics as a function of water-mist properties. NGP modeling of submicron mists in premixed flames was done in collaboration with Prof. Robert Kee at the Colorado School 

of Mines. The suppression effects of gas-phase agents were modeled using PREMIX90 and OPPDIFF.91 

The chemical reaction mechanism and the associated thermodynamic and transport properties were taken 

from GRI-Mech 3.0,92 but with the nitrogen chemistry removed. The inhibition effects of water aerosols 

on premixed flames were calculated using a multiphase combustion model.89 The multi-phase model 

solves the conservation equations for mass, momentum, and energy for the water droplets within a 

Lagrangian framework. One-dimensional conservation equations in an Eulerian framework were used to 

solve the physics of the gas-phase flame propagation. Coupling of the two phases was achieved through 

addition of drop-evaporation related source terms in the PREMIX software. Boundary conditions for the 

drop-evaporation dynamics were obtained from the PREMIX solution. 

Figure 4-16 illustrates the predicted dependence of the burning velocity on droplet size and water 

loading.89 For small drops, the burning velocity decreases monotonically as a function of water loading. 

For sufficiently small drops, the burning velocity becomes independent of droplet size but still depends 

on water loading. For stoichiometric, methane-air, premixed flames, this small-droplet limit is achieved 

for a droplet diameter of approximately 10 μm.89 

Results and Discussion 

Figure 4-17(a) shows the image of a premixed methane-air flame with no mist. Figure 4-17(b) shows 

mist of nominal 0.3 μm drops as they exit the burner nozzle in the absence of the flame. The drops are 

illuminated with a doubled Nd:YAG laser sheet. In Figure 4-17(c), both the illuminated drops and the 

flame image are shown although no mist drops are observed leaving the visible flame zone. The 

inhibition effect of this mist on the flame normalized burning velocity is plotted in Figure 4-18. Details 
Aerosol Flame Suppression Studies 367

of the experiment and results are reported in Reference 93 Also shown are the inhibition effect of N2, 

CF4, and CF3Br. The CF3Br data are modeling results considering full chemistry.

76,94 On a mass basis, 

the inhibition effectiveness of water mist was comparable to that of CF3Br. 

Figure 4-16. Predicted Burning 

Velocity for a Stoichiometric, 

Premixed, Methane-air Flame as a 

Function of Initial Droplet 

Diameter and Water Mass 

Loading. (Reference 90)

89

Figure 4-17. (a) Image of Premixed Methane-air Flame Stabilized on Burner; (b) Image of 

Water Mist (no flame) as it Exits the Burner; (c) Image of Laser-illuminated Water Mist 

and Methane-air Flame Showing Disappearance of the Mist at the Flame Boundary. 

(Reference 93) 

A summary of an evaluation of the thermodynamic properties of the agents investigated in Figure 4-18 is 

listed in Table 4–2. Columns 2 and 3 list the sensible enthalpy, per unit mass and per mole, respectively,

required to raise the temperature of each agent from 300 K to 1600 K.95 On a mass basis, N2 and CF4

should contribute similarly to the inhibition. The sensible enthalpy for water mist, including the heat of 

vaporization at 101 kPa (1 atm), is almost twice that of water vapor and 3.5 times higher than for the same 

mass of N2 or CF4. These predictions are consistent with the experimental observations. The mass of 

inhibitor required to reduce the burning velocity by a fixed amount (20 % was chosen here) is listed in

Column 4 and the molar amount in Column 5. 
368 Aerosol Properties 

Figure 4-18. Normalized 

Burning Velocity Reduction vs. 

Mass of Added Inhibitor for an 

Inhibited Methane-Air Premixed 

Flame. (Reference 93) 

The mass of liquid water needed to reduce the burning velocity by 20 % is one-third that required for N2

or CF4, which is in good agreement with the thermodynamic estimate. The value for water is comparable 

to published measurements for CF3Br. Thus, small water droplets acting thermally exhibit a comparable 

effectiveness on a mass basis in premixed flames as CF3Br with its chemical effect. Modeling prediction 

of the small-droplet limiting behavior is presented in Figure 4-19 along with the experimental results of 

Figure 4-18. The sub-micron mist inhibition results are in excellent agreement with modeling predictions 

of Yang and Kee, requiring no adjustable parameters.93 Water vapor was both observed and predicted to 

be less effective than water aerosols of droplet diameters smaller than ≈ 15 μm. An increase in the 

effectiveness of liquid water is predicted as the droplet size is reduced to ≈ 10 μm. Below this size limit, 

no increase in suppression effectiveness with decreasing droplet size is predicted. This limiting droplet 

size correlates with those droplets that completely evaporate in this flow field and is consistent with the 

observations in Figure 4-17.

Table 4–2. Thermodynamic Quantities for Physical Suppressants Compared to CF3Br. 

(Hf

1600K – Hf300K)

a

Agent kJ/g kJ/mol 

Mass Fraction to Reduce 

Burning Velocity by 20 % 

Mole Fraction to Reduce 

Burning Velocity by 20 % 

N2 1.5 42 0.063 ± 0.004 0.062 ± 0.004 

CF4 1.4 122 0.055 ± 0.004 0.019 ± 0.002 

H2O (vapor) 2.9 53 0.033 ± 0.002 0.050 ± 0.003 

H2O (mist) 5.2 93 0.017 ± 0.001 0.026 ± 0.002 

CF3Br 0.85 126 0.019 b 0.004 b

a

 Calculated from data in Chase et al.95

b

 Data from Noto et al.96
Aerosol Flame Suppression Studies 369

on of position above the burner were recorded for the 

a) and number density vs.

er (generated using the VOAG) were

perature ≈ 2100 K, determined

roplets generated with a 

 evaporated at the flame front. Thus, the

s flame is between 6 μm and 27 μm. 

These measurements are consistent with the predicted nominal 10 multiphase model. 

Figure 4-19. Premixed Methane-air 

Burning Velocity as a Function of Agent 

Mass Fraction, Normalized by the 

Uninhibited Burning Velocity. Water 

data from Figure 4-18 (symbols), 

multiphase flame modeling results for 

humidified flames inhibited with mists of 

experimental droplet number density decay rates are much slower than the 

model predicts; note that the distance plotted in Figure 4-20b is four times that of Figure 4-21b. The

so e

experiments, droplets are traveling at the hase, a nedimensional, freely gatin mod he sam but

this is about 30 cm/s. Thus, the prediction of oration rate (shorter me) for the 

longer residence tim the mo consiste y the effect of larger droplets on

the burning velocity needed. 

T predicts ning-p a f moderate

thane flames) ning velocity as seen in

erivative of burning velocity with respect to the water mass loading is infinite. 

urning points are caused by the droplet dynamics within the flame. Specifically, longer droplet 

residence time in a suppressed flame causes further increase of efficiency in flame suppression. 

Mathematically, at a turning point, an infinitesimally small increase of water mass loading causes the

PDPA measurements for larger droplets as a functi

methane-air premixed flame: diameter vs. position are shown in Figure 4-20(

position in Figure 4-20(b). Droplets larger than 27 μm in diamet

reasonably monodisperse and survived travel through this flame (peak tem

using a radiation-corrected, coated Pt/Pt-Rh thermocouple). The smaller d

nebulizer (average diameter 6 μm, Dv0.9 < 10 μm) completely

experimental droplet size for onset of complete evaporation in thi

μm limiting droplet size of the

indicated drop size (solid lines), and 

PREMIX modeling results for humidified 

flames inhibited with additional water 

vapor (dashed line).93

Modeling predictions for the experimental conditions presented in Figure 4-20 are shown in Figure 4-21. 

Droplet size vs. position is shown in Figure 4-21a, and relative droplet number density vs. position is 

shown in Figure 4-21b. The water mass fraction is 0.05. (The 0 mm position is the 300 K left boundary

for the freely propagating flame solution.) Droplet size behavior is in reasonable agreement in the post 

flame region. However, the

urce of the difference lies in the different residence times of the droplets in the flame. In th

same speed as the gas p

el, droplets also travel at t

bout 140 cm/s. In the o

propa g flame e speed as the gas phase,

a higher evap evaporation ti

e in del is nt. Experiments to quantif

are

he model a tur oint beh vior for the effect o

 on bur

and larger sized droplets (above 

25 μm for stoichiometric, premixed, me Figure 4-16. At the

turning point, the d

T
370 Aerosol Properties 

burning velocity to fall to the lower branch of the suppression curve. Such very-low-burning-velocity

branches are experimentally not achievable due to the various heat-loss mechanisms that would prevent

these flames from stabilizing. Therefore, it may be safe to interpret such turning points as flame

extinction points. 

The theoretical and computational analyses predict that the burning-velocity curves join together in the

lower-burning-velocity region. This behavior can be attributed to a greater droplet residence time in the 

slow-burning flames. If the burning velocity is sufficiently low, any droplet can totally evaporate within

the flame, leading to equivalent efficiency for flame suppression. 

let Diameter and Number Density

ixed Methane-air Flame as a Function of Height above the 

75

Figure 4-20. Experimental Measurements of Water Drop

in a Stoichiometric Prem

Burner Exit.

Figure 4-21. Multi-phase Predictions of Water Mist Droplet Diameter and Number Density

in a Stoichiometric Premixed Methane-air Flame.75
Aerosol Flame Suppression Studies 371

The turning points predicted for larger droplets seen in Figure 4-16 can be important in fire suppression. 

Unfortunately, no direct observation or measurement of such behavior for inhibited premixed flames has 

been reported. The challenges to realizing such experiments include the generation of the monodisperse 

mists, the high water-mass loading required at turning points, and controlling the non-ideal conditions 

such as buoyancy and heat loss to the environment. Experiments that focus on the inhibition effects of

odel predicted temperature, species 

concentration, number density, and droplet history profiles throughout the flame. These variables 

represent the underlying factors that control the overall burning velocity. Therefore, detailed 

Modeling of Counterflow, Non-premixed Flames

m

imm

m

m

chemistry. Stable gas-phase flames could be modeled. The calculation was very sensitive to the 

larger droplets are needed.

Understanding the inhibited flame structure and droplet dynamics provided considerable insight into the

mechanisms of flame suppression. The computational m

measurements of flame structure were extraordinarily valuable. Unfortunately, both creating the flame 

and developing the required diagnostics are significant challenges. For these studies, flame structure, 

burning velocity, and extinction characteristics were modeled as an ideal flat flame. The experiment was

a Bunsen-like configuration that allowed droplets to be introduced through the open tube. The burning

velocity in these flames was inferred from analysis of the flame area. In addition, flame profile data were 

measured, for example, on the centerline. This configuration and modeling approach are valid for submicron droplets that fully evaporate prior to entering the flame. However, for larger droplets or particles,

data from such experiments cannot be compared directly to one-dimensional flat-flame models. The gas 

and particle/droplet velocities, for example, were much greater in the Bunsen-like flame. It appears that 

there are two choices for modeling the inhibition effect of larger particles/droplets in premixed flames:

configure a “flat” premixed, burner-stabilized flame that can be seeded with large water particles/drops, 

or directly model the three-dimensional Bunsen-like flame. The success of uniformly seeding 

particles/droplets into a flat flame is highly unlikely considering the particle/droplet sizes to be studied. 

Success seems more likely for alternative modeling approaches. This area is one that still requires further

development.

i

One of the NGP objectives was to understand the trajectories of condensed-phase, monodisperse aerosols 

in a reacting counterflow field and investigate their effect on the flame extinction condition. Several 

odeling approaches were explored to aid in the understanding of the observed suppression behavior. 

One publicly available code advertised as capable of modeling suppression of counterflow flames by 

aerosols is RUN-1DL (Rogg and Wang, Cambridge, England). Available documentation on the treatment

of droplets was very sparse. The code as delivered only treated a single droplet stream and was not

ediately capable of modeling flame conditions involving an ensemble of water drops. Further 

development and code modification to model the range of NGP experimental conditions were deemed too

time consuming in light of the limited documentation. 

PHOENICS (CHAM, London), a finite volume code for calculating fluid flow, was also considered for 

odeling the interaction of droplets in the counterflow flame. NRL staff had previously used this to 

odel the flow field in this counterflow burner.97 PHOENICS was configured to treat the full flame 

 i The model described here was developed and predictions made before much of the experimental data was available. 

Subsequent experimental work presented later in this chapter benefited greatly from the modeling studies. The experimental 

studies in large part validated the modeling approach and subsequent predictions.
372 Aerosol Properties 

inclusion of the aerosol/flame interaction due to the time and spatial scale dynamic ranges. Further 

development and code modification were not pursued as part of this project. 

Chelliah and co-workers at the University of Virginia developed a two-phase hybrid Eulerian–Lagrangian 

formulation for treating the interaction of droplets and counterflow non-premixed flames.80,98 99 , The 

details of the hybrid Eulerian-Lagrangian formulation and the solution algorithm developed to predict the 

interaction between water mist and a laminar counterflow flame are described in Reference 80. The 

thin the mixing layer. The exit velocity profiles of both air and fuel streams were assumed 

to be of the plug flow type. The droplets were introduced though the upper nozzle at a steady rate with 

oplet velocities at the nozzle exit were assumed 

considered here, the gas velocity can differ from 

temperature, T , for a methane-air flame as a function of the flow strain rate, a, for some selected water 

o 

a ric pressure (i.e., partial pressure of 

model predictions were validated against experimental non-premixed counterflow flame extinction 

conditions and low-pressure flame structure data for both propane and methane flames, as well as indirect 

comparison to experimental methane-air burning velocity results.100 A reduced chemical kinetic 

mechanism was used, which was derived from a much more comprehensive one proposed by Wang and 

co-workers.101,102 The Wang mechanism had been validated based on global flame 

propagation/extinction experiments and flame structure measurements for a variety of fuels.101 For water 

mist suppression of propane or methane counterflow flames, extinction strain rate predictions using both 

mechanisms agreeded within 5 s-1. Both mechanisms predicted a non-premixed propane-air flame 

extinction strain rate of ≈ 580 s-1, in excellent agreement with the experimental measurements. 

For the model, water droplets were introduced into the air stream. The water droplet loadings were small, 

i.e., the mass of liquid water was less than 3 % of the air flow. Thus, the ratio of droplet separation 

distance to droplet diameter was greater than 20, and collisional effects between droplets were neglected. 

The reduced gas-phase chemical kinetics mechanism involed 17 species and 39 elementary reactions with 

variable thermochemical and transport data. The discretized gas phase equations were solved using an 

algorithm that includes Newton and fixed point iteration steps. 

The flow configuration was shown earlier in Figure 4-10. When ignited, a steady laminar flame was 

established wi

the air stream and were assumed to be monodisperse. Dr

to be the same as that of air. (For the two-phase system 

that of droplets.) The response of droplet trajectories for different flow strain rates was addressed by 

varying the nozzle exit velocities of methane and air streams. For plug flow boundary conditions 

(corresponding to viscous, rotational flow formulation), the flow strain rate, a, is defined as the negative 

of the axial velocity gradient of the gas in the air side (i.e., 2dv/dx), just outside the mixing layer 

(Figure 4-10). At flame extinction, the critical value of the flow strain rate was identified here as aext. In 

the experiments, the outer air flow was known to deviate slightly from the plug flow boundary 

conditions,103 but explicit detailed laser Doppler velocimetry data at the boundary can be easily 

implemented in the model. 

Effect of Water Vapor on Flame Extinction 

In a real-life fire situation, or in laboratory experiments, when water droplets are introduced into a dry air 

stream, there will be some evaporation. Consequently, the fraction of oxygen displaced (or diluted) by 

the water vapor can reduce the flame strength, or the extinction strain rate, aext, of the counterflow flame 

considered here. Figure 4-22 shows a comparison of the predicted variation of maximum flame 

max

vapor concentrations. The circle symbol corresponds to a dry case, whereas the star symbol correspond t

fully saturated case at a room temperature of 300 K and atmosphe
Aerosol Flame Suppression Studies 373

water pH2O,sat = 0.0351 atm or, in mass fractions, YH2O,u = 0.0222, with YO2,u = 0.2278 and YN2,u = 0.7500). 

Based on this figure, if the asymptotic value of highest strain rate for each case is taken as the extinction 

strain rate, a , then the results indicate that saturated water vapor alone can decrease the flame extinction 

Inflow

Water Vapor Partial Pressures.80

ext

strain rate by 25 %. This enhanced effect is not solely due to dilution, but is partly due to the high 

thermal capacity of water. 

Figure 4-22. Predicted Comparison

of the Maximum Flame 

Temperature, Tmax, vs. Flow Strain 

Rate, a, for a Dry Flame and for 

Four Flames with Different

Dynamics of Water Droplets without Interaction 

If the amount of water being injected into the air stream is greater than pH2O,sat = 0.0351 atm (or YH2O > 

0.0222), then the excess water will be in condensed form, i.e., in the form of droplets. The mass fraction

of water associated with such droplets leaving the air nozzle can be defined using 

d d d u u

d d d

m n v v

m n v Y

+ ρ =

0 0 0

0 0 0

0 (4-7)

where md0 is the initial mass of the droplet, nd0 the initial number density, vd0 the initial velocity of the 

droplet, and ρu and vu are the density and velocity of air leaving the nozzle, respectively. All the results

shown here are obtained by assuming that the droplet velocity leaving the air nozzle is same as that of the 

gas, i.e., vd0 ≡ vu, but other inflow conditions can be easily analyzed with the model as developed. 

Various monodisperse water mists for three different fixed water mass fractions in air (Y0 = 0.01, 0.02,

he number density of water droplets leaving the air nozzle,

n0 (number/cm ), for three different values of Y0 and for the three different monodisperse droplet sizes

and 0.03) were considered. Table 4–3 shows t

3

selected. 

For all the combinations of water mass fraction in condensed phase and droplet sizes, the ratio of droplet 

separation length to the droplet diameter, l0/d, was fairly large (>20) and was of uniform order. It was 

thus reasonable to neglect droplet–droplet interactions for all the cases considered. Further justification is 
374 Aerosol Properties 

presented below.ii For dilutely loaded droplets, the gas displacement by water droplets can also be

neglected because of the large density ratio between the condensed water and the gas. Under such

conditions, the only interaction between the droplets and the gas is from source terms due to mass, 

momentum and enthalpy transfer per unit volume from the droplets to the gas. (See Equations 2 through 

5 in Reference 80.) When these terms are included in the numerical calculations, the flame structure and 

extinction strain rate are affected dramatically, depending on the amount of condensed water added and

paring the trajectories of various droplet 

s, a fixed, low strain rate case (a = 130 s-1) 

erms turned “off” in the numerics) is

 the profiles of droplet trajectories and source terms 

s well. 

n

Mass Fractions.80

the droplet sizes. Thus, for the purpose of illustrating and com

sizes in the counterflow field and their associated source term

where the droplet source terms are excluded (or the source t

considered first. It should be mentioned here that

presented below remain essentially the same for high-strain rates a

 Table 4–3. Number Density and Droplet Separatio Lengths for Some Selected Water

Yo d (μm) No (cm-3) No

1/3 (cm-1) lo (cm) lo/d 

5 178,000 56.3 0.0177 35

0.01 20 2,790 14.1 0.0711 36

50 178 5.6 0.178 35 

5 360,000 71.2 0.0141 28 

0.02 20 5,630 17.8 0.0562 28 

50 360 7.1 0.141 28 

5 546,000 81.7 0.0122 24 

0.03 20 8,530 20.4 0.0485 24 

50 546 8.2 0.122 24 

Figure 4-23 shows a comparison of the axial velocity along the axis of symmetry of the gas and droplets 

of different sizes for the selected low-strain rate case. Although the coupling arising from the droplet

source terms on gas-phase structure was neglected, in solving equations for droplet trajectory shown in 

Figure 4-23, the droplet vaporization effects were included. (See Equations 18 through 23 in Reference 

0.) In Figure 4-23, the 5 μm droplets followed the gas fairly closely, whereas large droplets deviated

considerably. The droplet lag became slightly worse for high-strain rates, especially near extinction 

8

conditions. The 5 μm droplets were completely vaporized soon after they entered the hot mixing layer,

whereas 50 μm droplets penetrated through the flame and also the stagnation plane and then reversed

their direction somewhere in the fuel stream. When the droplet vaporization was suppressed, then the 

penetration of 50 μm droplets became much more pronounced, with multiple crossings at the stagnation 

plane. These predictions are consistent with those reported earlier by Chen et al. 64

 ii In actual water mist systems, it is highly unlikely that the water droplets generated are monodisperse. However, the predictions 

performed here with monodisperse droplets provide a better mechanism to analyze and understand the basic droplet dynamics 

and interactions with the gas-phase processes, which can be easily extended to polydisperse droplet flows in the future. 
Aerosol Flame Suppression Studies 375

Figure 4-23. Calculated 

Gas Velocity, Velocities for 

Varied Water Droplet 

Diameters, and Gas 

Temperature (thick solid 

line) in a 130 s-1

Methane/air Flame with

Droplet Source Terms 

turned “off.”80

The droplet response to ase flow conditions was characterized by the Stokes 

numbe defining flow resid e in the outer air flow as /a, and deriving an 

expression for the droplet response time (td

for St: 

the variation in gas-ph

r, St. By the ence tim

) based on Stokes drag, the following expression was derived

tf = 1

2

18

a d

St ρ d

μ = (4-8)

by Li et al.

igure 4-25 shows that the temperature of 5 μm droplets (Td) followed that obtained assuming thermal

equilibrium (solid line), whereas other droplets showed a small thermal lag. However, this lag had

significant effect on the flame structure and extinction condition considering the small temperature

difference and the energy associated with it. The variation of flux fraction, F, as a function of the axial 

location, shown in Figure 4-26, however, played a very important role in the droplet source terms (see 

Equation 17 in Reference 80), and therefore the state of the flame. The reduction in F seen as the droplet

approached the stagnation plane was because of the flow divergence or straining effect. 

where μ is the gas viscosity, d is the aerodynamic diameter of the drop and ρd is the drop density. For the 

strain rate of 130 s-1 considered in Figure 4-23 the variation of St for different droplet sizes in the outer

oxidizer flow is shown in Table 4–4. The 5 μm droplets have truly small St number, consistent with the 

results shown in Figure 4-23. As the droplet size approaches about 15 μm, the St number approaches 

about 0.1, indicating slower response of the droplet to the gas, as seen in Figure 4-23. 

Figure 4-24 through Figure 4-26 show the variation of other variables associated with the droplet, 

(evaluated from Equations 20, 21, and 23 in Reference 80), along the axis of symmetry. These results

were obtained under the same conditions as in Figure 4-23. Figure 4-24, for example, shows that Ud was 

always less than U, which is expected, as the gas was accelerating radially from the axis of symmetry and 

is consistent with experimental measurements of methanol droplets in a counterflow 104

F

in
376 Aerosol Properties 

a = 130 s-1.

80 Table 4–4. Stokes Number (St) for Different Droplet Sizes for 

d (μm) td (s-1) tf (s-1 x 102) St

5 0.75 x 10-4 0.77 0.0098

10 0.30 x 10-3 0.77 0.039

15 0.68 x 10-3 0.77 0.089

20 0.12 x 10-2 0.77 0.16

25 0.19 x 10-2 0.77 0.25

30 0.27 x 10-2 0.77 0.35

40 0.48 x 10-2 0.77 0.63

50 0.75 x 10-2 0.77 0.99

60 0.108 x 10-1 0.77 1.42

If the droplet vaporization were neglected, the curves shown for different droplet sizes would shift 

somewhat, but the dramatic reduction of F seen with droplet location does not change. Similar results for 

axial liquid methanol volume flux were observed experimentally by Li et al.104 with and without a flame. 

ure.

Figure 4-24. Comparison of 

U of Gas and Ud of Different 

Droplet Diameters, with

Droplet Source Terms 

Turned “off.”80 Thick line: 

gas temperat
Aerosol Flame Suppression Studies 377

Figure 4-25. Comparison of 

Droplet T rature, Td, of 

Different Droplet Diameters, 

with Droplet Source Terms 

Turned “o Thick line: 

gas temperature.

f

F, of 

ifferent Droplet Diameters,

ith Droplet Source Terms 

urned “off.”80 Thick line: 

 

CE TERMS WITHOUT INTERACTION. Once the solution of droplet variables was obtained

.e., d, vd, Ud, Td, and F), the source terms contributing to the gas phase were evaluated. (See Equation 

17 in Reference 80.) Here, such contributions were shown, once again by turning “off” the interaction of

such source terms with the gas-phase calculations to keep the flame structure condition the same for

empe

ff.”80

Figure 4-26. Comparison o

Flux Fraction Function, 

D

w

T

gas temperature.

SOUR

(i
378 Aerosol Properties 

different droplet sizes considered further in the next section. Figure 4-27 t

variation of mass, momentum, and energy source terms, i.e., Sm, SU -

symmetry for different droplet sizes. Figure 4-27 shows apparently

side before the flame, where evaporation effects were negligible. 

greater radial force per unit volume on the gas than larg

the fact that larger droplets shoot through that region faster than smal

hrough Figure 4-29 show the 

USm, and Sh 2 hiSm, along the axis of

 strange profiles, especially on the air

 It is seen that smaller droplets exerted a

er droplets. This is, however, easily explained by

le

shorter residence times. Comparison of Figure 4-28 and Figure 4-29

Sh - hiSm to be very similar (except for the change of the sign). This suggested that heat transfer to the

droplets and evaporation were well correlated; the implications of such

in the next section. 

EFFECT OF WATER DROPLETS ON FLAME EXTINCTION. An important question that was 

probed by the two-phase model developed here was the effect of water roplets (i.e., their size and mass

loading) on the flame extinction condition, with the source terms turned “on.” Such effects were shown 

by plotting the variation of maximum flame temperature as a function of the flow strain rate, as shown in 

Figure 4-30. Figure 4-30 also shows the effect of water vapor. Profiles for Tmax are plotted for a single 

mass fraction (Y0) of 0.01 for different droplet sizes. For comparison, the saturated water vapor at room 

conditions is shown as a solid line. 

ndicates that 50 μm droplets were the least effective of those sizes studied, with a relatively

igh extinction strain rate. They were still more effective than air saturated with water vapor (mass 

the lowest extinction strain rate. As 

m, they became less effective. The

m approached a δ-function, as seen

in Figure 4-27 and Figure 4-28, introducing severe convergence problems of the numerical model. 

Fi

of ,

wi plet

Source Terms Turned “off.” 

Thick line: gas temperature. 80

r droplets, so that they have much

indicates the profiles of Sm and

correlations are discussed further

d

Figure 4-30 i

h

fraction of 0.022). The 15 μm droplets were the most effective, with 

the diameter of monodisperse droplets was decreased below 15 μ

lowest size considered was 5 μm, as the mass and enthalpy source ter s

gure 4-27. Typical Sm Profiles

Different Droplet Diameters

th Y0 = 0.02 and Dro
Aerosol Flame Suppression Studies 379

Figure 4-28. Typical SU - USm Profiles 

of Different Droplet Diameters, with 

Y0 = 0.02 and Droplet Source Terms 

Turned “off.” Thick line: gas 

temperature. 80

line: g

.

strain rate a, for Different Droplet 

Diameters, with water mass fraction = 

0.01. 80

Figure 4-29. Typical Sh - hiSm 

Profiles of Different Droplet

Diameters, with Y0 = 0.02 and Droplet

Source Terms Turned “off.” 80 Thick 

as temperature.

Figure 4-30. Comparison of Tmax vs
380 Aerosol Properties 

In fact, droplets of 5 μm and below should approach the lim

vapor phase at the air side boundary of the mixing laye

affected. Irrespective of this difficulty, this most interesting nonm

captured well by the two-phase model developed here

it where all the water considered is in the 

r, provided the inflow temperature of air is not 

onotonic effect of droplet diameter was 

 for stead

flames. 

For higher water mass fractions in the condensed phase, s a

function of flow strain rate, as shown in Figure 4-31 and Figu eased 

significantly with increasing water droplet loading. 

Figure 4-31. Comparison of Tmax vs. 

train rate a, for Different Droplet 

iameters, with water mass fraction = 

.02.

 80

Figure 4-32. Comparison of Tmax vs. 

train rate a , for Different Droplet

y, laminar, counterflow, nonpremixed

imilar profiles were obtained for Tmax as

re 4-32. The extinction strain rate decr

s

D

0

s

Diameters, with water mass fraction = 

0.03. 80
Aerosol Flame Suppression Studies 381

For the mass fractions considered here, the nonmonotonic effect of droplet diameter on the extinction

strain rate can be better illustrated by Figure 4-33. For all the conditions considered here, 15 μm to

20 μm droplets were the most effective.

Figure 4-33. Comparison of aext vs. 

Droplet Diameter for Different Droplet 

Mass Fractions in the Condensed Phase 

at Inflow.

 80

 be explained by the location of the mass

m droplets, the peak value of Sm, and also of Sh - hiSm, 

yer. The effect of Sh - hiSm on flame 

and hence early flame extinction.105,106 The exact

me extinction condition in the presence

f water droplets are being investigated separately, with comparison to that of traditional chemical agents, 

such as halon 1301. 

The concept of flow strain rate as defined here, i.e., a = -dv/dx, is strictly applicable to constant density or 

ul fluid dynamical parameter for comparison with 

ned with nonintrusive laser Doppler velocimetry

d dynamical parameter is the diffusion time scale

This superior effectiveness of 15 μm to 20 μm droplets can

source term seen in Figure 4-27. In the case of 15 μ

occurs at the oxygen consumption or radical production la

temperature resulted in lower radical production 

physical, thermal, and chemical contributions responsible for fla

o

nonreactive mixing layer flows. However, it is a usef

near extinction experimental velocity data obtai

techniques.103

 For theoretical studies, a more useful flui

defined by the inverse of the scalar dissipation rate:106,107

2

χ = 2D ∇Z (4-

λ/(ρc

9)

where D is an appropriate diffusion coefficient (e.g., D = the

mixture fraction. For small stoichiometric mixture fractio

stoichiometric scalar dissipation rate, χs, with a taking int u t the thermal expansion effects, has

been derived by Kim and Williams.108 Instead of using such an analytical expression, the scalar 

dissipation rate at the stoichiometric point at extinction was directly computed here by solving for a

conservation equation for mixture fraction. 

The resulting variation of (χs)ext as a function of the initial droplet size and water mass fraction in 

condensed phase is shown in Figure 4-34, indicating a trend similar to that of aext shown in Figure 4-33. 

p) for unity Lewis number) and Z is

ns, Zs, an analytical expression relating the

o acco n
382 Aerosol Properties 

Figure 4-34. Comparison of (χs)ext vs. 

Droplet Diameter for Different Droplet 

Mass Fractions in Condensed Phase 

at Inflow.

 80

reful experimentation. Furthermore, the generation 

and methods of extending the present modeling work 

seen in Table 4–3, the fraction of droplets colliding before evaporation can be 

neglected. 

sidered here, as seen in Figure 4-23. 

igure 4-35 shows the calculated trajectory of 50 μm droplets along the x-axis as a function of tim

s roughly estimated as 

The results presented here need to be validated with ca

of truly monodisperse droplets is a challenging task, 

to polydisperse droplets also need to be pursued. 

DROPLET-DROPLET COLLISIONS. A key assumption in the numerical model developed was that

droplets are transported without collision with other droplets. This assumption allowed for a fairly simple 

description of the droplet flow field and derivation of a self-similar solution for both droplets and the gas

phase. It relied on the idea that the droplets leaving the air nozzle were monodisperse with the same exit

velocity (nd0) and, hence, similar dynamics in the counterflow field. Although such ideal conditions are 

difficult to accomplish in experimental situations, if the initial droplet separation distance to droplet

diameter is large, as

It is valuable to test the validity of this assumption under the worst case for this type of laboratory

configuration. This occurs when large droplets at high droplet loading penetrate the stagnation plane and 

reverse their trajectory, creating a region where opposed flow of droplets exists near the axis of 

symmetry. Such a region existed for 50 μm droplets con

F e and

the corresponding flux fraction function for Y0 = 0.03. With vd0 ≈ 60 cm/s and an average distance l0 = 

0.12 cm (see Table 4–3), each droplet was estimated to be preceded and followed by a droplet before and

after a time 1.8 x 10-3 s. Thus, assuming that droplets were leaving the air nozzle at every 1.8 x 10-3 s,

then the region between points 1 and 2 (the region where each droplet group would see a group of

droplets flowing in the opposing direction) in Figure 4-35 consists of ≈ 17 layers. By taking the average

value of droplet flux fraction, F, in region 1 → 2 as 1/2 and in region 2 → 3 as 1/4, the probability of

droplet collision in this worst case region wa
Aerosol Flame Suppression Studies 383

0.4%

4

1

2

1 17

2

≈ ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

<

o

coll l

d P (4-10)

This estimate for collision probability was even smaller for smaller droplets because of evaporation and 

the increased straining of droplet flux. Hence, the assu ption of negligible droplet collision was 

reasonable. 

Several other simplifications and approximations were als

model. The implication of the approximations introduced

approximations were addressed in detail in Reference 80. 

Y FOR MODELING AEROSOL SUPPRESSANTS IN COUNTERFLOW FLAMES.

This work developed a new two-phase numerical model to analyze the effectiveness of monodisperse, 

ber density

equation. The model was applied to a case where monodisperse water droplets were introduced with the 

ed to be the most effective. Addition 

of 3 % of water by mass in condensed phase (in addition to 2.22 % as water vapor due to saturation) was 

shown to reduce the extinction strain rate to 130 s-1 from ≈ 400 s-1 for the dry air case. The nonmonotonic 

m

o used in the development of the analytical

and possible inaccuracies arising from these

Figure 4-35. Calculated Droplet 

Trajectory as a Function of Time, for 

d = 50 μm and Y0 = 0.03. 80

SUMMAR

condensed phase fire suppressants in extinguishing counterflow flames. The model overcame the 

singularities associated with solving the droplet number density equation in a counterflow field accurately

and in a very robust manner. This was achieved using a hybrid Eulerian–Lagrangian formulation for the

gas-droplet flow, with the introduction of droplet flux fraction to describe the droplet num

air stream to a nonpremixed counterflow methane/air flame. Several cases with droplet diameters ranging

from 5 μm to 50 μm and mass loadings of liquid water in the air stream from 0.01 to 0.03 were

considered. The model predicted that small droplets (< 20 μm) tended to follow the gas closely and went

through rapid evaporation in the hot mixing layer, never crossing the stagnation plane. The 50 μm droplet 

crossed the stagnation plane several times before being completely vaporized. The important finding was

that the flame extinction strain rate showed nonmonotonic behavior for different monodisperse droplet 

sizes considered for several water mass loadings. Assuming that the droplets were introduced with the air

stream at the same velocity as the gas, 15 μm droplets were predict
384 Aerosol Properties 

ppression effect of droplet size favoring 15 μm droplets was attributed to the droplet dynamics in the

counterflow field and to the large mass evaporation and heat sink o served near the oxygen consu

layer. Sensitivity analyses of the approximate analytical models used to simplify the numerical

NRL researchers quantified water mist suppression effectiveness in non-premixed flames by determining 

109 A schematic of the experimental 

su

b mption

integrations showed that mass evaporation and heat transfer terms needed to be more accurately modeled.

Methods of improving the accuracy of these models were left for future pursuits.

Non-Premixed Counterflow Flame Studies 

Experimental Setup 

the extinction strain rate in counterflow propane/air flames.

configuration is shown in Figure 4–36. 

Figure 4-36. Apparatus for Aerosol-inhibited, Non-premixed Counterflow Flames.

75

The burner consisted of two 50 cm long stainless steel tubes (1 cm inner diameter), with outer concentric 

tubes for a co-flow. The upper tube had a jacket for cooling water. The two tubes were aligned vertically

and collinearly with a separation of 1.00 cm ± 0.05 cm. Air or fuel exited from a single tube with a 

parabolic velocity profile. Flow straighteners were not used. The air and fuel velocity profiles near the

tube exits when opposing each other were flattened slightly from a parabolic profile due to the presence 
Aerosol Flame Suppression Studies 385

of the other opposing flow. In the configuration used for the experiments, the luminous flame zone was 

fairly flat. The end of the each tube was plumbed through the center of stainless steel plates on the top 

and bottom of a 22 cm inner diameter, 19 cm long acrylic tube. The acrylic tube had flat windows for 

good optical access. These windows were critical to both the position of the laser beam overlap for the 

aerosol velocity and sizing measurements and the detection angle of the scattered light. A ≈ 4 SLPM 

purge flow of nitrogen was introduced into the bottom of the chamber. All gases exited through the top 

 diameter exhaust port. Representative centerline 

temperature and axial velocity profiles are indicated in the “blow-out” section of Figure 4–36, as is the 

us reactants in this configuration, a relationship between the local strain rate (i.e., 

the maximum gradient of axial velocity on the air side of the reaction zone), the burner gap size distance, 

 air stream. For a condensed phase agent, the situation became 

more complicated, because under some conditions the condensed phase comprised a substantial fraction 

of the total momentum of the flow, and did not have the same velocity as the surrounding gas. In 

determining strain rates in the present study, it was assumed that the gas flow field was unchanged by the 

presence of the water mist. This approximation is only valid if the mass fraction of water in the air stream 

is small. For larger water mass fractions, the gas flow field must be measured in the presence of the water 

drops. The velocities of the water droplets themselves could not be used to determine the gas flow 

velocity, however, because they were too large to follow the gas flow field. For these reasons, the water 

mass fractions in the air stream were no larger than 0.03. 

Liquid mist aerosols were generated and seeded in the air flow in the lower tube. Air side introduction of 

the mist is relevant to real fire, total flooding fire suppression scenarios. Fuel was introduced from the 

top, water-cooled tube. Flames were started using an arc igniter for a standard fuel and air flow 

configuration and stabilized in a 1 cm gap between the two tubes. For hydrocarbon fuels burning in air, 

stoichiometry dictates that the non-premixed flame reside on the air side of the stagnation plane. Phase 

Doppler Particle Anemometry was used to measure droplet size distribution, number density, and droplet 

velocities through the flow field by vertically translating the burner assembly. The aerosol generator 

required a ≈ 1 L/min to 1.5 L/min flow of air for proper operation. In order to examine flames of lower 

strain rate, part of the air stream containing droplets was diverted from the burner. Droplets from the 

excess stream were removed by filtering, and the resulting air flow was measured to determine the 

amount of air sent to the burner. 

The mists were produced using a vibrating orifice aerosol generator (TSI Inc. Model 3450), based on the 

design of Berglund and Liu.111 A schematic of the droplet generator is shown in Figure 4-13. Water was 

forced through a pinhole that was acoustically excited by a piezoelectric ceramic. At specific resonant 

frequencies, the water jet broke up into a monodisperse droplet stream. This stream exited the generator 

rough a hole in the dispersion cap. By forcing air to exit through this same hole, the droplet stream was 

disp be. 

Measurements of droplet number density as a function of radial position at the tube exit indicated that the 

plate of the combustion chamber through a 5 cm

strain rate. For gaseo

reactant velocities and densities were determined.110 This burner-specific relationship was used to 

calculate local strain rates, and had been previously shown to remain valid for the addition of gaseous 

agents of high molecular weight to the

th

ersed into a cloud as it exited the droplet generator and entered the counterflow burner’s bottom tu

droplets were evenly distributed except near the tube wall. The mass flow of water was adjusted 

primarily by controlling the backing pressure of water entering the orifice. The use of the small orifice in 

this type of droplet generator placed limitations on the achievable water flow. In general, the smaller the 

orifice used, the smaller the maximum flow of water that could be obtained. 
386 Aerosol Properties 

To obtain flame conditions of high strain and low water mass fraction, the aerosol was mixed with a 

secondary (dry) air stream. The mixture was then introduced up the tube toward the reaction zone. For 

low strain rate conditions, the air flow through the droplet generator required to accomplish droplet 

dispersal often exceeded the desired total air flow to the burner. Under these conditions, a portion of the 

air/droplet stream was diverted to bypass the burner. The analysis below assumes that no collisions 

between droplets occurred. The volume fraction occupied by droplets for these operating conditions was 

typically ≈ 2 x 10-5. The assumption that a significant number of collisions did not occur is consistent 

with the droplet size histograms, which did not show appreciable droplet growth, as would be the case if 

collisions were to occur. 

The aerosol generator used in the present study produced droplets having very narrow size distributions. 

When the dispersion air is used, 95 % of the droplets had diameters within + 2.5 μm of the mean. (The 

size distribution was even narrower in the absence of dispersion air.) 

Droplet size and velocity distributions were monitored using a Phase Doppler Particle Anemometer 

(PDPA-Dantec Measurement Technology). For the mists studied, all droplets in the original droplet 

stream were characterized by a single 4 μm wide PDPA bin size. In order to distribute the stream into a 

mist, a dispersion cap and entrainment assembly were employed. Mist dispersion and air entrainment of 

the highly-monodisperse droplet stream into a mist in the burner tube degraded the monodispersity 

slightly. As shown in Table 4–5, > 50 % of the water mass for each mist was still contained in three 4 μm 

wide bins, centered on the title bin. For the 14 μm mist, this value was 90 %. 

Table 4–5. PDPA Derived Diameter Information for the Experimental Mists Studied 

Measured at 2 mm from the Air Tube Exit. 113

Mist Name 

% Water Droplets within 

± 2 μm of Nominal 

Diameter 

% Water Mass within 

± 2 μm of Nominal 

Diameter 

% Water Mass within 

± 6 μm of Nominal 

Diameter 

14 μm 63 47 90 

30 μm 61 60 72 

42 μm 53 38 51 

One disadvantage of this generator was that the droplet diameter could not be continuously adjusted: 

monodisperse droplets could only be produced at specific sizes, corresponding to piezoelectric driver 

frequencies that match acoustic resonances of the orifice. Using a 5 μm diameter pinhole, for example, 

on zone. 

monodisperse droplet streams with a size distribution peak at diameters of 14 μm, 18 μm, and 24 μm 

were produced. With a 10 μm diameter pinhole, monodisperse streams of 25 μm, 30 μm, and 37 μm 

droplets were obtained. If the piezoelectric driver frequency were nonresonant, a bimodal or multimodal 

droplet size distribution was generally produced. 

Using the PDPA, droplet diameters, axial velocities, and number densities were measured at discrete 

points at a single point in the flame by recording each droplet which crossed the laser probe volume 

during a specified time period. Laser light scattered by the droplets was collected through a window 

mounted in the Plexiglas chamber. The burner was mounted on a three-axis translation stage, such that 

the laser probe volume could be positioned anywhere in the gap between the opposed tubes to record 

droplet characteristics as a function of position. The axial position of the flame was determined by 

centering the PDPA probe volume in the middle of the flame’s visible emissi
Aerosol Flame Suppression Studies 387

The droplet size distribution was monitored during the experiments to ensure that the piezoelectric driving 

frequency was correctly chosen to yield a monodisperse droplet distribution. The droplet volume density

determined by the PDPA was the primary determination of the amount of liquid water delivered to the 

flame. Comparison runs between the PDPA determination of the water delivery rate, and direct

measurements of the accumulated mass of water exiting the droplet generator, yielded agreement within a

few percent. The PDPA system was capable of acquiring data at higher droplet loadings than those 

reported here. This was verified in the case of the larger droplet sizes (the achievable mass fraction of the

smaller droplets was limited by the generator). The limit imposed on the water mass fraction was due to

the effect of the water droplets on the strain rate, not a limitation of the PDPA diagnostic. 

distribution, in these propane/air

counterflow flames, of initially monodisperse water mists of 30 μm and 18 μm, respectively. The figures

8 μm droplets dominating

the size distribution. 

Figure 4-37. Droplet

Distribution Evolution of a 30 μm 

Droplet Behavior 

Figure 4-37 and Figure 4-38 show the evolution of the droplet size

plot number densities of droplets in various size ranges as a function of axial position (x) along the 

burner’s axis (r = 0 mm). The local axial strain rate, a, imposed on the flames corresponds to

approximately 30 % of the extinction strain rate measured in the present apparatus110

 for the uninhibited 

flame (aext = 608 s-1± 65 s-1). The air and droplets exited the lower tube at x = 0 mm; the propane exited

the upper tube at x = 10 mm. The luminous zones of the flames were located at x = 5.0 mm and 4.5 mm, 

respectively, in the experiments employing 30 μm and 18 μm mists. For both initial sizes, the diameter of

the droplets changed very little until the flame was reached, with the 30 μm or 1

Diameter 

Mist in a 170 s–1 Strain Rate 

Flame; the luminous flame is 

centered at 5 mm. 109
388 Aerosol Properties 

Figure 4-38 Droplet Diameter 

Distribution Evolution of an 18 μm 

Mist in a 170 s-1 Strain Rate 

Propane/air Counterflow Flame. 

The luminous flame is centered at 

4.5 mm. 109

reached. For the 30 μm initial size, the total number density de s dramatically, in 

passing through the flame zone. The droplets that were detected in or beyond the flame zone had a broad 

size distribution and a much smaller average size than did the incident drops. These observations indicate 

that, for this flame condition, the 18 μm mist underwent essentially complete vaporization once it entered 

the reaction zone, while the 30 μm mist appeared to be near the th shold size above which droplets were 

not completely evaporated. When incident droplets of 42 μm diameter were used, a larger number of 

droplets were detected beyond the flame zone, supporting the assertion that 30 μm is close to the 

minimum size capable of just penetrating the flame. 

Figure 4-39 and Figure 4-40 focus specifically on the number de

droplets respectively. Number density is plotted vs. axial position first 

increases with axial position; then quickly drops in the flame regio

Three effects combine to explain the shape of the number density main effect is related to 

the velocity profiles of the 30 μm and 18 μm drops, provid 4-39 and Figure 4-40, 

respectively. At the lower tube exit, the droplets had roughly the same velocity as the gas stream. As the 

gas stream’s axial velocity changed in the counterflow field, the equilibrium in velocity between the 

liquid and gas phases was lost, and the drag forces acted to rees blish it. The droplet velocity profile 

therefore followed that of the gases; the velocity initially dropped as the gases moved towards the 

stagnation plane. It then increased when the hot gases expanded n zone, before it dropped 

down again, close to the stagnation plane. In regions where the droplets were decelerating, faster droplets 

caught up to slower ones, and the number density tended to increase. In the flame region, the droplets 

accelerated, which tended to reduce their number density. Figure 4-39 and Figure 4-40 show that the 

impact of axial velocity gradients on droplet number density was significant, with variations in number 

density well correlated with variations in velocity. 

In the flame region, the two droplet sizes showed somewhat different behaviors. In both cases, the 

droplets evaporated; and the total number density of droplets, summed over all size ranges, decreased. 

For the 18 μm initial droplet size, virtually no droplets of any size were detected once the flame was 

creased, though not a

re

nsity profiles of the 30 μm and 18 μm 

. In both experiments, the density 

n. 

 profiles. The 

ed in Figure

ta

 in the reactio
Aerosol Flame Suppression Studies 389

Furthermore, as the air exited the lower tube, the fl

field, producing radial drag forces on the mist. Du

burner axis. The divergence of the air flow therefore acted to redu

centerline. The third effect is evaporation in the flame region, w

and thus also contributed to the decrease in the number densities of the 30 

ow streamlines began to diverge in the counterflow 

e to this effect, the droplets moved away from the

ce the droplet number density along the 

hich caused the droplet size to decrease, 

μm and 18 μm droplets. 

ber Density (triangles) 

and Velocity (circles) of

30 μm Droplets vs. Location 

in a 170 s –1 Strain Rate

Propane/air/30 μm Mist 

Counterflow Flame. 109

Figure 4-40. Profiles of 

in a Propane/air/water mist 

109

m the PDPA software by summing over all droplets 

detected within the measurement time, with the sum weighted by the velocity of each drop. The droplet 

Figure 4-39. Profiles of 

Num

Number Density (triangles) 

and Velocity (circles) of

18 μm Droplets vs. Location 

Counterflow Flame. 

When the data of Figure 4-39 and Figure 4-40 are plotted in terms of droplet flux rather than number 

density, the peak just before the reaction zone is not present. Figure 4-41 shows the droplet flux profile 

for the 30 μm mist. The flux was determined fro
390 Aerosol Properties 

flux decreased slightly as the flame was approached, under the effects of the diverging flow and 

evaporation. The flux then decreased dramatically in passing through the high temperature zone. For the

18 μm incident droplet size, the plot of flux vs. position is qualitatively similar to that of the 30 μm drops,

except that the flux is reduced to essentially zero once the reaction zone is reached. In the plots of droplet 

ainties, in the PDPA concentration 

the flame over the course of the

flux, scatter in the data was attributed in part to experimental uncert

measurements in particular. Slight variations in the position of

experiment also contributed to the scatter. 

Figure 4-41. Droplet Flux Profile for 

a 30 μm

Propan

rain rates of non-premixed propane/air 

 droplet sizes of 14 μm, 30 μm, and

air 

tube exit, at strain rates within 10 % of extinction. Local strain rates w s

as discussed above. The results are plotted in Figure 4-42. For comp f

halon 1301 (CF3Br) vs. local strain rate in the propane/air counterfl e

42 μm water droplets were clearly much less effective in reducing the he

14 μm or 30 μm droplets. The 14 μm droplets were slightly more eff

comparison was difficult because only a very limited mass fraction of water was obtainable in the 14 μm 

droplet size with the present generator. Both of the smaller droplet sizes appeared to be more effective in

reducing the extinction strain rate than was halon 1301. 

Mist in a 170 s-1 Strain Rate 

e/air Flame.

Flame Extinction 

The core data set comprised measurements of the extinction st

counterflow flames at various water mass fractions (up to 0.03) at

42 μm. Droplet fluxes were measured by the PDPA instrument along the burner axis 2.0 mm from the

ere determined from the gas flow

arison, extinction mass fractions o

ow flame

110

 are also plotted. Th

extinction strain rate than were t

ective than the 30 μm, but direct 
Aerosol Flame Suppression Studies 391

Figure 4-42. Water Mist Effects on the Extinction of Non-premixed, Counterflow

Propane/air Flames: (a) Flame Extinction Strain Rate as a Function of Mass of Water or 

Halon 1301 in the Air Flow; (b) Mass Fraction of Water Needed to Extinguish the Flames 

as a Function of Droplet Diameter, Compared to Mass of Gas-phase Halon 1301. 75

Maximum Droplet Evaporation Diameter

Li, Libby, and Williams

104 performed both numerical and analytical modeling, as well as experimental 

measurements, of the behavior of methanol droplets in opposed flows, both in the presence and absence 

me. Clearly, there are differences between the interaction of a fuel droplet with a flame and that of

suppressant droplet. Nevertheless, the effects of evaporation and of viscous drag imparted by the local 

analogies between the two cases. Li et al.104 analyzed the phenomenon of "pushback"

 Chen et al.64 in which droplets in certain size ranges exhibited oscillatory motion 

of a fla

a

gas flow field have

previously described by

in the vicinity of the stagnation plane. For a given flow field, large droplets tended to oscillate, while 

small droplets asymptotically approached an equilibrium position slightly below the stagnation plane. 

Under the assumptions of the Stokes drag law, neglect of evaporation, and an axial strain rate independent 

of axial position, the minimum droplet radius for oscillation to occur is given by 

2

1

min 8

9

⎟

⎟

⎠

⎞ ⎜

⎜

⎝

⎛ = K

R

ρ

μ (4-11)

where μ is the absolute viscosity of the surrounding gas, ρ is the droplet density, and K is the axial strain

rate. For a water droplet in air at a strain rate of 150 s-1, the minimum diameter for oscillation to occur 

was approximately 30 μm. In practice, this threshold represents a lower bound because all droplets lose 

mass by evaporation, and water droplets with initial sizes smaller than the oscillation threshold usually 

ndergo complete vaporization upon entering the flame zone.80 The size ranges of droplets investigated 

in the present study bracketed the threshold size for oscillation for the flow field conditions. Furthermore,

the threshold size was similar to the size predicted by Lentati and Chelliah to be most effective at

u
392 Aerosol Properties 

suppression, although this could be a consequence of the relative importance of viscous drag and 

evaporation for water in particular. 

The experimental results were consistent with the numerical predictions of Lentati and Chelliah80 for 

monodisperse droplet streams in a number of respects. For a methane/air counterflow flame with a local 

strain rate of 130 s-1, Lentati and Chelliah80 predicted that a 30 μm water droplet should just pass through 

the reaction zone before evaporating completely, while droplets having diameters 15 μm or less should 

completely evaporate before reaching the location of maximum temperature. The data shown in

Figure 4-37 and Figure 4-38 were taken with a different fuel (propane vs. methane) and at a slightly 

higher strain rate (165 s-1 to 175 s-1 vs. 130 s-1), but show the same qualitative behavior as a function of

droplet size. Furthermore, Lentati and Chelliah80 predicted that appropriately sized water droplets could 

be more effective on a mass basis than halon 1301 in suppressing combustion, but that the effectiveness 

of water was likely to decrease significantly with increasing droplet size for diameters >30 μm. Both of

these predictions were consistent with the data presented in Figure 4-42. There were approximations

made in the modeling as well as non-ideality in the present experiment. Nevertheless, the results of these

experiments supported the predictions of Lentati and Chelliah80 in those aspects for which there was an

dequate comparison. 

The d n the 

r

d

Solu l can 

premixed and opposed flame configurations for the onset of 

e temperature field, droplets begin to evaporate at the lower

of the strain rate.80 For a

rain rate of 450 s

-1 (high strain rate for methane, mid strain rate for propane) the residence time was 

 2.2 ms. The estimated droplet size for the onset of comp te evaporation was ≈ 25 μm. This value 

agreed with the experimental water droplet size survival properties in propane-air flames

63 and was

flame suppression model further. 

a

egree of evaporation (or decomposition) of an aerosol particle or droplet in a flame depends o

esidence time of the aerosol in the high temperature environment. In the flame, the evaporation rate and

roplet size are spatially dependent since the aerosol is moving in a non-uniform temperature field.

tion of the problem requires full consideration of the flow field. However, the evaporation mode

be used to estimate a droplet size in the

complete droplet evaporation. In the flam

temperature and the resulting smaller drop, traveling into the higher temperature region, will experience a 

higher evaporation rate and consequently a shorter total evaporation time. 

Consideration of the burning velocity and flame thickness of the methane-air premixed flames reported

earlier in this section suggested that the water droplets should experience a residence time in the range of

1 ms to 2 ms. Consideration of the temperature field and the results of Figure 4-6 indicated a droplet size 

limit for complete evaporation (and the corresponding onset of water droplet limiting effectiveness) at ~ 

15 μm in premixed flames. This upper limit estimate was consistent with the water droplet size profiles 

observed in the premixed flame and with predictions of the multiphase model. 

For counterflow flames, transit time in the flame was estimated from the inverse

st

≈ le

consistent with the modeling predictions for methane-air flames.80

Modeling the Effect of Water Vapor 

NRL staff collaborated with Chelliah and co-workers to understand the observed suppression behavior of

water vapor and water droplets on propane-air non-premixed counterflow flames. Comparison of the

predictions of the model described in Section 3.5.1 with experimental results provided valuable insights 

regarding the experimental observations.112 The collaboration also served to validate the multi-phase 
Aerosol Flame Suppression Studies 393

During the application of water mist to a fire or a laboratory flame, an enclosed or semi-enclosed chamber 

will be saturated with water vapor. The time required to attain the saturation condition depends on 

several factors, including the air temperature, original humidity level, and the total surface area of the 

 in the figure are flame temperature vs. strain rate curves for flames with dry 300 K 

air, saturated 300 K air, and saturated 300 K air containing a 14 μm, 30 μm, or 42 μm diameter 

droplets applied. For the same condensed water mass, smaller droplets have a higher surface area per unit 

volume (or mass), leading to a more rapid saturation of the initially dry air, while larger droplets with 

lower surface area/volume are expected to take considerably longer. Consequently, different levels of 

water vapor content in air are realizable (i.e., saturated to partially saturated), depending on the droplet 

size, droplet number density or mass added, gas temperature, and the elapsed time. 

The model assumed that the room temperature was constant at 300 K and that the air feed stream was 

completely saturated with water vapor. The latter implied that droplets traveled with the air stream long 

enough to attain equilibrium conditions. At 300 K, the mole fraction of water vapor in saturated air is 

0.0352; the mass fraction is 0.0224. This corresponds to mole fractions of oxygen of 0.2026, nitrogen of 

0.762, and water of 0.0352 in the oxidizer stream. 

Results of the two-phase model for non-premixed counterflow propane-air flames are shown in 

Figure 4-43. Indicated

monodisperse water mist. The water masses for the individual aerosols were based on the experimental 

determinations shown in Table 4–5. The extinction condition is defined by the curve turning point; the 

value of the strain rate at the turning point is the predicted extinction strain rate for each suppressed flame 

condition. 

Figure 4-43. Variation of the Predicted Flame Temperature as a Function of the Flow 

Strain Rate for Propane-air, Non-premixed Counterflow Flames for Dry Air, Air Saturated 

with Water Vapor, and Saturated Air with Monodisperse Water Droplets. 113
394 Aerosol Properties 

Modeling the Effects of Condensed Water and Droplet Size 

he methane-air system. Unlike the methane-air

counterflow system where the optimal droplet size was predicted to be between 15 μm and 20 μm, the

A comparison of the modeling and experimental results is presented in Figure 4-44. 

The reduction of oxygen mole fraction from 0.209 (dry air) to 0.2026 (saturated air) due to displacement 

by water vapor at 300 K yields a decrease in the predicted flame extinction strain rate from 584 s-1 to 

458 s-1, a 22 % reduction in the extinction condition, as shown in Figure 4-43. At elevated oxidizer 

stream temperatures, assuming complete saturation, the increase in partial pressure of water vapor and the 

resulting decrease in oxygen concentration would lead to an even lower flame extinction strain rate. 

Likewise, a lower level of humidification in the incoming air stream would result in a smaller reduction in 

the extinction strain rate than that of the completely saturated air condition. 

When condensed water droplets were added to an already saturated air stream, the interaction of droplets 

with the flame front reduced the flame extinction condition further, as seen in Figure 4-44. However, the 

degree of reduction of the flame extinction condition depended on the droplet size, consistent with 

previous modeling and experiments involving t 80

present predictions and experiments indicated that the optimal droplet size for the propane-air 

counterflow system was below 15 μm. The difference in optimal size was related to the higher flame

extinction condition of propane-air flames, about 50 % greater than that of methane-air counterflow

flames. The lower flow residence times at higher strain rates provided less time for a given droplet size to

completely evaporate. For higher strain rates, droplets need to be correspondingly smaller to optimally

interact with the flame front. 

Figure 4-44. Measured Extinction Strain Rate for Water Mist Suppressed Propane-air,

Non-premixed Counterflow Flames vs. % Mass of Condensed-phase Water in the Air109

and Calculated Results for 14 μm, 30 μm, and 42 μm Diameter Monodisperse Water Mists 

in Saturated Air at 300 K. 113
Aerosol Flame Suppression Studies 395

Comparison of the results for the three different mists suggests the occurrence of several phenomena. In 

the experiments, droplets were entrained into dry air. The very rapid reduction in extinction strain rate 

with only a very small mass of water droplets at the flame, followed by a less dramatic fall off with

subsequent water droplet addition suggested a fair degree of humidification. This was particularly

noticeable for the 14 μm mist. The experimental data for 14 μm mist at greater than 0.005 mass fraction 

of added water and for the 30 μm mist data at greater than 0.015 mass fraction of added water droplets 

were in very good agreement with the model predictions. The 42 μm data were ≈ 12 % higher than the 

calculation predicted at 0.015 mass fraction of added water drops. This disagreement was consistent with 

quantitative prediction of the degree of air 

saturation expected for the experimental conditions follows below. 

an ensemble of monodisperse droplets at a specified initial mass fraction in

dry air was determined by direct time integration. For each 1 ms time step, the amount of vaporization 

ity, and the 

D-squared evaporation law. The amount of water vaporized impacted the relative humidity and

consequently the evaporation rate for the next time step. 

The predictions of the total amount of evaporation for a residence time of 750 ms are shown in

Figure 4-45. 

Figure 4-45 Calculated 

Humidification Levels of 

a 296 K Air Stream 

Containing 14 μm, 30 μm, 

or 42 μm droplets, 

Following a Residence 

Time of 750 ms Starting 

in Dry Air, as a Function 

of the Mass Fraction of 

Liquid Water Remaining 

after Droplet 

Evaporation. 113

the degree of humidification for this larger droplet size. A

Humidification Effects of Air Stream 

In order to quantify the degree of droplet evaporation for the different droplet sizes and the effect on 

flame extinction, evaporation times for the 14 μm, 30 μm, and 42 μm diameter droplets were calculated.

The calculation incorporates the droplet evaporation equations given in Chin and Lefebvre.113 Droplet

evaporation times in air at 296 K were determined for a range of relative humidity for each of the 

experimental droplet sizes. For each droplet size, the variation in droplet evaporation times between 0 %

and 100 % relative humidity was fit to a second order polynomial. 

The evaporation over time of

was calculated based on the instantaneous relative humidity, the droplet number dens
396 Aerosol Properties 

us, the assumption of saturated air in

the extinction calculation was valid and was consistent with the rapid reduction in extinction strain rate 

e 30 μm drops, two sets of experimental points were collected, 

clustered at liquid mass fractions of 0.015 and 0.028. For the 0.015 mass fraction tests, the relative 

ri

2. For this condition, the calculation

tion of velocity and size 

s ixed propane/air

counterflow flames. For both size mists, the peak in the droplet size distri he

flame zone was reached. The peak then shifted to smaller diameters due

number density with axial position were strongly correlated with variations in droplet axial velocity

fluxes of both 30 μm and 18 μm droplets decreased between the air tube

due to the effects of the diverging flow and evaporation. None of the 18 of the

30 μm droplets survived the flame, suggesting that for this range of droplet sizes in a counterflow flame at

moderate strain rate, most of the suppression potential of the mist is used. 

On a mass basis, both 14 μm and 30 μm diameter mists were found to be m 1

at suppressing non-premixed propane/air counterflow flames. The flam

42 μm diameter mist were considerably poorer than those of the 14 μm The lower

suppression efficiency of the 42 μm mist paralleled previous prediction odeling, and 

appeared to be caused by incomplete droplet vaporization during passage through the reaction zone. 

Comparison of the computational extinction strain rates with the reported experimental measurements

showed good agreement in consideration of the degree of mist evaporation in the experimental air stream

entering the flame. These results explained the observed droplet dependent suppression behavior in the

The residence time considered here was the time required to transport the mist after entrainment in

originally dry air up the 50 cm long inlet tube. A residence time of 750 ms is the shortest value studied

and corresponds to the air exit velocity necessary to produce flames with a local strain rate of ≈ 600 s-1. 

The shortest residence time provided a lower bound for the degree of humidification. Figure 4-45 plots 

the relative humidity (saturated air having a water mass fraction of 0.018 at the experimental temperature 

of 296 K) against the mass fraction of liquid water remaining after the specified residence time. The 

liquid water delivered to the flame as determined via PDPA measurements 2 mm above the air tube exit 

was reported by Zegers et al.110

, rather than the initial mass fraction of water introduced by the droplet

generator. Figure 4-45 shows that in the case of 14 μm drops, for any liquid water to remain after 

750 ms, the air stream relative humidity must be at least 95 %. Th

observed for this droplet size. For th

humidity after a 750 ms residence time was predicted to be approximately 85 %, while for the 0.028 mass 

fraction tests it was 95 %. Since the extinction strain rates for the flames inhibited by the 30 μm droplets 

were approximately 350 s-1, this implies a droplet residence time in the air stream prior to ar val at the

flame of approximately 1250 ms. The degree of saturation for a 0.015 mass fraction of residual liquid at

1250 ms was predicted to be 96 % (data not shown). For the 42 μm droplets the experimental 

measurements all had liquid mass fractions in the vicinity of 0.01

predicted a relative humidity of only 55 %; assumption of a residence time of 1000 ms, based on the 

measured extinction strain rate, gives a relative humidity of 66 %. The calculation assuming 100 % 

relative humidity was expected to overpredict the extinction strain reduction (lower extinction strain rate),

due to the additional heat capacity from the water vapor that was not present in the experiment. This

underestimate in the extinction strain rate was in quantitative agreement with the results for the 42 μm 

mist, as shown in Figure 4-44. 

Conclusions 

Using piezoelectric generation of aerosol drops, NRL investigated the evolu

distributions of initially monodisperse, 30 μm and 18 μm water mist in non-prem

bution did not change until t

to evaporation. Variations in

. The

exit and the stagnation plane, 

μ droplets and very few

ore effective than halon 130

e inhibition properties of the

 or 30 μm mists.

s of numerical m
Aerosol Flame Suppression Studies 397

range of 14 μm to 42 μm in suppressing counterflow propane air flames. The agreement between the 

predictions and experiments was remarkable, considering the high effectiveness of the small amount of 

water mass introduced to the flame with the air stream. The hybrid Eulerian-Lagrangian multi-phase 

flame suppression model and chemical kinetic propane combustion mechanism well predicted the 

measured local propane-air extinction strain rate. Water droplet suppression modeling of the flames and 

consideration of the amount of humidification of the air stream for the reported experimental mists 

quantitatively accounted for the observed response of the flames to the added liquid water mass as well as 

the effect of water droplets in this size range. Because of the higher strain rate at extinction, the optimum 

size water droplet was < 15 μm, somewhat smaller than for methane air flames. This smaller optimum 

size may be anticipated for other fuels considering the similarity of the extinction properties of propane to 

that of many other higher hydrocarbon “typical” fuels. The present findings indicated that, if the delivery 

issues inherent to a condensed phase fire suppressant were successfully addressed, aerosols of water 

composed of fairly small droplets have the potential to be comparable to that of CF3Br. 

4.5.2 Alkali Metal Bicarbonates, Chemically Active Fire Suppressants 

Fire suppression by chemically active particulates such as the alkali metal bicarbonates had been 

investigated for a number of years prior to the NGP.1,2,3,4 NaHCO3 powder (effective diameter < 10 μm) 

was known to be far more effective than any other gaseous chemical agent being considered to replace 

halon 1301, but the exact physical and chemical mechanism was unquantified. Explicit understanding of 

the suppression behavior was also lacking. Two NGP projects explored the suppression behavior of these 

powders in counterflow flames, providing quantitative information that contributed to the development of 

suitable chemical and physical models describing the behavior of these aerosols. 

Study I: Extinguishment of Flames with KHCO3 and NaHCO3 

 whose uninhibited strain rates 

-1 -1 -1

Experimental

NRL staff examined the fire suppression effectiveness of solid aerosols of sodium bicarbonate (NaHCO3) 

and potassium bicarbonate (KHCO3) in extinguishing non-premixed counterflow propane air flames using 

the counterflow burner described in Section 4.5.1 and shown schematically in Figure 4–36.114,115 For the 

powder aerosol suppression studies, air was supplied through the top tube and propane through the 

bottom tube as is shown in Figure 4-46. This figure also shows the location, relative to the stagnation 

plane, of the visible flame, and predicted profiles of the temperature and the flame radicals H, O and OH. 

Experimental strain rates were evaluated using laser Doppler velocimetry (LDV) following the procedure 

outlined by Zegers et al.109 Uninhibited propane/air flames had an extinction strain rate of ≈ 600 s-1. 

Powder suppression efficiency was evaluated by adding powder to flames

were 180 s , 305 s , and 475 s , referred to as low, medium, and high. The flows and strain rates are 

summarized in Table 4–6. 

The powder delivery system provided a steady flow of agent, with controlled variability from 

≈ 10 mg/min to 300 mg/min. The system, shown in Figure 4-47, consisted of a glass metering valve with 

an adjustable orifice and a tee above the valve where the powder could be introduced. The valve was 

made from a tapered brass rod inserted into a matching tapered glass orifice. The system was airtight and 

the orifice could be changed by screwing the tapered rod in or out. A hose clamp holding a piece of 

teflon was fastened to the glass metering valve at about the position of the valve. The teflon was used to 
398 Aerosol Properties 

contact a rotating wheel of variable frequency. A modulated helium-neon laser beam, monitored with a 

lock-in detector, was used to measure the powder flow. The laser beam was positioned perpendicular to 

the flow of powder just below the top tube exit. Scattered laser light created by the powder flow was 

collected by a photomultiplier tube and collection lens at 90°. The intensity of the scattered light was 

calibrated to the powder flow by collecting and massing powder with no flame. Because the particle 

scattering intensity is a function of the air velocity and the size distribution of the particles, a calibration 

was obtained for each powder size and at each flow tested. 

Figure 4-46. Schematic of the Non-premixed Counterflow Burner and Calculated 

Temperature, O, H, and OH Profiles for the Low Strain Flame. 114

Table 4–6. Flow and Strain Rate Conditions for Extinction Experiments. 114

Low Strain Flame Medium Strain Flame High Strain Flame 

Air Flow (L/min) 1.5 2.4 3.5 

Fuel Flow (L/min) 1.0 1.7 2.8 

Global Strain Ratea (s-1) 110 180 290 

Measured Strain Rate (s-1) 180 305 475 

a

 Calculated strain rate following Zegers et al.110
Aerosol Flame Suppression Studies 399

Air Inlet

Powder Fill Point

Rotating Knob to 

Adjust Orifice

#10-32 Threaded Brass 

Rod Modification Taper on Rod for

Orifice Variations

Glass Tubing

Vibration Transferred

to Glass at this point

via Hose Clamp 

Assembly Clamping Point 

Teflon Stopcock 

To Upper Burner Tube

Transition to 1/4" OD

Glass Tubing

A video system was used to determine the exact time that the flame was extinguished. Using a fiber optic 

positioned on the video screen, the intensity of a portion of the flame was monitored and recorded 

simultaneously with the powder scattering intensity using a LabView controlled A/D board. The addition

of powder to the flame dramatically increased the excited C2 emission from the flame. A suitable optical

filter was used to block this emission from the video monitor. Extinction of the flame was evidenced by a

dramatic change in light intensity from the video signal. The corresponding powder scattering signal was 

ignited and established midway between the two tubes, and the 

strain rate per Table 4–6. Once the system was acquiring d

evaluated to determine the powder flow rate at extinction. 

flows adjusted to produce the desired

ata, the powder was turned on, mixing into the 

ir stream from above the flame to simulate total flooding agent delivery systems. Particle delivery from

e top also minimized any size dependent particle entrainment/flow issues. The flow was gradually 

ramp the 

propane/air fla ached to the

upper tub nsistent

with othe air

flam s de suff it.

The flat horizontal burning region of the flame was centered along the axis and positioned at the location 

d enta of the flows. However, the flame was rounded at the edges an d a bowll e flame curved around and attached to the outer rim of the upp be. 

Upon extinction, the flat horizontal region of the flame disappeared, but the upper curved region of the 

burner tube. The flame would continue to burn in the annular 

Figure 4-47. Powder Delivery

System. 115

Prior to each test, the chamber and tubes were cleaned to remove any excess powder. The flame was

a

th

ed up until the flame extinguished. In this flow configuration (air on top, fuel on bottom)

mes tended to extinguish only along the centerline, leaving an annular flame att

e. However, the extinction strain rates determined from the centerline extinction were co

r studies and have been shown to be valid.

110 Total extinguishment occurred for the propane/

es only after an excessive amount of powder wa livered to the flame to iciently overwhelm

etermined by the mom

ike shape. The top of th

d ha

er burner tu

flame remained stabilized to the upper
400 Aerosol Properties 

onfiguration even after the flow of powder was stopped as the flame had attached and stabilized itself on 

the upper tube. The air and fuel flows were decreased to destabilize the annular flame and recover the

rs seemed to flow very freely with almost no

agglomeration noticed. Representative micrographs of the NaHCO3 powders are shown in Figure 4-48; 

micrographs for the KHCO3 powders, ground and sieved at 45 μm to 53 μm are shown in Figure 4-49. 

38 μm 45 μm to 53 μm 63 μm to 75 μm

c

original flame structure. The process was repeated until 8 to 10 extinction points were obtained. This 

procedure was repeated for the same powder at the remaining two strain rates. 

Tests were conducted with both sodium bicarbonate (NaHCO3) and potassium bicarbonate (KHCO3). 

Both powders were mechanically sieved into size ranges of <38 μm, 38 μm to 45 μm, 45 μm to 53 μm, 

53 μm to 63 μm, and 63 μm to 75 μm using a commercial sieving unit. Silica was added to the powder 

samples (2 % by mass) as a drying additive. Both powde

Figure 4-48. Typical Micrographs of NaHCO3 Powder Samples. 115

Figure 4-49. Typical Micrographs of KHCO3 Powder. 115

Results and Discussion 

Extinction mass concentration measurements as a function of the measured strain rate and particle size

bin are tabulated in Table 4–7 and plotted in Figure 4-50. The halon 1301 data were derived from Zegers 

et al.109 
Aerosol Flame Suppression Studies 401

Table 4–7. Propane-air Counterflow Non-premixed Flame Extinction Mass Concentrations 

for Powders and Halon 1301. 114

Extinction Mass Concentration for Specified Size Bin (g/m3) 

Chemical 

Measured 

Strain rate

(s-1) < 38 μm 

38 μm to 

45 μm 

45 μm to 

53 μm 

53 μm to 

63 μm 

63 μm to 

75 μm 

180 70 52 150 270 730 

NaHCO3 310 8 15 53 56 80 

480 4 1 7 7 4

180 12 18 70 100 370

KHCO 310 3 12 20 18 51 3

480 1 6 3 1 2

180 110

310 56 Halon 

1301

480 15 

Figure 4-50. Extinction Mass Concentration as a Function of Strain Rate for Various 

Particle Sizes R , Counterflow, 

s concentration varied inversely with the strain rate of the flame and directly

anges of NaHCO3 and KHCO3 Powders in a Propane/air

Non-premixed Flame. 114

In general, the extinction mas

with the particle size of the powder. KHCO3 was approximately 2.5 times more effective than NaHCO3 

on a mass basis in extinguishing the flame (3 times more effective on a molar basis). The data in

Figure 4-50 are replotted in Figure 4-51, showing the extinction mass concentration vs. average particle
402 Aerosol Properties 

diameter for each strain rate flame for NaHCO3 and KHCO3. The most dramatic increase in performance 

for the smaller particles was observed for the low strain rate flame. 

Figure 4-51. Data from Figure 4-50 Plotted to Show Extinction Mass Concentration as a 

Function of Particle Size Range for Each Flame Strain Rate and Added Powder. 114

A number of factors have been postulated to contribute to the inhibition of fires by powders although the 

exact mode of the alkali metal bicarbonates' effectiveness is still not completely understood. The powder 

adds a great deal of heat capacity to the flame, and solid particles are often effective infrared radiators, 

leading to radiation cooling of the flame zone. In addition, the alkali metal bicarbonates undergo 

endothermic decomposition at elevated temperatures, releasing carbon dioxide. The residual alkali 

hydroxide vaporizes and decomposes at a much higher temperature. The difference in effectiveness 

between sodium and potassium is also partially due to differences in decomposition temperatures and 

enthalpies, and possibly to differences in catalytic scavenging cycles between the two elements. The 

possibility of free radical scavenging processes involving either the solid surface of the particle 

(heterogeneous) or gaseous by-products of the decomposition (homogeneous) have been postulated. 

Catalytic reactions would involve the alkali metal ion (i.e., Na, K) and proceed as A + R + M → AR + M

followed by AR + R′ → A + RR′ where A is the alkali metal ion, R and R′ are H, O, or OH, and M is a 

ird body. These reactions are exothermic and add heat to the flame. However, under favorable 

ther it. 

Th 6

Plots of the extinction mass concentration as a function of average particle diameter and surface area can 

th

mochemistry, scavenging of the radicals inhibits the flame more than the additional heat promotes 

e possibility of catalytic recombination on the surface of the solid particle has also been suggested.11,11

be found in Figure 4-52a and Figure 4-52b, respectively, for the medium strain rate flames for both 

powders. The average particle diameter was calculated as the arithmetic mean of each particle size bin, 
Aerosol Flame Suppression Studies 403

Ewing et al. concluded that the fire extinguishing effectiveness of dry chemicals could be fully

nd 

3

le

h a

strain

ass

assuming an even distribution. This average particle diameter was then used in the calculation of the

average particle surface area. 

16

explained on the basis of particle size and distribution. They found that the effectiveness of the agent

increased gradually as the particle diameter decreased until reaching a critical diameter. At this diameter

there was a dramatic increase in the flame suppression efficiency and the effectiveness remained constant

for all particles below the limit size. The limit sizes reported were 16 μm and 22 μm for NaHCO3 a

KHCO3, respectively. NRL noticed a dramatic increase in the efficiency at about 45 μm for the NaHCO

sample at medium strain rate. A sizable increase was also noted for KHCO3, but at a much larger partic

size of about 65 μm. The higher value found for KHCO3 may be influenced by the large percentage of

very small particles in all size bins of this agent, as evidenced in representative photomicrographs of the 

various size bins. 

Figure 4-52a shows that the measured extinction concentration increases with the diameter, althoug

linear correlation between extinction mass concentration and diameter was not found at the medium

rate. There did appear to be a linear correlation at the medium strain rate case for the extinction m

concentration vs. average particle surface area as seen in Figure 4-52b. For this strain rate, the particle 

surface area available to interact with the flame more closely correlated with suppression effectiveness 

than particle diameter. 

Figure 4-52. Extinction Mass Concentration as a Function of (a) Average Particle

Diameter and Average Particle Surface Area for NaHCO3 and KHCO3 Powders in a 

Medium Strain Rate (310 s-1) Propane/air Counterflow Non-premixed Flame. 114
404 Aerosol Properties 

For the high and low strain rate cases, the correlation between surface area and effectiveness was less

clear. For the low strain rate flame, the largest particles fell directly through the flame and into the 

bottom tube, as they were too large to remain suspended near the stagnation plane by the opposing fuel 

ace area (e.g., by gas adsorption), and a more monodisperse size distribution 

would be useful. However, the extinction results for the 1-D flames of uniform strain rate are consistent 

mass, small particles are expected to be more effective at fire suppression, in

accord with our experimental observations. Any dependence on surface area does not, however, allow 

3. 

ttributed to the chemical effectiveness of alkali metal bicarbonates.117

For t tant.

As t at

differe i

e particles behave as spherical bodies, the smaller, less massive particles are expected to follow the flow

flow. A dramatic falloff in effectiveness was observed for the two largest particle bins, which likely 

resulted from the reduced residence time near the flame for these particle sizes. For the high strain rate

case, the amount of powder required to extinguish the flame was so small that the particle delivery was 

difficult to accurately measure and control with our seeding apparatus, leading to a larger uncertainty. It 

was thus difficult to quantify a relationship between particle size or nominal surface area and extinction 

effectiveness for these two strain rates, although smaller particles were more effective in both cases.

Effectiveness appeared to correlate more closely with surface area than with diameter. The determination

of surface area in this study was based on the assumption of a spherical particle of the average diameter

for the sieve fraction. A more accurate determination of the effect of surface area on suppression requires

direct measurement of surf

with “real world” observations that smaller particles are preferable presumably due to the increased

surface area. 

Suppression effectiveness should scale with particle surface area for most powder suppression modes.

This includes heat transfer, radiation, decomposition/vaporization, and surface catalysis. Therefore, for a

specified total powder

determination of the primary means by which these particles extinguish fires, due to the interrelated

operating suppression modes. 

Sodium bicarbonate and potassium bicarbonate have different decomposition temperatures. KHCO3 has a

decomposition temperature between 100 °C and 200 °C, whereas NaHCO3 decomposes at approximately

270 °C. Thus, KHCO3 would require less time to reach its decomposition phase than NaHCO3, under 

similar strain rate and particle size distribution. Also, according to data compiled by Ewing, Hughes, and 

Carhart15, the reaction heat sinks, those associated with the decomposition, vaporization and dissociation, 

are reported to be 59.7 kcal/mole and 128.6 kcal/mole for NaHCO3 and KHCO3, respectively, based on

vaporization of non-decrepitating solids. Non-decrepitating solids are defined as those which decompose 

or vaporize without breaking up. The sensible enthalpies of the decomposition/vaporization products of 

the two powders over a temperature range from ambient to the extinction temperature of a methane fire,

were reported by Ewing, Hughes, and Carhart15 to be 57.3 kcal/mole and 47.2 kcal/mole for NaHCO3 and

KHCO3, respectively. The total sensible enthalpies (phase change and heating of the

decomposition/evaporation products) are 117.0 kcal/mole for NaHCO3 and 175.8 kcal/mole for KHCO

On a mass basis, the values are 1.39 kcal/g for NaHCO3 and 1.76 for KHCO3. For comparable powder

sizes based on this thermal analysis, KHCO3 should be some 30 % more effective than NaHCO3. The 

comparable suppression effectiveness of small particles of sodium or potassium bicarbonate compared to 

halon 1301 and the much higher effectiveness of potassium vs. sodium shown in Figure 4-50 can be

a

he counterflow flame configuration, where the particles reside as they decompose is very impor

he particles exit the top tube of the burner, they position themselves in the counterflowing field

nt locations based on their size and mass and the magnitude of the velocity field. Assum ng that 

th
Aerosol Flame Suppression Studies 405

field as it changes direction to flow radially outward because of the matching opposed fuel flow. Li, 

Libby, and Williams

104 mathematically examined the behavior of droplets in an opposed flow field.

They observed that large droplets fall through the stagnation plane and are turned around by the opposing

flow, if the drag force is greater than the gravitational force. The smallest particles always followed the 

gas velocity streamlines (assuming minimal thermophorectic effects). Slightly larger particles penetrated

the stagnation plane and could be “pushed back” by the opposing flow and could exhibit a decaying

oscillatory behavior near the stagnation plane. The amplitude for this oscillation was a function of the

particle size. Lentati and Chelliah80 calculated different trajectories for water droplets as a function of 

size in a counterflow burner. Their calculations indicated that the oscillatory amplitude is essentially zero 

for a 30 μm (non-evaporating) water drop. For smaller drops, the resulting oscillatory motion was

essentially confined to a plane. It should be noted that these calculations are valid only for the particular 

conditions (e.g., strain rate, spherical drops, fuel density) of the models used. However, comparison to 

particles is qualitatively similar. 

The proximity of the oscillatory “plane” to the stagnation plane depends on the particle shape and size 

and the strain rate of the flow but not on the original direction of the particle trajectory. Figure 4-53 

shows the axial positions of the oscillatory planes relevant to bicarbonate spheres, 10 μm to 70 μm in

, H, and OH radical profiles for the

diameter, in a potential non-reacting counterflow field model assuming the experimental flows for the 

medium strain rate flame studied. The drag forces on the particles were assumed to obey Stokes law. The

0 cm and 1 cm positions corresponded to the propane and air tube exits, respectively. Ratios of 

experimental flow to tube cross section were used as model exit velocities, yielding a stagnation plane at 

0.39 cm. 

Figure 4-53. Calculated 

Particle Asymptotic Location 

as a Function of Particle 

Diameter for the Flow

Conditions of the Low Strain 

Propane/air Flame. 114

Also shown in Figure 4-53 are calculated axial temperature and O

propane/air flame at the medium strain rate. For particles introduced into the upper flow or the lower 

flow, the contribution of gravity dictates that the particle will asymptotically oscillate in a plane below the

stagnation plane. Smaller particles were predicted to lie closer to the stagnation plane. The location of 

the oscillatory plane becomes significant if the opposing flow supports a flame. Stoichiometry dictates 

that the propane/air non-premixed flame be located on the air side, as shown in Figure 4-53. With air
406 Aerosol Properties 

(and powder particles) on the top and propane on the bottom, the flame is located above the stagnation 

plane. All particles will then pass through the flame and those particles within a certain range of 

diameters (~ 20 μm to 60 μm for the low strain rate flows) can oscillate below the stagnation plane and 

hence below the center of the flame but still in an elevated temperature region. As the particles 

decompose, they become smaller, and the opposing flow will move them closer to the stagnation plane, 

but never in the flame center if delivered with the air from the top. 

With propane on top and air on the bottom, the flame is located below the stagnation plane on the air side. 

In this configuration, some of the particles oscillate at the flame maximum temperature, and hence also at 

the peak of the key flame radicals. This oscillatory location should lead to greater decomposition because 

the particle would experience a greater residence time in the higher temperature region and ultimately 

generate a higher inhibitor concentration closer to the maxima of the key flame radicals. Following this 

analysis, one might expect that there should be an optimum delivery particle size which would allow the 

particle to locate at the flame maximum. Since the distance of the flame center to the stagnation plane is 

Flame extinguishment by dry chemical agents is due to a combinat thermal and chemical 

mechanisms. The relative contribution of each mechanism is difficult to determine since effectiveness 

depends on the nature of the particle and the particle size and the flame conditions. Both suppression 

mechanisms rely on the degradation of the solid particles to achieve the maximum effect on the flame. 

The degree of degradation is a function of the particle size which with the gas velocity determines 

eventual location of the particles in and near the flame. Smaller particles decompose faster and this 

degradation increases suppression effectiveness in two ways: heat abstraction due to highly endothermic 

processes, including decomposition and vaporization; and product t can 

promote scavenging of the important flame propagation radicals. Sin

through the flame, the location of the thermal extraction in the flow is 

the chemical players (if any) are released. 

Conclusions 

In propane/air counterflow non-premixed flames, KHCO3 was shown to be more effective at 

extinguishment on a mass basis than NaHCO3 for all particle sizes tested. The effectiveness of the 

powder agent varied inversely with the particle size for the size range tested; smaller particles were more 

effective in suppressing the flames than the larger ones. Model predictions addressing the behavior of 

particles in counterflowing fields were consistent with the observed behavior for the suppression 

effectiveness when the fuel is on the bottom. An explanation of the high effectiveness observed for the 

a function of the strain rate, one might predict that the optimum particle size might also vary with strain 

rate. Experimentally, delivering powders via the lower tube is more complicated than delivering powders 

from the top. Top delivery is essentially “dropping” the particles on the flame; there is no restriction on 

how big a particle can be and entrainment is straightforward. Bottom, upward delivery, for the gas 

velocities and particles in this study, cannot deliver the largest particles to the flame since these particles 

cannot be supported and carried “up” by the entrainment flow velocity. 

ion of both

ion of gas-phase atoms tha

ce all of the flow is forced to go 

probably less important than where 

alkali metal bicarbonate particles required a significant contribution from a chemical suppression 

mechanism. 
Aerosol Flame Suppression Studies 407

Study II: Modeling and Flame Extinguishment Measurements of NaHCO3

Contrary to the extinction behavior observered for NaHCO3 reported in the prior section, Trees and 

Seshadri118 found a non-monotonic variation of NaHCO3 particle mass fraction with particle size required 

to extinguish a non-premixed counterflow flame for the size range ≈ 0 μm to 30 μm. Chelliah and coworkers conducted NGP studies to address the uncertainties regarding the effect of NaHCO3 particle size 

on flame extinction and developed a rigorous method of quantifying various physical, thermal, and 

O3 particles, 

Figure 4-54. Schematic of the Counterflow Flame 

Apparatus with the Particle Seeder. 119

chemical contributions.119 

Flame Suppression with NaHCO3

Figure 4-54 shows the apparatus for generating non-premixed, methane air counterflow flames. The 

NaHCO3 particles were separated into five size ranges: <10 μm, 10 μm to 20 μm, 20 μm to 30 μm, 

30 μm to 40 μm, and 40 μm to 60 μm using a vortex separation technique. Particles of a given size were 

introduced with the air stream at a relatively steady rate. The typical particle mass fraction was about 

0.01, corresponding to a particle flow of about 0.1 g/min. For a selected NaHCO3 loading, the flame 

extinction effectiveness of NaHCO3 particles was characterized by the extinction flow strain rate, 

measured by two methods: a global formula based on the nozzle exit velocities and nozzle separation 

distance81 and local velocity field measurements using laser Doppler velocimetry (LDV). 

Because of the need to establish the accuracy of previously reported data with NaHC

considerable attention was devoted to the development of a particle seeder that could deliver steady feed 

rates of particles of specified sizes. Two approaches were used. 

Water Fluidizer air

Nitrogen 

Observation 

windows

Co-flow

Fluidizer 

Ejection

Methane 

Nitrogen 

Laser 

diode

and the 

detector
408 Aerosol Properties 

For the small air flows (typically 0.1 L/s), the fluidized bed approach shown in Figure 4-54 gave 

satisfactory seeder performance, but only for particles less than 30 μm. The fluidized bed seeder 

e mass fraction of particles attained a steady state.

The particle feed rate was continuously monitored using a Mie scattering detection system located at the

igure 4-54. This detection system consists of a laser diode operating 

and after each

arrangement consisted of a multi-start Teflon s

was driven by speed-locked variable speed motor.

This seeder was vibrated pneumatically to assist the particle flow, but the particle feed rate was primarily

controlled by the motor speed. Part of the total air flow was diverted through a screw-driven particle 

seeder and was then mixed with the remaining air prior to exiting through the air nozzle. As before, air 

flow through the seeder was held constant while the total air flow through the nozzle was varied to attain

different flame strain rates. For particle sizes μm, the Mie scattering signal varied less than

± 10 %. Particles smaller than 20 μm had a tendency to clump, and consequently significant variation of

the scattering signal was observed. Hamins et al.20 had overcome such clumping by adding about 1 % by

mass of 10 nm silica particles to the NaHCO3 particles. Due to the uncertainty of the effect of adding the

silica particles and the associated safety issues, no silica was added. 

Without particles, the measured global flame e

local LDV strain rate was about 390 s-1. Thes reviously by 

Chelliah et al.

103

 As NaHCO3 particles were added to the air stream, the flame extinction strain rate was 

expected to decrease because of the physical, thermal, and chemical effects of the particles. 

The measured variation of flame extinction strain rate vs. the NaHCO3 particle mass fraction in the air

stream is shown in Figure 4-55 for the five particle size groups. All five groups demonstrated a clear 

decrease in the extinction strain rate with increasing particle loading. For a given particle mass fraction,

the smallest particles (0 μm to 10 μm and μm to 20 μm, which gave indistinguishable results) 

produced the lowest extinction strain rate, indicating their highest effectiveness as a flame suppressant. 

consisted of a 19 mm diameter glass tube with two porous plugs connected at either end. The air inflow

to the seeder, controlled through a mass flow controller, was split into two streams and connected to 

either end of the glass tube. The flow passing through the bottom, which controlled the fluidization level,

was monitored through another flow controller. The fluidized particles and the total air entering the 

fluidization tube were ejected through a small tube located toward the middle of the tube. By maintaining 

the total air flow through the fluidizer tube constant, th

exit of the seeder air, as shown in F

at 671 nm, a Thor Labs DET1-SI High Speed Silicon Detector with collection by a 3 cm focal length lens, 

background radiation filtering using an Edmund Scientific 671 nm interference filter, and a HewlettPackard Model 7132 strip chart recorder. Because the seeder air flow was held fixed, the scattering signal 

was independent of the total air flow exiting through the air nozzle and was primarily dependent on the 

particle delivery rate of the seeder. The scattered signal was observed to decay slightly over 20 min to

30 min periods, but could be controlled by shifting part of the bottom fluidizing air to the top. Different 

particle flows were obtained by controlling the fraction of seeder air flowing through the top and the 

bottom. The actual mass fraction of the particles in the air stream was determined by calibrating the

recorded scattering signal by a separate gravimetric analysis, usually performed before

flame extinction experiment. 

For particles greater than 20 μm, a positive feed auger system (not shown) was employed. This seeder

crew (7 mm outer diameter and 5 mm inner diameter) in a 

high-precision stainless steel tube, where the screw

above 20 

xtinction strain rate was about 470 s

-1, while the measured

e values are consistent with those reported p

10
Aerosol Flame Suppression Studies 409

Trees and Seshadri 118 had reported a similar variation of the mass of particle required with varying 

extinction strain rate, for each size group. However, they found that while 0 μm to 10 μm particles were 

mpose into oxides of Na in two 

age, Na2CO3

is known to decompose at 1170 K, leading to formation of Na2O and CO2. In the presence of water, 

the most effective, 10 μm to 20 μm particles were the least effective, with 20 μm to 30 μm particles 

showing an intermediate effectiveness. There were two differences between the studies that could have 

affected the results: (a) The fuels were different. Seshadri used a liquid heptane fuel pool, while the 

present experiments used gaseous methane. (b) Seshadri added about 1 % by mass of 10 nm silica 

particles to minimize the clumping of particles and help fluidize the smaller particles. 

Figure 4-55. % NaHCO3 by Mass in air as a Function of Extinction Strain Rate for Different 

Size Groups of Particles. 119

When sodium bicarbonate particles are heated, they are known to deco

stages. In the first stage, NaHCO3 decomposes to solid sodium carbonate (Na2CO3), H2O, and CO2

around 543 K. (Note that the global decomposition temperature can vary depending on the particle 

heating rate, e.g., smaller particles are expected to heat rapidly when exposed to the flame environment 

because of their higher surface area/volume ratios than the larger particles.) In the second st

however, Na2CO3 and Na2O are known to react to form sodium hydroxide (NaOH).120 The NaOH formed 

promotes catalytic recombination of the radicals species needed for flame propagation, causing the flame 

to extinguish readily.121 

0

0.5

100 150 200 250 300 350 400

Extinction Strain (1/s)

1

1.5

2

2.5

3

NaHCO3 Mass Fraction (%)

0-10 um

10-20 um

20-30 um

30-40 um

40-60 um
410 Aerosol Properties 

Numerical Simulation of NaHCO3 Particle Effects 

The numerical model described earlier in Section 4.5.1 was extended to include the effects of sodium 

bicarbonate particle transport, heating, decomposition and subsequent interaction of decomposed gaseous 

species with the homogeneous flame chemistry. The detailed gas-phase kinetic model for methane-air

was augmented with a 20-step elementary mechanism for Na from Jensen and Jones.121

 The model 

to be responsible for reduction of the radical pool 

developed for sodium bicarbonate particles involved considerable additional uncertainties. These 

included the non-spherical nature of the particles, the transient heating effect of the particles, the

NaOH + H → H2O + Na,

Na + OH+M → NaOH + M

heterogeneous decomposition mechanism of sodium bicarbonate, and the gas-phase chemical inhibition 

mechanism by intermediate species formed from the decomposition. 

Previous studies had indicated that sodium bicarbonate decomposes at a low temperature of about 543 K

to form sodium carbonate through the reaction:10

 2 (NaHCO3)s → (Na2CO) s + CO2 + H2O. 

For particle temperatures above 1124 K, the sodium carbonate formed can decompose further to form 

other reaction intermediates of Na. Based on these observations, Friedman and Levy120

 had postulated 

that NaOH is the primary inhibiting compound responsible for flame extinction and proposed the 

following reaction pathways leading to formation of NaOH 

 (Na2CO3)s → (Na2O) liq + CO2, 

 (Na2O)liq + H2O → 2NaOH, 

 (Na2O) liq → Na2O, 

followed by the homogeneous reaction: 

 Na2O + H2O → 2NaOH. 

Once the alkali metal hydroxide was formed, the following catalytic radical scavenging cycle was known 

,

NaOH + OH → H2O + NaO. 

The relative importance of these reactions was not well established and the literature indicates some 

conflicting results.117,120,121,

In the initial NGP simulations, a fast global decomposition reaction of the form, 

 2(NaHCO ) → 2NaOH + 2CO

122

3 s 2
Aerosol Flame Suppression Studies 411

was employed. The particle temperature at which the decomposition occurred (defined here as Tdecomp) 

was varied to obtain the sensitivity of the particle size. Although such a global model can be fine tuned to

n, a plot of equivalent NaOH mass 

fraction for different extinction strain rates is also shown. The predicted curve with NaOH indicates the

rve, but the predicted trends were not expected to

change. The results clearly indicate that particles above 20 μm have a small effect on flame extinction,

, the extinction strain rate decreased significantly,

Figure 4-56. Numerical 

Prediction of NaHCO3 Mass 

Fraction as a Function of 

Extinction Strain Rate, for 

Selected Particle Sizes. 119

y vary , the location at which these particles were rapidly decomposing in 

e terms were selectively turned “off” in the model. Unlike previously reported 

ffect in the presence of

3 were found to be 

ing the particle heating and subsequent heterogeneous particle 

decomposition need to be developed in order to provide more accurate flame extinction predictions. 

a narrow range of conditions (e.g., narrow particle size range), it should be cautioned that extension of 

such models to a wider range of conditions may lead to physically unrealistic results. For the

homogeneous chemistry associated with NaOH, a detailed reaction model involving 6 species in

20 elementary reactions proposed by Jensen and Jones121 was employed, in conjunction with the detailed 

chemistry model for methane oxidation.80

Figure 4-56 shows the predicted variation of NaHCO3 mass fraction as a function of flame extinction

strain rate, for selected NaHCO3 particle size groups. For compariso

absolute maximum inhibition possible, based on the homogeneous model employed. Different NaOH

reaction mechanisms were expected to modify this cu

but as the particle size decreased below 20 μm

indicating a remarkable increase in the effectiveness. This effect was attributed to the rapid heating of the

particle and fast decomposition processes.

B ing the Tdecomp described earlier

the flow field was varied. As a consequence, the amount of Na released to the homogeneous phase is 

affected, leading to a varying degree of chemical inhibition. To quantify the contributing thermal and

chemical effects, the sourc

fine-water droplet simulations, these results indicated a significant chemical e

NaHCO particles. For the dilute particle loading considered here, the physical effects 

negligible. Improved methods of treat
412 Aerosol Properties 

Conclusions 

All the NaHCO3 particle sizes considered (< 10 μm, 10μm to 20μm, 20μm to 30μm, and 40μm to 60 μm) 

indicated a decrease in flame extinction strain rate with increasing particle mass fraction, consistent with 

the NRL results. The effectiveness of flame suppression (quantified here based on the NaHCO3 mass 

fraction needed) varied monotonically with particle size, with the smallest size group being the most 

effective. Two independent approaches to particle seeding yielded identical results for the overlapping 

particle size range of 20 μm to 30 μm, providing considerable confidence in the measured results. 

Co-flow, Non-premixed Flames – Cup Burners 

The effectiveness of aerosol agents in a cup burner apparatus was not explicitly investigated under the 

NGP. What was investigated and is reported in Chapter 3, was the role of particle formation in the flame 

after agent introduction which limits the effectiveness of chemical agents. This is a fundamental 

thermodynamic limitation regardless of whether the agent enters the flame as a gaseous species or a 

condensed liquid or solid. The cup burner flame still remains a challenging yet relevant configuration for 

understanding aerosol effectiveness in general. Thus, work continu ding the 

NRL84 to understand the behavior of aerosols agents in a cup burne me 

configurations. 

4.5.3 Aqueous Metal Salt Solutions, Combined Chemical 

Non-Premixed Flames 

Introduction 

Chelliah and coworkers, using both experimental and modeling approaches, examined the interactions of 

small droplets of water and water/NaOH solutions with steady, laminar counterflow methane/air nonpremixed flames.98 To compare to the experimental studies, the previously developed hybrid EulerianNumerical predictions taking into account the detailed homogeneous chemistry and global particle 

decomposition model indicated a trend in particle effectiveness very similar to that obtained from the 

experiments. By varying the temperature at which the particles decompose to gaseous products, the 

predictions showed that the smallest particles are more effective because of their rapid heating, leading to 

complete decomposition of the particles. By manipulating various source terms in the model, the superior 

effectiveness of NaHCO3 particles was attributed to the homogeneous catalytic radical scavenging by 

NaOH formed, while thermal effects associated with particle heating played a minor role. 

es in a number of labs inclu

r apparatus and other co-flow fla

and Physical Suppression 

Lagrangian numerical model for monodisperse droplet distributions was generalized to account for more 

realistic, polydisperse size distributions. 
Aerosol Flame Suppression Studies 413

Experimental Method 

e post 

combustion gases eliminated secondary flames. The nozzle tubes entered the chamber through vacuum 

An oil-free shop compressor followed by a series of desiccant beds provided the air; the fuel gas was 

methane (BOC grade 4.0, 99.99% purity). Mass flow meters (Teledyne Hastings–Raydist, factory 

calibrated with a reported accuracy of ± 1% of full-scale reading) indicated the volumetric flows of the 

methane and air. For experiments involving water vapor or water droplets, the metered dry air was 

saturated with water vapor prior to the nozzle exit by bubbling the air through a water bath. A bath 

The counterflow burner, shown in Figure 4-57, was similar to that described in Study II in Section 1.5.2. 

A steady, planar, non-premixed flame was established in the mixing layer of counter flowing methane and 

air streams. The fuel and air nozzles (Pyrex glass) each had an area contraction ratio of 6.5 and an exit 

diameter of 1.5 cm, producing nearly plug flow velocity profiles at their exits. Co-flowing nitrogen 

streams on both fuel and air sides helped to maintain a very stable planar flame disk. The nozzles were 

enclosed in a cylindrical burner chamber, in which water cooling coils and air dilution of th

fittings which permitted easy adjustment of the nozzle separation distance, which is typically set to 12 

mm. An air-driven mass flow ejector evacuated the product gases, while a differential pressure gauge 

monitored the chamber pressure, which was typically 120 Pa (12 mm water) below atmospheric pressure. 

Figure 4-57. Schematic of the Counterflow Burner 

Configuration with the Water Droplet Atomizer.98
414 Aerosol Properties 

heated to 5 oC above ambient produced an air stream at the nozzle exit at ambient temperature and at least 

97 % relative humidity (verified with a hygrometer, Testo 605-H1).

clogging and erosion. The second atomizer w

the droplets generated had a much wider s varied by

frequency. All the experimental flame extinction 

ator. 

stem consisted of the ultrasonic nozzle and a broadband ultrasonic

nstech Model 2000) with a plastic 10 cc syringe fed water to the 

as fixed, and the air and fuel flows were increased

e was found to be very stable until the extinction point. The water 

red gravimetrically by carefully collecting all of the saturated air

a collection surface. These tests showed that while

atomizer reached the nozzle exit, the operation of the 

nt over the entire range of water and air flows of the tests. 

u

nozzle was characterized by the manufacturer

The droplet atomizer was located at the base of the bottom (air) tube, and the droplets were entrained in 

the airflow. The glass tubes allow easy detection of any water condensation on the tube walls. The 

droplet size distribution is measured using a PDPA (Aerometrics, two-component; model, DSA). 

Two types of droplet generators were tested: (1) a piezoelectrically excited fluid jet atomization system

(Fluid Jet Associates)123 and (2) an ultrasonic fluid surface breakup system (Sono-Tek, Model 8700-120). 

The first atomizer was capable of generating truly monodisperse droplets (variable by changing the orifice 

diameter), but the small, 10 μm holes needed for generating 20 μm droplets proved susceptible to

as found to be relatively simple to use and did not clog, but

ize distribution. The median droplet size was

selecting a different length nozzle tip and resonance 

data were obtained using the ultrasonic droplet gener

The ultrasonic droplet generation sy

frequency generator. A syringe pump (I

atomizer. In the experiments, the water mass flow w

until the flame extinguished. The flam

mass flow at the nozzle exit was measu

stream (with droplets) and impinging the droplets on 

only about 70 % of the mass of water injected to the 

droplet injection system was very consiste

The droplet number distribution ( f ) as a f nction of size (d) in the vicinity of the ultrasonic atomizer

and was believed to follow a log-normal distribution:

( ) ⎥

⎦

⎤ ⎢

⎣

⎡ − = − 2 2

ln ln

exp 2

1

σ

μ

π σ

d

d f d (4-12)

with a median diameter (l) of about 20 μm and a Sauter mean diameter (SMD) of about 30 μm (for σ = 

0.6). This normalized distribution, separated into 12 discrete sections for modeling purposes, is shown in

Figure 4-58a. Nonetheless, based on the esti ated settling velocities of different droplet sizes and the 

flow velocities in the air tube, it was found that not all the droplets produced by the atomizer would be 

convected out of the air nozzle. For example, for an airflow corresponding to a flow strain rate of about

285 s-1, the estimated maximum droplet size that could be carried in the air stream was about 75 μm,

whereas for a strain rate of 160 s-1, the corresponding maximum droplet size was about 55 μm. Thus, the 

air flow velocity up the air-nozzle would modify the generated droplet size distribution. PDPA 

measurements verified the droplet sizes at the air nozzle exit. Figure 4-58b shows the measured droplet

size distribution for a high strain rate of 285 s , and Figure 4-58c shows a similar plot for a lower strain 

rate of 160 s-1. The shift in maximum size of droplets was consistent with estimates, and the general 

hape of the distribution remained log-normal. 

m

-1

s
Aerosol Flame Suppression Studies 415

-1

-1 98

Figure 4-58. (a) Normalized Droplet 

Diameter Distribution of the 

Ultrasonic Atomizer as Reported by 

the Manufacturer; (b) and (c) 

Normalized Water Droplet Diameter 

Distributions Measured Using a 

PDPA at the Exit of the Air Nozzle for 

an Air Flow Corresponding to 

Counterflow Strain Rates of 285 s

and 160 s , respectively.
416 Aerosol Properties 

Flame extinction experiments were conducted by in

such that the momentum of the two streams was balanced, that is, 

creasing the air and methane nozzle exit velocities 

( ) ( ) 4

2 2

air CH ρv = ρv (4-13)

where ρ is the density and v is the axial velocity. Knowing the nozzle separation distance L, the flow 

strain rate was defined by the global formula81

L

v

a

air = 4 (4-14)

For non-premixed methane/air flames, the measured global flame extinction strain rate of 470 s-1 was 

obtained, while the measured local flow velocity using a laser Doppler velocimetry system yielded a local

flow strain rate of about 390 s-1. Both of these numbers were highly reproducible and consistent with

previous experiments and modeling efforts.103

 The experimental flame extinction results are all based on 

the global strain rate formula unless otherwise noted. 

Results 

The presence of condensed-phase water implied that at equ th

water vapor, and this was verified using a hygrometer. At a e

of 300 K, the saturated water vapor mole fraction in air is

saturated water vapor can have a significant effect on the flam d

modeling efforts had indicated that the predicted local flam 20 s

-1

for a methane/air flame to 365 s-1 for methane and air s

extinction strain of 12 %).

80 Present experiments with satur

global extinction strain rate of 405 s-1 (i.e., a reduction of 13 d

dry air), in excellent agreement with the predictions. 

Lazzarini et al.

98 also examined the effect of water droplets of 20 μm median diameter on the extinction 

condition for the flames. With increasing droplet number density (or mass fraction of water droplets in the

air stream), it was expected that flame extinction would occur more readily, resulting in a lower extinction 

strain rate. Figure 4-59 plots the mass fraction of water in the condensed phase (Y0) as a function of the 

flame extinction strain rate. 

Note that the zero water droplet mass fraction corresponds to the case where the air stream is saturated

with water vapor (i.e., Yvap = 0.0224). The predicted variations of water droplet mass fraction as a 

function of global extinction strain rates assumed 20 μm monodisperse droplets. Irrespective of the

assumption of monodisperse droplet size distribution in the simulations, the predicted trend is seen to be 

in reasonable agreement with experiments. 

The experiments and predictions agreed well for pure water vapor. The agreement grew poorer with 

increasing droplet mass loading. In predictions, relaxation of the monodisperse size distribution 

approximation based on the measured distributions shown in Figure 4-58 was not expected to rectify this 

difference because 20 μm monodisperse droplets have been predicted to be the most effective. As 

discussed later in the numerical section, any broadening of the size distribution about 20 μm led to

predicted higher mass fraction of water in the condensed phase for the same flame extinction strain rate. 

hus, differences seen in Figure 4-59 between experiments and modeling were likely due to other causes. 

ilibrium, the air stream was saturated wi

tmospheric pressure and a room temperatur

0.0351 (mass fraction Yvap = 0.0224). The 

e extinction condition. Previously, detaile

e extinction strain rate reduced from 4

aturated with water vapor (a reduction in

ated water vapor in the air stream yielded a 

% from 470 s-1 measured for methane an

T
Aerosol Flame Suppression Studies 417

In experiments, calibration of the water droplet mass flow through gravimetric analysis introduced errors; 

however, considerable care was taken to address uncertainties associated with the approach adopted. 

Figure 4-59. Variation of Water Droplet 

ding a chemically active fire-suppressing compound 

in water. NaOH was selected, as it was known to be the primary compound in the catalytic radical 

lutions of NaOH and investigate their effect on suppressing counterflow 

xpected to affect the evaporation of NaOH. The measured 

Mass Fraction in the Condensed Phase as 

a Function of Flame Extinction Strain 

Rate. 98

Experiments and simulations also examined the effect of the addition of NaOH to the water droplets. The 

primary mechanism of flame extinction by pure water droplets is through the thermal cooling of the flame 

front, leading to slower chemical reaction rates. For example, in a counterflow field of methane and air, 

the 20 μm droplets were shown to be most effective because most of the droplet mass was predicted to 

evaporate near the oxygen consumption or radical species production region.105,106 This thermal effect of 

water droplets can be considerably enhanced by inclu

recombination path of sodium bicarbonate fire suppression.121 Since the solubility of NaOH in cold water 

is about 30 % of the total mass,124 a significant amount of NaOH can be released at the flame front, 

provided that the droplets consisting of water/NaOH are completely vaporized. As noted earlier in this 

chapter, 20 μm droplets were fully evaporated in this flame system. The same ultrasonic was used to 

deliver various aqueous so

methane/air flames. 

Figure 4-60 plots the experimentally measured water/NaOH mass fractions as a function of the flame 

extinction strain rate. As before, the air flowing into the nozzle tube was saturated with pure water vapor. 

Because the NaOH vapor pressure is very small, less than 130 kPa (1 mm Hg) at room temperature124, air 

saturated with pure water vapor was not e

results indicated that with increasing NaOH mass fraction in water, the amount of water/NaOH mass 

fraction needed for flame extinction decreased significantly - almost a factor of 5 for 17.5 % NaOH by 

mass in water at the lowest strain rate considered. Combining thermal and chemical suppression effects 

leads to a significant increase in fire suppression ability, as shown in Figure 4-60. 

For a selected extinction strain rate, Figure 4-60 also shows an interesting nonlinear decrease in total 

water/NaOH mass fraction with increasing NaOH fraction in the solution. Such a phenomenon may be 

related to saturation of NaOH in the vapor phase and must be addressed through future modeling efforts. 
418 Aerosol Properties 

Figure 4-60. Droplet Mass Fraction in 

the Condensed Phase as a Function of 

Flame Extinction Strain Rate for 

Different Mass Loadings of NaOH in 

Water. 98

he hybrid Eulerian-Lagrangian formulation for gas and liquid phases described in Section 4.5.3 was 

used here.80 This model includes a detailed reaction model for methane oxidation and transport of species

Numerical Predictions 

T

and energy across the mixing layer. In this approach, knowing the gas-phase solution, the Lagrangian

equations for mass, momentum, energy, and particle flux fraction (normalized by that at the air nozzle

exit) were integrated in time to determine the droplet location and source terms contributing to the gasphase conservation equations. With these new source terms, the Eulerian equations describing the gas

phase were then integrated using a standard approach.125 Finally, the two sets of coupled equations were 

iterated until a predetermined convergence criterion was reached. 

As described in the experimental section, the flame extinction was realized by increasing the flow of

methane and air (saturated with water vapor), while the mass flow of condensed water droplets (mH2O,c) 

was held constant by the syringe pump. During this process, the mass fraction of water droplets 

air H O c

mH O,c 2 = m m

Y

,

0

2 + (4-15)

or a constant mH2O,c = 

2.65 mg/s/cm generated the innermost closed oval in Figure 4-61. Starting at some strain rate away

from extinction (e.g. point P), extinction conditions were predicted depending on whether strain rate was 

in air changed because of the varying air mass flow (mair). In previous numerical predictions, it was 

assumed that the condensed phase water mass fraction (Y0) was held constant at all strain rates. The 

predicted flame temperature variation as a function of flow strain rate for constant mH2O,c and for constant 

Y0 are shown in Figure 4-61 for 20 μm monodisperse droplets. In the absence of radiative losses, usually

only one extinction condition is realized for constant Y0. Two flame extinction points are predicted if 

strain rate is changed in a proportional manner to keep the momentum balanced as depicted in

Figure 4-61. Changing the strain rate by varying only the air and fuel flows f

2
Aerosol Flame Suppression Studies 419

increased (point E1) or strain rate was decreased (point E2)

of the two-phase system considered, and its occurrence

practice, this second extinction point (E2) can be realized

droplet transport is held constant. 

Figure 4-61. Variation of the 

Maximum Flame Temperature as a 

Function of Local Flow Strain 

Rate for Constant Droplet Mass 

he model developed previously to simulate the interaction of fine water droplets with the counter flow 

e assumed monodisperse droplets.80 This assumption made the computations less demanding, but 

ses of the results on droplet size effects became considerably simpler. 

ntal difficulties in realizing truly monodisperse water droplets, the above model

. This phenomenon was purely a consequence 

was observed for water/NaOH solutions. In 

by decreasing the oxidizer transport while the 

Fractions (lines) and Constant 

Water Mass Flux Rates (Lines 

with Symbols). 98

T

flam

more importantly, the analy

Because of the experime

was generalized here to include the polydisperse size effects. It was shown previously that for water mass

loadings similar to those considered here, the ratio of droplet-to-droplet separation distance to droplet size

was over 20. Under such conditions, the equation for droplet mass flux (I) (i.e., the spray equation), 

written in the form for the present quasi-one-dimensional counterflow field80

+ I = Γ I Ud dt

d 2 (4-16

127

)

was simplified by setting the droplet collision source term Γ = 0. Thus, by treating the droplet size 

distribution as composed of several discrete size groups126, , the simulation simplified to solving several

ith the three different discrete polydisperse size distributions shown in Figure 4-58. A comparison of

these flame extinction results with the ideal 20 μm droplet size for this flow configuration is sh

Table 4–8. The predicted extinction strain rates indicate that any broadening of the size distribution of 

roplets from the ideal 20 μm size led to a higher flame extinction strain and a further deviation from the 

Lagrangian equations for these discrete droplet size groups, with initial conditions corresponding to each 

discrete size group specified at the air nozzle exit. 

For 0.01 mass fraction of water in the condensed phase, flame extinction calculations were performed 

w

own in

d
420 Aerosol Properties 

experiments. At higher water mass loadings and decreasing flow strain rates, the thermal radiation effects

are known to increase and may also have an influence in the present predictions. However, the strain rates 

indicated in Table 4–8 are rather moderate, and it is unlikely that the significant differences between 

experiments and modeling can be solely attributed to radiation heat loss effects. 

Table 4–8. Experimental and Predicted Global Flame Extinction Strain Rates (s-1) for 

Different Droplet Size Distributions and Water Droplet Mass Fraction of 0.01. 98

Experiment 

20 μm 

Monodisperse

Droplets 

Distribution

from Figure 

4-58a 

Distribution

from Figure 

4-58b 

Distribution

from Figure 

4-58c 

Global strain rate (s-1) 180 286 297 291 292

Conclusions and Summary 

The main purpose of the work reported in this section was to 

predictions of the effectiveness of fine water droplet

On a mass basis, the predicted ability of fine water mi

effectiveness than halon 1301 was verified. Although the ori

monodisperse droplets as assumed in previous theoretical inv

challenging because of difficulties associated with clo

ultrasonic water atomizer generating log–normal distribution of 

provide experimental data to validate 

s in extinguishing counterflow non-premixed flames. 

st to suppress gaseous fires with similar or better

ginal goal was to obtain results using

estigations, this task became rather

gging of very small orifices (≈ 10 μm). Instead, an

et size

of about 20 μm, was employed. The actual droplet size distrib

prescribed distribution at the atomizer depending on the con he

counterflow burner, and this variation was characterized using a n

numerical model was extended to include such polydisperse d the

assumption that the droplet collisions were negligible based on the large separation distance between

droplets compared to their diameter. 

When the air stream was saturated with water vapor only, the counterflow non-premixed flame extinction

condition measurements agreed well with the corresponding numerical predictions. With addition of

condensed phase water droplets, the trends agreed well; however, considerable quantitative differences 

exist between the experimental results and the modeling predi tions. The observed differences were

shown to be mildly affected by the polydispersivity of the atomizer employed. This led to the conclusion 

that other submodels in the numerical model, including radiative heat losses, need to be evaluated. 

h vior

lame inhibition and extinguishment have been studied in a wide variety of laboratory flames, often with 

conflicting results. Chelliah and co-workers examined the effectiveness of water droplets and drop

droplet sizes, with a median dropl

ution was expected to deviate from the

vective velocity in the air tube in t

PDPA. The hybrid Eulerian–Lagrangia

roplet size distributions, subject to

c

Addition of NOH to water complemented the thermal fire suppression effectivenss of water. At low 

strain rates (≈ 125 s-1), 17.5% by mass of NaOH in the solution reduced the flame extinction strain rate by 

almost a factor of five. Inclusion of alkali metal compounds to water mist may not be desired in many

applications because of corrosive effects. Yet combining the chemical-acting powder agent with the

physical-acting liquid can enhance the suppression effectiveness of the liquid and provide a potentially

more convenient delivery method for the powder agent. 

Comparing Non-Premixed and Premixed Flame Suppression Be a

F

lets of
Aerosol Flame Suppression Studies 421

water with three alkali metal compounds in both premixed and non-premixed opposed flow flames, with 

the intent of constructing a unified phenomenology.128,129

Experimental 

Two apparatus were used, producing non-premixed and premixed methane/air flames. The burners were 

designed to be modular, so that the same water-generation and gas-supply systems could be used with 

either. The counterflow burner was as described in the previous section.98 As before, the water mass flow 

was fixed, and the air and methane flows were increased together until the flame was extinguished. The 

fuel and air flows were adjusted so as to balance their momentum; according to Equation (4-13). The 

global flow strain rate was defined by Equation (4-14) providing a suitable parameter that describes the 

non-premixed flame extinction condition.81 The total water (or solution) droplet mass fraction in the 

air + saturated water vapor stream was denoted as Y0, while the mass fraction of the additive (i.e., NaOH, 

KOH, NaCl, or FeCl2) was denoted as yadd. In the experiments, variation of aext with Y0 as well as yadd

were measured. 

The pr ebratype nozzle. A schlieren imaging system provided the flam , w to e 

burning velocity.77 The nozzle exit f as d b ple

measurements. The typ angle of the flame wa t 20o

. With addition of the solution 

d ght wa ntained constant at about 2.0 cm by adjusting total flow o tants 

hile maintaining the equivalence ratio at the desired value. For the present data, the expanded 

e factor of 2) in the burning velocity were ± 6.0 %. 

 a diameter range of 10 μm to 20 μm, except as noted. 

ver the 

0.055 and 0.112 NaOH cases (i.e., Y0 = 0.005 for yNaOH = 0.175 vs. Y0 = 0.010 for yNaOH = 0.055 and 

emixed burner consisted of a straight-sided conical flame established at the exit of a Mache-H

e cone angle

 characterize

s abou

hich was used

y laser Dop

 determine th

low profile w r velocimetry

ical half-cone

roplets, the flame hei s mai the f reac

w

uncertainties (with a coverag

The inhibition of a laminar premixed flame was characterized by the relative decrease in the burning 

velocity (denoted here by SL) with respect to that of an uninhibited flame (SL).The simple conical 

premixed flame stabilized above the Mache-Hebra nozzle, and the associated flame cone angle 

measurement technique adopted here is one of many traditional methods of obtaining the laminar flame 

speed.77 While curvature and stretch effects did exist in the flame, they were considered minor, 

particularly since all results were reported as normalized flame speed, S /S0. The uninhibited methane and 

air flame speed obtained from the present burner was 35 cm/s ± 1.2 cm/s. Because the ultrasonic 

atomizer and the length of the Pyrex tube delivering the reactant gas mixture were the same as in the 

counterflow burner, the median droplet diameter was not expected to be different from that of the 

counterflow flames. 

The ultrasonic atomizer was that described above. Droplets of pure water and water containing NaOH, 

KOH, and FeCl2 were generated with

Extinction of Non-Premixed Flames with Water Solutions 

Figure 4-62 shows a comparison of droplet mass fraction in air (Y0) as a function of the global extinction 

strain rate (aext), with varying NaOH mass fraction in water (yNaOH). Similar results were reported above 

for NaOH mass fractions in water of yNaOH = 0.055, 0.112, and 0.175.98 Earlier experiments at the lowest 

flow strain rate of 125 s-1 and 0.175 NaOH in water showed a significantly higher effectiveness o

0.112). However, for higher extinction strain rates (>125 s-1), the earlier experiments showed a negligible 

differences in extinction condition with varying NaOH mass fraction. 
422 Aerosol Properties 

The difference between the new yNaOH = 0.175 data and previous results was found to be caused by a 

temperature dependence of the ultrasonic nozzle performance. When the nozzle tip was allowed to reach 

a steady operating temperature, typically about 40 oC after about 15 min of continuous operation, the 

resulting flame extinction data were consistent and reproducible. While the revision of yNaOH = 0.175 data 

can be attributed to ultrasonic atomizer nozzle performance, the observation that the mass fraction of 

ting. The differences 

in the effectiveness with varying NaOH mass fractions pointed to the importance of the residence time in 

fixed strain rate (i.e. drop and resulting particle 

residence time) higher additive concentrations were expected to take longer to release all of the chemical 

Mass Fractions in Water. 129

OH mass fraction in the droplets. Also shown in Figure 4-63 is the mole fraction of 

NaOH in air (XNaOH), assuming that all of the NaOH in the water droplets was completely released to the 

perature and 

thermodynamics.

NaOH above 0.055 yields no apparent increase in flame suppression is rather interes

the flow field and the requirement that both the droplets evaporate and the residue particles evaporate or 

decompose. Only after the drops have evaporated can the residue particles begin to decompose in order 

to release the chemical suppression agent. Higher concentration of additives for the same drop size 

resulted in a larger particle after drop evaporation. At a 

agent. 

Figure 4-62. Droplet Mass Fraction 

as a Function of Extinction Strain 

Rate of a Methane/air Nonpremixed Flame for Several NaOH 

Also important to the degree of suppression effectiveness was a thermodynamic limit to the amount of 

resposnsible chemical catalytic scavenger that could be achieved in the gas phase at the flame conditions. 

Agent saturation behavior has been reported in premixed flames with iron, manganese, and tin 

organometallic compounds130,131 and metal salts added as particles.13 For sodium compounds, NaOH is 

the active chemical catalytic species for Na-based chemical suppression. Thus, the gas phase 

concentration of NaOH controlled the amount of suppression achievable. The flame extinction results 

shown in Figure 4-62 at the strain rate of 125 s-1 are replotted in Figure 4-63, according to the variation of 

droplet mass fraction (Y0) vs. NaOH mass fraction in water (yNaOH). As the yNaOH in the water was 

increased above 0.08, the total droplet mass fraction (Y0) needed for flame extinguishment became 

invariant with the Na

gas phase. As pointed out above, the amount of additive released is controlled by the residence time. In 

addition, the maximum XNaOH that can exist in the gas phase is controlled by tem
Aerosol Flame Suppression Studies 423

0

ction of NaOH in Air as a 

achievable. Assuming that NaOH exists as a monomer in the gas phase, an equilibrium liquid-vapor 

elded an NaOH saturation temperature of 

is about 400 K to 500 K below the characteristic 

tion, radical-production region of the flame.106

 

ating droplet, it is conceivable that an 

the saturation of excess NaOH vapor.

ay be controlled by a compound with a 

t

controls saturation of condensation of the relevant species is

detailed numerical simulations can provide better understanding

Although numerical simulations that include detailed interactio

gas phase have recently been developed to describe water dro odels for

droplet evaporation of binary mixtures in flames have not. In this case, modeling the evaporation of 

binary liquid droplets with highly disparate boiling temperatures (373 K for water vs. 1663 K for NaOH) 

may require further simplifications. Only experimental results are presented here. 

Besides NaOH, other alkali metal compounds are known to be chemically effective fire suppressants. In 

particular, potassium-containing compounds have been shown to be more effective than sodium.

13 

Figure 4-64 shows experimental extinction strain rates of a methane/air non-premixed flame with similar 

sized water/KOH droplets containing varying mass fractions of KOH. For the lowest extinction strain 

Figure 4-63. Y and Mole 

Fra

Function of yNaOH for the 

Extinction Strain of 125 s-1 in 

Figure 4-62. 129

Thermochemical data for NaOH were used to estimate the maximum amount of gas phase NaOH

calculation based on the Clausius-Clapeyron equation yi

1125 K.132 This vapor-liquid equilibrium temperature 

temperature within the rate-limiting oxygen-consump

Because of the thermal boundary layer surrounding each evapor

effective temperature below the flame temperature may control 

Alternatively, loss of active inhibiting species to condensation m

lower vapor pressure than NaOH. In experiments, resolving he appropriate effective temperature that

perhaps impossible, and therefore, only

of the saturation phenomenon occurring.

ns between the condensed phase and the

plet interactions,

80 analytical m

rate considered (i.e., 125 s

-1), the data show increasing flame suppression up to about yKOH = 0.112. 

Further increase in the KOH mass fraction in water up to 0.175, however, yielded no further increase in 

the flame suppression ability of the water/KOH solution. Following the analysis of water/NaOH 

saturation vapor conditions, current water/KOH results indicated that KOH saturates at a gas-phase mole
424 Aerosol Properties 

fraction of about xKOH = 0.0003 for yKOH = 0.112. Liquid-vapor equilibrium data for KOH indicated that 

this mole fraction corresponds to a gas temperature of about 1025 K, which is about 100 K lower than that 

obtained for the NaOH case. The lower effective saturation temperature for KOH was consistent with the 

lower boiling temperature of KOH (1597 K) compared to that of NaOH (1663 K). 

Figure 4-64. Droplet Mass 

Fraction as a Function of 

Extinction Strain Rate of a 

Methane/air Non-premixed 

Flame for Several KOH Mass 

Fractions in Water. 129

Non-premixed methane/air flame extinction experiments were also conducted with NaCl. Water/NaCl 

droplets have been considered previously by Zheng et al.,42 but only in the context of extinction of

counterflow premixed flames as a function of fuel-air equivalence ratio. The flame suppression trends 

shown in Figure 4-64, with increasing NaCl mass fraction in water, were consistent with previous data

with NaOH and KOH. 

Compounds containing the metal Fe has previously been shown to be very effective at flame inhibition.

133 

Here, exploratory experiments were conducted to test the efficacy of an Fe2+ compound dissolved in

water. An FeCl2 mass fraction of an yFeCl2 = 0.15 in water was tested, and this water/FeCl2 solution 

clearly indicated a chemical inhibition effect, as seen in Figure 4-65. Although FeCl2 was not as effective

as NaCl on a mass basis, experiments with lower mass fractions of FeCl2 must be performed to evaluate 

the occurrence of saturation phenomenon, as observed in water/NaOH solutions. 

In order to relate the chemical inhibition of the agents considered here to previous studies, the nonpremixed flame extinction results were analyzed on a molar basis of the additive. To avoid uncertainties 

related to saturation effects, only additive mass fractions below the saturation condition were considered. 

Since water/NaOH extinction results showed saturation of NaOH vapor was approached for additive mass 

fractions above 0.055, the molar comparisons between NaOH and KOH were performed at yadd = 0.055.

For the extinction data points considered, the mole fraction of alkali metal hydroxide (AmOH) was 

evaluated and is plotted as a function of flame extinction strain rate, as shown in the upper panel of

Figure 4-66. 
Aerosol Flame Suppression Studies 425

Figure 4-65. Droplet Mass

Fraction as a Function of 

These molar comparisons indicated a roughly double performance benefit of KOH over NaOH as an

additive to water droplets, similar to what has been reported based on flames inhibited with particulates.13 

Kinetic studies indicated that recombination of K with OH is about 30 % faster than Na with OH and that 

e/air Flame

Extinction Strain Rate of a

Methane/air Non-premixed 

Flame for NaCl and FeCl2 in 

Water. 129

recombination of K with O2 is in fact 2 to 3 times faster than Na with O2.

134 Thus, differences in the

reaction kinetics was perhaps the primary reason for the more effective flame suppression of KOH over

NaOH.

117

Figure 4-66. Upper panel: Mole 

Fraction of Alkali Metal atom in Air 

as a Function of Non-premixed 

Methane/air Flame Extinction Strain 

Rate for NaOH and KOH Mass 

Fractions of 0.055 shown in Figure 

4-63 and. Lower panel: Mole Faction 

of Am in Air as a Function of Nonpremixed Methan

Extinction Strain Rate for NaOH and 

NaCl Mass Fractions of 0.03 Shown 

in Figure 4-63 and Figure 4-66 

(upper panel). 129
426 Aerosol Properties 

Comparison of molar plots of NaCl vs. NaOH, shown in the lower panel of Figure 4-66, indicates a 

surprisingly superior flame suppression ability of NaCl over NaOH. The difference in effectiveness is not

due to the vapor pressure of the parent compounds, since NaCl has a lower vapor pressure than NaOH.

titative comparison between NaCl and 

derstand their suppression differences.

Although Cl can display an inhibiting effect similar to Br, a quan

NaOH requires additional studies and more detailed modeling to un

For a droplet mass loading of Y0 =0.014 and an extinction strain rate of 125 s-1, a comparison of the mole

fraction of chemical agent needed in air (Xi) yielded the following order: XKOH = 0.00021 < XNaCl = 

0.00023 < XNaOH = 0.00034 < XFeCl2 = 0.00056 (assuming linear interpolation between available data). 

These values indicated that on a molar basis KOH is the most effec ve chemical agent, followed by NaCl 

and NaOH. For water/FeCl2, the interpolation assumed that yFeCl2 = 0.15 is unsaturated at this high molar 

loading, which is certainly questionable. If FeCl2 does condense a lower additive mass fraction (e.g., 

yFeCl2 < 0.05), then on a molar basis water/FeCl2 would be the m would 

approach 0.00018 (instead of 0.00056 above). 

Extinction of Premixed Flames with Water Solutions 

The inhibition of premixed flame propagation with various che

documented.130,135,136 Chelliah and coworkers129

 investigated the inhibition of a conical premixed flame

with fine water/NaOH droplets. Based on phenomenological reasoning,137 it is well known that the

burning velocity is proportional to the square root of the chemical reaction rate. 

ti

at

ost effective solution as XFeCl2

mical fire-suppressing agents is well

2

0 ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ =

L

L

S

S S (4-17)

nsidered.

 of chemical inhibition 

Consequently, Figure 4-67 shows the square of normalized burning velocity vs. droplet mass fraction (Y0) 

for varying the NaOH mass fraction in water (yNaOH). With increasing yNaOH, these results do not show any

significant increase in flame inhibition compared to the inhibition with pure water droplets. This was a

rather unexpected finding because of the close similarities (described below) between the premixed and

non-premixed flames co

The insensitivity of higher NaOH mole fractions in water droplets in the premixed flame configuration 

can be explained by considering the droplet residence times in the flame. Because of the vast difference

in the boiling temperature between water and NaOH (373 K vs. 1663 K), water is expected to evaporate 

first. If the flow residence time of droplets through the premixed flame were less than that through a

counterflow flame, a partially vaporized droplet (smaller than the initial ~ 20 μm at the inflow boundary) 

with a much higher concentration of NaOH would emerge. Therefore, the lack

observed in the premixed experiments with water/NaOH solutions may be caused by partially vaporized

droplets. To test this hypothesis, the flow residence times of droplets through premixed and nonpremixed flames, inhibited with pure water droplets, were investigated. 
Aerosol Flame Suppression Studies 427

Figure 4-67. Square of the 

Normalized Burning Velocity of a 

P th 

Fine Droplets of Water Solution 

fo ions

in

mical reaction rate. A formal asymptotic analysis, assuming that the overall reaction of the 

form fuel + oxidizer → products is applicable to both premixed and non-premixed flames, yielded the 

following relationship between the burning velocity and extinction strain rate 

remixed Flame Inhibited wi

r Several NaOH Mass Fract

Water. 129

Comparison of Premixed and Non-Premixed Flames with Pure Water Droplets 

As described above, the premixed flame burning velocity is proportional to the square root of the overall

chemical reaction rate. The non-premixed flame extinction strain rate, however, is directly proportional

to the che 107

( ) ext

P st

L a

c Z

S ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

∝ 2

2 0

0

ρλ ρ (4-18)

where Zst is the stoichiometric mixture fraction (assumed l

conductivity, and cp the specific heat. Based on this relations ,

transport, and thermodynamic properties were not affected by

added,138 a direct comparison of the extinction/inhibition o

accomplished by defining normalized flame strength as

to be a small parameter), k the therma

hip and assuming that the mixture fraction

the small fraction of condensed-phase agent

f non-premixed and premixed flames was

⎟

⎟

⎠

⎞

hib

ib

Figure 4-68, which shows the normalized

function of the mass fraction of wate

equally effective i

⎜

⎜

⎝

⎛

= ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ =

ext unin

ext inh

L

L

a

a

S

S S

,

,

2

0 (4-19)

Its applicability for droplets of pure water was examined in

flame strength for the premixed and non-premixed flames as a r. The 

data indicate that the 13 μm median diameter droplets are not n inhibiting the premixed

flame. (Note that for proper normalizations, SL

0

 corresponds to a case in which the premixed methane/air

stream is saturated with water vapor, and aext,uninhib to a case in which the non-pre-mixed air stream is

saturated with water vapor.) 
428 Aerosol Properties 

Strength of Nonhe aforementioned differences between the inhibition of non-premixed and premixed flames by pure 

water droplets, as well as the disparate results for droplets of water/NaOH in premixed flame

explained based on the flow residence time as associated with each flame structure and its effect on the 

vaporation of fine water droplets. Figure 4-69 shows the numerically obtained flame structure of a 

iffered substantially: 4 ms for the premixed vs. 14 ms for the nonpremixed. This implied that the time available for the droplets to vaporize in premixed flames was 

Figure 4-68. 

Normalized Flame 

premixed and 

Premixed 

Methane/air Flames 

Inhibited with 

Droplets of Water 

with a Median 

Diameter of 

20 μm. 129

T

s, was 

e

premixed and a non-premixed flame, corresponding to conditions in which the flame was inhibited by 

13 μm droplets with a droplet mass fraction of Y0 = 0.01. 

The estimated flow residence time of these droplets through the thermal layer, from the cold boundary up 

to the peak flame temperature, d

considerably less, and that the 13 μm droplets were not necessarily the ideal size for inhibiting the 

premixed flame considered. Consequently, the maximum thermal cooling was not achieved compared to 

the counterflow flame, resulting in the lower flame inhibition observed in Figure 4-68 for premixed 

flames. 

Residence time considerations may also explain the much poorer relative effectiveness of either NaOH or 

KOH at a = 240 s-1 vs. a = 125 s-1, as shown in Figure 4-62 and Figure 4-64. The higher strain flames 

present lower residence times, possibly preventing release of all of the additive to the gas phase. 
Aerosol Flame Suppression Studies 429

Figure 4-69. Non

premixed and 

Premixed Flame 

Structures 

Corresponding to 

In

-

hibited Conditions of

Water Droplet Mass 

Fraction of Y0 = 0.01. 129

.5 % by mass) significantly

hanced the fire suppression ability of water by complementing its thermal effects with chemical

talytic radical recombination effects of NaOH. When the mas metal hydroxide in 

water was below this condensation limit, comparison of the chemical effectiveness indicated tha

ween flame types, as well as for evaluating the 

ficacy of chemically active additives. Comparison of the effects of water/NaOH droplets on the 

 non-premixed and inhibition of premixed flames implied that if the droplets we

completely evaporated before reaching the chemical reaction lay r (because of non-optimal dropl

fference was the cau mes. 

Conclusions and Summary

The low vapor pressure of alkali metal salts at normal room temperature requires their delivery to a flame

as a powder or as a droplet spray, if suppression is to be efficient. However, efforts to combine the 

thermal fire suppression ability of fine water droplets with the chemical inhibition of alkali metal

compounds indicated the existence of an upper limit for agent effectiveness perhaps due to condensation 

of some intermediate inhibitor species. Inclusion of NaOH in water (up to 17

en

ca s fraction of an alkali 

t KOH

was twice as effective as NaOH on a molar basis for a wide range of flow strain rates. Comparison of the 

effectiveness of water/NaOH with water/NaCl on a molar basis indicated the superior effectiveness of

NaCl over NaOH. 

The relationship between droplet size and flow residence time was found to be important, both for

comparing the behavior of condensed-phase agents bet

ef

extinction of re not

e et size

or insufficient droplet residence times), then the full chemical effectiveness of the agent was not realized.

When the same water droplets were introduced to non-premixed and premixed flames, with no velocity

lag between the droplets and the gas phase, the characteristic flame extinction/inhibition conditions of the 

two flames differed. Flame structure analysis revealed a distinct flow residence time of droplets (with a

median diameter of 13 μm) through each flame structure: 14 ms for the non-premixed flame, 4 ms for the 

premixed flame. This time di se for the observed differences in the two fla
430 Aerosol Properties 

The predicted results for polydisperse size distributions were consistent wit monodisperse 

predictions. However, differences observed between experiments and modeling could not be explained 

solely on the basis of the size distribution effects. 

The results illustrate the importance of understanding the particular reacting flow field and temperature 

conditions in order to assess the intricate coupling between droplet size and its residence ti

flame structure. In real fires with turbulent flow fields, the flow residence time ite 

varied; optimum droplet sizes might be different than the 10 μm to 20 μm observed in these studies. 

Therefore, design of optimal fire-suppression systems using droplets of water so r the 

flow residence times and flame structures of each application carefully. 

4.5.4 Dendrimers, Combined Chemical and Physical Flame Suppre

Introduction 

At the beginning of the NGP, dendritic polymers or dendrimers were a new po coming 

to fruition as commercial products. They are ultra-branched, three-dimensional polymeric molecules 

possessing a low-density core surrounded by a crowded, high density surface.139,140 Typical dendrimers 

were of the order of 5 nm to 10 nm in diameter. The interior of the dendrimer has tens to hundreds of 

branching junctions, depending on the overall degree of polymerization of the dendrimer. In the 

chemistries most commonly synthesized, these branching points were tertiary amines. Thus, the nitrogen 

at the branch junction has a lone pair of electrons that could be used for “complexation.” The ends of 

these branching paths are located on the dendrimer's surface and may also number in the tens to hundreds. 

The chemistry of these end groups is easily tailored to incorporate a desired chemical moiety or to 

l ion inside the dendrimer or by attaching it to a 

dendrimer surface, it might be possible to create a super-effective agent encapsulated in an inert, biols 

-received 

our derivatives of G4.0 dendrimers were synthesized and tested. They were designated 

4.SA (no metal added) or G4.SA-X, where X indicated the added metal ion: Cr, Fe, or K. Each of the 

 

h previous 

me through the 

s of droplets can be qu

lutions must conside

ssioniii

lymer technology

encourage complexation with a specific ion. 

Previous research had shown that transition metal complexes and alkali metal salts may act as superagents for fire suppression, performing as much as 60 times more efficiently than halon 1301.141,142 

Unfortunately, the most effective of the complexes also suffer from such problems as high toxicity and 

reactivity with metal surfaces. By encapsulating the meta

compatible shell that would decompose upon contact with flame. 

Beck Tan et al.

143 of the Army Research Laboratory (ARL) assessed the applicability of aqueous aerosols 

of dendritic polymerics for fire suppression effectiveness on gaseous laboratory flames (using the 

dispersed liquid agent fire fighting screen described in Chapter 8) and on liquid pool fires. 

Materia

Poly(amidoamine) dendrimers were purchased from Dendritech, Inc., of Midland, Michigan. Two types 

of dendrimers were used in this study. The G3.5-ONa dendrimers had 64 -COONa end groups per 

molecule and a nominal molecular weight of 14,000 g/mole. These materials were used in the as

condition. F

G

iii Text and figures in Section 4.5.4 are taken mainly from Reference 143 unless otherwise noted. 
Aerosol Flame Suppression Studies 431

added ions had indicated significant flame inhibition potential (Chapter 3). The base 

the G4.SA dendrimer was 19,300 g/mole. Each potassium dendrimer containe

other metal dendrimer contained 21 metal ions. Elemental analyses of the fiv

Table 4–9. 

Table 4–9. Elemental Composition of Pure Dendrimers and Dendrimer Salt Co

molecular weight of

d 64 potassium ions; each 

e dendrimers are given in

mplexes. 

Dendrimer 

Element Quantity G4.SA G4.SA-Cr G4.Sa-Fe G4.SA-K G3.5-ONa 

Mass % 8.08 6.94 6.81 6.7 7.38 H 

Atom % 53.09 51.2 51 50.9 52.5

Mass % 48.03 42.07 37.9 42.2 43.37 C 

Atom % 26.7 25.6 23.7 26.7 25.7

Mass % 14.7 12.76 16.63 12.61 12.9 N 

Atom % 7.0 6.72 7.29 6.84 6.55 

Mass % 29.1 33.6 37.3 28.7 29.4 O 

Atom % 12.5 15.5 17.5 13.6 13.1 

Mass % -- 4.63 -- -- -- Cr 

Atom % -- 0.66 -- -- -- 

Mass % -- -- 4.37 -- -- Fe 

Atom % -- -- 0.6 -- -- 

Mass % -- -- -- 9.8 -- K 

Atom % -- -- -- 1.9 -- 

Mass % -- -- -- -- 13.1 Na 

Atom % -- -- -- -- 2.14 

Fire Suppression Experimentation 

The fire suppression effectiveness for aqueous solutions of the five dendrimers was evaluated using two 

techniques. 

One set of tests involved the dispersed liquid agent fire suppression screen (DLAFSS) apparatus

developed at NIST for screening experimental fire suppressants; it is described fully in Chapter 8.

Briefly, the apparatus consists of a cylindrical, opposed flow, propane burner located in a vertical wind 

tunnel. A nebulizer was used as a droplet generation device. Liquid agents were entrained into the air

flow in the wind tunnel and traveled upward where they impinged on the propane flame. When the air 

sample without causing the sample delivery device to stall. 

velocity reached a critical value, the stable, blue enveloped flame at the burner was blown off, forming a 

wake flame. The transition to a wake flame was equivalent to extinguishing the fire. All experiments 

using the DLAFSS apparatus were conducted with the propane flow set at 2.0 L/min and the nebulizer air 

flow at 0.25 L/min. 

Some data were collected for solutions of 25 % (by weight) dendrimer in water and a sample delivery rate 

of 0.3 mL/min. However, due to the high viscosity of the solutions, the sample delivery pump stalled 

repeatedly. The remainder of the tests were conducted using more dilute solutions (15 % by weight) and 

increased sample delivery rate of 0.5 mL/min. These settings allowed for the same delivery rate of 
432 Aerosol Properties 

Although this was by far the best method found for screening experimental liquid fire suppression agents, 

the dendrimer residue frequently clogged the burner pores, necessitating changing the burner. This also 

added substantial variation to the test results. (The blow-off velocity for air alone was 220 cm/s. The 

elocity and water blow-off velocity for a given burner 

collected on a given day were averaged and used in calculating blow-off velocity ratios of water to air, 

the agent directly onto a 160 mm 

diameter pan containing 200 mL of JP-8. A Binks Model 15 paint spray gun was used as the agent 

 gun produced a 

fine stream of fluid droplets entrained in a forced air flow from a relatively small volume (50 mL) of 

ratus was adjusted so that the droplet spray distribution pattern covered the entire surface 

the spray gun was 4.6 mL/s, while the rate of collection 

sults 

dendrimer, dendrimer Na salts, dendrimer K salts, dendrimer Fe complexes, dendrimer Cr complexes, 

NaAc, K2CO3, chromium (+3) chloride hexahydrate, and ammonium iron (+2) sulfate hexahydrate. In the 

metal 

range of scatter between runs and for the different burners was about ± 20 cm/s. For water droplets, the 

blow-off velocity was 200 cm/s, with scatter of ± 20 cm/s.) To improve the ability to discern differences 

among the tested aerosols, data for air blow-off v

sample to air, and sample to water for samples that were evaluated using that same burner. Then, the 

ratios calculated for each sample were averaged. These average ratios were the parameters used to 

evaluate the candidate suppressant’s performance. 

The second method used was a modification of the JP-8 fuel fire suppression agent screening method 

developed by Finnerty at ARL.144 A test involved spraying a stream of 

delivery device, with a 1.58 mm nozzle operating at a pressure of 275 kPa (40 psi). This

fluid. The appa

of the fuel container. For water, the flow out of 

in the fuel container was 1.7 mL/s. 

The time to extinguish the fire was the criterion used to evaluate the effectiveness of various agents. 

Additional data included the volume of agent fluid remaining in the spray gun, and the volume of fluid 

remaining in the fuel pan. All experiments were recorded using a video camera, and exact fire-out times 

were extracted from the recordings. A photograph of an experiment in progress is shown in Figure 4-70. 

Figure 4-70. Setup for JP-8 Fire 

Suppression Experiments. 

DLAFSS Re

Fire suppression screening experiments were conducted on water and aqueous solutions of pure 

case of the salt solutions, the salt concentrations were adjusted so that the atom concentration of the 
Aerosol Flame Suppression Studies 433

ion was either equivalent to that in the corresponding dendrimer-metal complex solution (Na, Fe, and Cr 

salts) or set at 10 times the equivalent weight (K salts). 

The values of average blow-off velocities calculated for each agent screened are shown in Table 4–10. 

These values are similar to those collected by Yang et al.145,146 for a variety of potassium and sodium

small molecule salts. A lower blow-off velocity indicates a more efficient flame suppressant. 

Table 4–10. Average Blow-off Velocities for of Various Candidate Agents. 

Agent 

Blow-off Velocity (cm/s) 

@ 0.3 mL/min

Blow-off Velocity (cm/s) 

@ 0.5 mL/min

Sample-to-water 

Blow-off Ratio 

Water 1.00

10 % G4.SA 170 0.91

25 % G4.SA-K 152 0.82

4.8 % K2C03 189 0.98 

15 % G4.SA 193 0.96

15% G4.SA-K 180 0.94

3.23 % K2C03 204 0.98 

15 % G3.5-ONa 152 0.78

0.65 % CH3C02Na 171 0.87 

The results of the DLAFSS flame suppression experiments are summarized in Figure 4-71 and 

Figure 4-72. isinterpretation of the data, the water-to-air blow-off locity ratio m

during the sa itions as sample-to-air blow-off velocity ratio is given in each case. The 

effectiveness of the agent should be interpreted based on that reference. 

s the results of fire suppression experime conducted with a sample delivery rate

of 0.3 mL/min st interes g result from .3 mL/min ments wa om compar f the

5 % G4.SA-K and the 4.8 % K2 3 solutions. Although the K2 3 aerosol contained 10 times the K 

ns of the G4.SA-K dendrimer salt solution, the dendrimer solution was more efficient as a fire

the pure dendrimer solution, 10 % G4.SA, had fire 

to pure water. This was significant because there were initial

S results for the 15 % solutions and a sample delivery rate of 0.5 mL/min are summarized in

city ratio for water measured under the same conditions. The 0.65% NaAc, which

To avoid m ve easured

me cond the

Figure 4-71 summarize nts

. The mo tin

C0

the 0 experi

CO

s fr ison o

2

io

suppression agent. Another important finding was that 

suppression capacity that was comparable

concerns that the polymer might act as a fuel and thereby be detrimental to fire suppression effectiveness. 

The DLAFS

Figure 4-72. The results from small molecule K and Na salts are also included for comparison. The

results for the 15 % G4.SA solution indicate that once again, the addition of the pure dendrimer to water 

was not detrimental to fire suppression effectiveness. The fire suppression effectiveness of the 15 % 

G4.SA and the 15 % G4.SA-K samples was the same as water, within experimental uncertainty.

Comparison of G4.SA-K to the K2C03 solution showed that the dendrimer salt solution was perhaps a

little better than the pure salt, but the former contains one-tenth the K ion concentration. 

The agent that showed the most significantly enhanced fire suppression capability was the G3.5ONa 

dendrimer K salt. The blow-off velocity of the dendrimer-sodium salt was reduced nearly 20 % relative 

to the blow-off velo

had the same Na ion concentration as the G3.5-ONa dendrimer salt, also showed promise for enhancing 

the fire suppression effectiveness of water, but to a lesser extent than the dendritic salt solution. 
434 Aerosol Properties 

Reference Ratios for Water 

to Air Measured under the 

Same Conditions. Solid 

bars: Water to air; lined bars: 

sample to air.

The blow-off velocity ratios of sample fire suppression effectiveness to water fire suppression 

effectiveness measured when the same conditions were used are tabulated in Table 4–10. All ratios were 

comparable to or less than 1.0, indicating that all the experimental agents performed at least as well as 

water. In particular, the Na salts (both dendrimer and small molecule) and the dendrimer-potassium salt 

delivered at 0.3 mL/min from 25% solution were considerably more effective than water. 

Figure 4-71. Blow-off 

Velocity Ratios for Various 

Candidate Agents to Air, and 

Reference Ratios for Water 

to Air Measured under the 

Same Conditions. Solid 

bars: water to air; lined bars: 

sample to air.

Figure 4-72. Blow-off 

Velocity Ratios for Various 

Candidate Agents to Air, and 
Aerosol Flame Suppression Studies 435

In the case of the Na salts, it appears that the Na cation was the ac

salt and the dendrimer salt showed promise as fire suppression ag

tive species. Both the small molecule

ents. It is noteworthy that the NaAc 

the cation with the dendrimer "carrier." 

This is the type of effect that motivated the project and offers some proof of the concept that the use of

the Na and

K through coupling wit y 

was that ely high 

local concentration of them, h th rec on

believed to be responsible for flam This type of ad been her

situations in which an active species was the surface of a dendrimer.

147 A second possibility 

was t nd size distribution s being delivered to the flame were by the 

dendr way as to improve the fire suppression effectiveness of the fluid, based on purely

eter from a nebulizer, which was the sample delivery device 

incorporated into the DLAFSS device, is known to be related to the properties of the nebulizing fluid,

such rface tension, and viscosity.

148 The density of the dendrimer solutions was not

substa ent from that of water, and the surface tension of polymer and salt solutions was

gener by more than about 10 % to 20 %. However, the viscosity of dendri olutions

was m that of water. Ot er recent work on poly( ine) dendrime ndicated

at the viscosity of 15 % to 25 % dendrimer solutions at room temperature should be on the order of tens 

The idea that the viscosity effects may alter the fire suppression effectiveness through changes in the 

ld be of similar

showed an effect when present at only ≈ 0.65 % by weight. Since this salt is inexpensive, non-hazardous 

and readily available, it may be prudent to investigate its fire suppression effectiveness more fully. The 

G3.5-ONa dendrimer performed somewhat better than the small molecule salt solution of equivalent Na

ion concentration. A similar effect was observed in comparing the 25 % G4.SA-K and the 4.8 % K2CO3

solutions. Since the pure dendrimer solutions were not found to be significantly more effective than 

water, the observed effect must be related to the combination of 

dendritic polymers as carriers for active fire suppressant species could be beneficial. 

The mechanism responsible for observed enhancement in the fire suppression effectiveness of

h the dendrimer carrier cannot be easily extracted from the data. One possibilit

the organization of the cations on the dendrimer surface, which resulted in an extrem

ad a beneficial effect on e kinetics of the radical ombination reacti

e suppression. kinetic effect h observed for ot

localized on

hat the size a of the droplet altered

imers in such a

geometrical considerations. The droplet diam

as density, su

ntially differ

ally not changed mer s

uch greater than h amidoam rs

149 i

th

of centipoise, an order of magnitude greater than that of pure water (~1.0 cP).150 A third possibility may 

the thermodynamic saturation limits to the amount of inhibitor species permitted in the gas phase 

discussed in Section 4.5.3. 

delivery spray may explain the apparent discrepancy between the fire suppression effectiveness observed

for 25 % G4.SA-K dendrimer solution delivered at 0.3 mL/min and the 15 % G4.SA-K solution delivered 

at 0.5 mL/min (Figure 4-71 vs. Figure 4-72). The mass flow of polymer into the flame was identical, but 

the fire suppression effectiveness of the 25 % solution was considerably better than the 15 % solution. 

This may have been a reflection of the viscosity difference between the two fluids, which could be as

much as a factor of 6 to 8, and its effect on the droplet size and distribution in the sample delivery spray.

Although results by Yang et al.151 indicated that the droplet size and distribution generated in the

DLAFSS experiment are not strongly dependent on fluid viscosity in low viscosity solutions (~1cP to 

4 cP), they did not measure droplet size distributions in polymer solutions that have viscosities that are 

much greater than that of water. 

Although it is likely that viscosity played a role in the determination of apparent fire suppression

effectiveness, in the sodium salt and sodium dendrimer systems, chemical mechanisms are also probably 

contributing significantly. Both the 15 % pure dendrimer solution and the 15 % G4.SA-K solution were 

much less effective than the 15 % G3.5-ONa solution, though the viscosities shou
436 Aerosol Properties 

magnitude (see Table 4–10). Given this evidence of chemical interaction, it is in the sodium-containing 

systems where the kinetic benefits of localizing the active species on the surface of a dendrimer should be 

most readily apparent. The results from the G3.5-ONa solution provide the most concrete evidence that 

using a dendrimer carrier to deliver active species into a flame can have a positive effect on the fire 

suppression effectiveness of that active species. 

Pan Fire Suppression Results 

The analysis of data from the JP-8 fuel fire screening experiments was straightforward. The relevant 

parameters recorded were the time to extinguish the fire and the ma agent that was dispersed in 

extinguishing the fire. Because a relatively large mass of agent was required to perform each experiment 

and the results were not found to be highly reproducible, these exp he 

dendrimer-iron complexes that could not be studied using the DLAFSS apparatus because of their high 

viscosity. The associated transition metal salt and pure dendrimer con

[III] chloride hexahydrate, ammonium iron [II] sulfatehexahydrate, and G4.SA). The dendrimer 

chromium complexes could not be studied because the supply of materi

The results from the JP-8 fuel fire screening experiments are summarize -74. 

Although the masses of the various agents required to extinguish the JP-8 fires varied from 40 g to 65 g 

(Figure 4-74), the measurement uncertainty precluded drawing definitive conclusions regarding the 

relative fire-out times for any of the agents studied. It was not possible to judge the relative efficiency of 

the agents from these experiments also due to the discrepancy in the gent delivery rates. These rates 

were calculated from the mass deployed and the time of deployment and are plotted in Figure 4-75. The 

paint sprayer was delivering higher viscosity solutions at slower rates ifficulty in uniform agent 

dispersion for solutions of varying viscosity is a challenge to be addressed during the development of 

screening techniques for experimental fire suppression agents. 

F

V

S

2

e

s

ss of 

eriments were used only for t

trols were also studied (chromium 

al was insufficient. 

d in Figure 4-73 and Figure 4

a

. This d

igure 4-73. Times for 

arious Experimental Fire 

uppressants to Extinguish a 

00 mL JP-8 Pan Fire. The 

rror bars represent ± one 

tandard deviation.
Aerosol Flame Suppression Studies 437

Figure 4-74. Mass of Various 

Experimental Fire 

Suppressants to Extinguish a 

200 ml JP-8 Pan Fire. The 

error bars represent ± one 

standard deviation. 

Figure 4-75. Mass flows for 

Various Fire Suppressants 

Deployed during JP-8 

Screening Experiments. 
438 Aerosol Properties 

Because of the difficulties encountered in investigating the transition metal complexes using the JP-8 fuel

fire experiments, it was not possible to draw any strong conclusions regarding the fire suppression

effectiveness of the dendrimer-transition metal complexes. However, the experiment was sensitive

ssion effectiveness of water and can marginally improve 

the flame suppression effectiveness of water. Any contribution of polymer combustibility 

e suppression contribution. 

a

happen.152 Droplets with sufficient momentum to penetrate the flame,

be consumed in the flame (participate directly in the suppression

chemical mechanisms) or reach the fuel surface. Droplets with mentum (e.g., very small

droplets) may not penetrate the fire and will be deflected away by the rising hot plume.153 These 

deflected droplets will eventually evaporate or strike adjacent objects. For droplets landing outside the

burning area, the cooling of the adjacent and surrounding surfaces could mitigate or contain the flame 

spread. Therefore, droplet interaction with surfaces is an integral part of the overall suppression process. 

 

enough so that the effect of a super-agent would have been noticeable. This study indicated that the 

dendrimer-iron complexes did not exhibit the super-agent performance previously observed for iron

pentacarbonyl and other volatile transition metal complexes.143

Important Findings 

This exploratory study found that: 

• Poly(amidoamine) dendrimers added to water at substantial concentrations have no 

deleterious effect on the fire suppre

was at least offset by the flam

• The use of dendritic polymer carriers to deliver agents containing potassium and sodium ions 

can increase the effectiveness of these agents on laboratory flames.

• Dendrimer-sodium salts were more effective flame suppression agents than dendrimer

potassium salts. This is in contrast to the geater observed effectiveness of potassium over 

sodium in all other (non-dendrimeter delivery) suppression studies and may be due to 

saturation limits in the dendrimeter experiments. 

• The addition of sodium acetate to water improved its flame suppression effectiveness 

significantly, even at loadings of less than 1 % by weight. 

• Potassium carbonate did not significantly improve the flame suppression effectiveness of

water at concentrations as great as 5 % by weight for the droplet size and flow conditions

studied. 

• More work is needed in the development of versatile experiments for screening experimental

aerosol fire suppressants. 

4.6 SURFACE COOLINGiv 

pproach a fire, several things can 

 depending on their sizes, can either 

processes through physical and/or

little mo

4.6.1 Introduction 

When liquid suppressant droplets released from a discharge port

iv Large portions of the text of Section 4.6, as well as the figures , have been used from . 

The reader should consult the original document for additional details. 

Reference 152 without further attribution
Surface Cooling 439

Surface heat extraction and burning surface area reduction are generally regarded as the dom

mechanism for extingu 154

inant

ishment of burning solid fuels by water sprays. The cooler liquid droplets can 

interact with the relatively hot fuel surface, extract heat during vaporization, and thereby reduce the 

urface temperature. With surface cooling, the burning rate falls, which leads to a decrease in heat

entually, a critical burning rate is reached beyond which combustion can

o longer be sustained. The deposited droplets also reduce the mass transfer area on the fuel surface, 

hich also causes a reduction in burning rate. 

In order to understand surface cooling without resorting to a detailed analysis of the coupling between the 

transient thermal response of the fuel bed as a result of the application of liquid agent and the heat

feedback from the flame to the fuel, the transient coupling process is considered as a series of quasisteady steps. At any instant in time, a steady-state energy balance at the fuel surface can be written as: 

(4-20)

where is the fuel burning mass flux, 

s

feedback to the fuel surface. Ev

n

w

m H Q Q T T q fuel fuel cond conv flame flame s surf ′′ ∆ = ′′ + ′′ + − − ′′ 4 4 & ε σ ε σ

m fuel & ′′ ∆H fuel is the effective heat of gasification of the fuel, Qcond ′′ is

the conductive heat flux to the fuel surface, Qconv

′′ is the convective heat flux to the fuel surface, flame ε is 

the flame emissivity, σ is the Stefan-Boltzmann constant, Tflame is the average flame temperature, s

surface emissivity, Tsurf is the fuel surface temperature, and q" is the heat flux removal from the fuel 

surface due to surface cooling. 

The surface cooling heat flux, q", depends on the interaction of the liquid droplets with the surface. There

are many different scenarios of droplet impact on a surface. The impacting processes depend on the 

thermophysical and geometrical properties of the droplet and of the target surface and on the magnitude 

and direction of the impact velocity vector. In addition, for a flaming surface, the droplet will traverse 

through an environment different from that of a non-burning surface before impact. However, in the 

literature, droplet/surface interaction studies focus almost exclusively on non-burning surfaces. 

When relatively cool liquid droplets impact a solid surface, cooling of the surface occurs as a result of

heat transfer between the surface and the droplets. A liquid droplet impacting a solid surface has been

extensively studied in the literature.155,156,157,158,159,160,161,162,163 Maps that identify the various impact and 

heat transfer regimes over a wide range of experimental conditions have been constructed. Spray cooling 

or quenching of hot metallic surfaces has also been extensively studied.164,165,166,167,168,169 Many empirical 

heat transfer correlations for spray cooling applications exist. 

For a liquid droplet interaction with a liquid surface, several scenarios need to be considered. The impact

dynamics are complicated by the ability of the liquid surface to deform and to displace in response to the 

impinging droplet. There is also internal fluid motion of the liquid pool initiated by the impact

processes.170 If the droplet has a very high impact velocity, it can cause a crater in the liquid pool and a

splash upon impact. For a burning liquid pool, the splash may result in a sudden increase in burning rate 

due to an increase in the liquid surface area caused by the splash droplets. If the droplet is soluble in the 

liquid, mass transfer between the penetrating droplet and the liquid will occur. Depending on droplet

impact parameters, differences in miscibilities and physical properties (surface tension and viscosity)

between the droplet and the target liquid, and the depth of the liquid layer, several phenomena are likely

following impact. The droplet may float on the surface, may coalesce into the liquid pool with little or no 

splashing and produce a vortex ring, may rebound or splash forming a crater and a small jet or column of

liquid at the surface, or may dissolve into the liquid pool. Most work in the literature has examined the 

ε is the 
440 Aerosol Properties 

isothermal impact of a liquid droplet onto a surface of

Rein170

 and the references therein). In a fire suppression scenario, 

different liquid target (suppression fluid vs. fuel) with different phy

the droplet can be much lower than that of the liquid surface (espec

contact between a cold droplet with a hot liquid may some

unwanted situation. All of these complicated interactions render

 the same liquid (e.g., see the review article of

the impacting droplet encounters a 

sical properties. The temperature of 

ially, in the case of burning), and the

times results in vapor explosion, 171 an 

t

between liquid droplets and a liquid surface an arduous and daunting task. Appropriate heat transfer 

correlations do not exist in the literature; therefore, liquid surface cooling calculations, in spite of the 

importance of the process, are not attempted here. 

This section discusses droplet evaporation and solid surface cooling the

surface cooling problem. If the droplets evaporate completely before reaching the fuel surface, the heat 

removal from the surface will decrease significantly due to the fact that the liquid latent heat of

vaporization does not play a role. Under this circumstance, the surface will only be cooled by a vapor jet,

which is not an effective cooling mechanism.

Another important aspect in surface cooling is the ability of the droplets to reduce the surface temperature 

of components in the vicinity of a fire to prevent re-ignition, which is a potential problem in suppressing 

aircraft engine nacelle fires. If sufficient cooling of any hot surfaces in the vicinity does not occur, 

re-ignition of residual fuel vapor/air mixture may result. 

In the following discussion, the five fluids (water, lactic acid, C F H3O, HFE-7100, and R338mccq) 

identified in Pitts et al.152 as having the highest mass-based al

he study of heat transfer processes

by droplets, an integral part of

3 5

tot ∆H (total heat absorbed including 

enthalpy of evaporation in going from room temperature to 1400 K) were considered to evaluate their 

. Low boiling-point fluids were not considered

 flash-vaporize soon after they leave the discharge 

ic let

diameter as a function of time as 

(4-21)

where D is the instantaneous droplet diameter, Do is the initial droplet diameter, t is the time, and K is the 

evaporation constant. The evaporation constant was calculated by

droplet vaporization and surface cooling characteristics

because under typical operating conditions, these fluids

opening and arrive at the surface in the form of a vapor. 

4.6.2 Evaporation of Liquid Droplets 

The estimation of liquid droplet evaporation is based on the class al d2-law, which gives the drop

, 2 2 D D Kt = o −

, ρ

8 ln(1 )

F pg

g

C

k B

K

+ = (4-22)

where B is the transfer number, kg is the gas-phase thermal conductivity, Cpg is the gas-phase heat 

capacity, and ρF is the liquid density. The B number based on mass transfer considerations is given by

, 1 FS

FS F

M Y

Y Y B −

− = ∞ (4-23)

here YFS and YF∞ are the vapor-phase mass fraction at the droplet surface and ambiance, respectively.

action, YFS, was evaluated by

w

The mass fr
Surface Cooling 441

, ( ) FS F FS A

FS F

FS P M P P M

P M Y

+ − = (4-24)

where PFS is the liquid vapor pressure at the droplet surface, P is the ambient pressure, which is equal to 

the sum of the liquid vapor pressure and partial pressure of air at the droplet surface, and MF and MA are

the molecular weights of liquid and air, respectively.

The B number based on heat transfer considerations was given by

, ( )

vap

pg s

T H

C T T

B

∆

− = ∞

(4-25)

where T T

heat of vapo

Equation (4

If Ts is kno

circumstanc calculation procedure can

be foun n

and Kanury

curves defin

and Lefebvr r approach.

To calc t

respectively te the gas phase thermophysical 

properties. 

∞, s, and ∆Hvap are the ambient temperature, droplet surface (wet-bulb) temperature, and latent 

rization of the liquid, respectively. Under steady-state conditions, BM = BT, and either 

-23) or Equation (4-25) may be used to calculate the evaporation constant in Equation (4-22). 

wn, the calculation of the evaporation constant is straightforward; however, under normal 

es it is necessary to determine Ts in order to evaluate B and K. The

d i References113, 172 and 173. The basic idea was to find Ts such that BM = BT. Spalding172

173

 used a graphical approach. By plotting BM and BT against Ts, the intersection of these two 

ed the wet-bulb temperature and the corresponding B. An iterative method was used by Chin 

e.

113 The calculations presented here used the latte

ula e K, the reference temperature (Tr) and mass fractions (YFr and YAir for the fluid and air, 

), based on the one-third rule,113 were used to estima

3 r s

T Ts T T − = + ∞ (4-26)

3

F FS

Fr FS

Y Y Y Y − = + ∞ (4-27)

1 . YAir = −YFr (4-28)

The equation for calculating the gas-phase heat capacity, Cpg, was given by

( ) ( ) ( ) Cpg Tr =YAirCpA Tr +YFrCpF Tr (4-29)

where CpA is the heat capacity of air and CpF is the gas-phase heat capacity of the fluid, all evaluated at Tr. 

The thermal conductivity of the gas phase was estimated using the following equation, 

( ) ( ) ( ), g r Air A r Fr F Tr k T =Y k T +Y k (4-30)

where kA is the thermal conductivity of air and kF is the gas-phase thermal conductivity of the liquid, all 

evaluated at Tr. 
442 Aerosol Properties 

The conditions used in the calculations were: (1) T∞ = 1400 K (to simulate a flaming ambiance),

(2) YF∞ = 0, and (3) P = 0.101 MPa. The thermophysical properties of air, water, and lactic acid were 

obtained from the database of the Design Institute for Physical Properties I.174 The properties of

C3F5H3O, HFE-7100, and R338mccq were estimated by methods described in Reid et al.,175 the modified

Rackett technique was used for ρF, the method of Joback for CpF, the method of Lee and Kesler for PFS, 

the method of Pitzer’s acentric factor correlation for ∆Hvap, and the method of Ely and Hanley for kF. 

Figure 4-76 shows droplet evaporation times, tb, as a function of initial droplet diameter. The droplet

evaporation time was easily obtained by assigning D = 0 in Equation (4-21) once K is known,

.

2

K

D

t o

b = (4-31)

For a given initial droplet diameter under the same ambient conditions, the evaporation time for a water 

roplet was the longest and for a R338mccq droplet the shortest. For the five fluids examin

bas was

d ed, the

ranking ed on droplet evaporation times

. , 338 , 7100 , , , b R mccq b HFE b C3F5H3O b lacticacid b water t <t <t < t <t (4-32)

How fast a droplet evaporates determines whether it will still remain as a droplet upon reaching the 

surface. However, if flame cooling is the dominant fire suppression mechanism, then it would be better to 

have the droplet completely vaporized in the flame. 

Figure 4-76. Droplet Evaporation Time as a Function of Initial Droplet Diameter for Five 

Fluids. 

Droplet Initial Diameter (μm)

0 10 20 30 40 50 60 70 80 90 100

Evaporation Time (s) 0.00

0.01

0.02

0.03

0.04

0.05

0.06

Water

Lactic acid

C3F5H3O 

HFE7100

R338mccq

Ambient temperature @ 1400 K
Surface Cooling 443

4.6.3 Spray Cooling 

In order to assess surface cooling, we consider the literature on spray cooling and quenching of hot 

metallic surfaces. The application of spray-cooling correlations to fuel surface cooling is only appropriate 

when the fuel surface is solid (e.g., a PMMA or wood slab). Even in this case, care should be exercised 

because subtle differences in cooling mechanisms exist. A burning PMMA surface in reality is in a 

molten state, and wood forms a porous char layer. The interactions of droplets with molten and porous 

surfaces remain very difficult problems to tackle. Differences also arise between fuel surface cooling and 

quenching of metallic surfaces. In the latter, radiative heat transfer to the droplets only occurs from the 

heated surface; however, in the former, it also occurs from the flame. Strictly speaking, the following 

s which can be followed along the “boiling curve.” Such a 

curve is a plot of surface heat flux vs. surface temperature and represents the strong relationship between 

ated with spray cooling of a hot surface. The 

general features of the curve are similar to those for a pool boiling curve. 

id 

Droplets. 

Since the boiling curve of one fluid may differ from that of other fluids, there are several im

calculations are more applicable to the cooling of a hot surface to prevent re-ignition than to the cooling 

of a burning fuel surface. 

Depending on the surface temperature, when a hot surface is being quenched by a liquid spray, it will 

experience several distinct heat transfer regime

heat flux and surface temperature during quenching. The characteristics of the boiling curve are classified 

as (with increasing surface temperature): (1) convective, (2) nucleate boiling, (3) critical heat flux, 

(4) transition boiling, (5) Leidenfrost point, and (6) film boiling. Detailed descriptions are available.164,165 

Figure 4-77 is an illustration of a boiling curve associ

Figure 4-77. Typical Boiling Curve Associated with Quenching of a Hot Surface by Liqu

portant 

assumptions that have to be made in the following analysis. 

Surface Temperature

 Transition boiling

Conv r e ective egim

Film boiling

Nucleate boiling

Surface Heat Flux 

Critical h

Onset of

nucleate boiling

eat flux (CHF

Leidenfrost point

)
444 Aerosol Properties 

• When a comparison of surface heat fluxes is made, it is assumed that for the prevailing 

surface temperature, these fluids are in the same me on the boiling curve. This 

prevailing surface temperature. Unfortunately, in 

most cases, it is not possible a priori to determine the demarcation point (surface 

r without resorting to experimental 

observations.

• The second assumption is that the dimensionless heat transfer correlations obtained

literature using water (in most cases) are applicable to the other fluids considered here. 

ot

s were

parameters, though encountered in spray quenching applications, may not be identical to

), and the Prandtl number (Pr) by the following expression, 

 regi

assumption may not be valid because, for example, the surface temperatures at which the 

critical heat flux and Leidenfrost phenomena occur may be significantly different among the 

fluids studied; therefore, while one fluid may be at the critical heat flux, the other fluid may 

be in the nucleate boiling regime at the

temperature) from one cooling regime to anothe

in the

• A third assumption is that the effect of surface roughness on surface heat flux is n

considered. Furthermore, the spray parameters used in the following calculation

chosen to be within the range of applicability of the empirical correlations. These operating

those appropriate for fire fighting applications. In addition, since different droplet generation 

techniques (sprays vs. mono-dispersed droplet streams) have been used to obtain the heat 

transfer correlations in various regimes, it is imperative that a comparison among the

calculated surface heat fluxes should be made with these fluids in the same regime. 

4.6.4 Convective Regime 

Mudawar and co-workers164,165 had performed extensive studies on cooling of metal surface in this regime 

using water sprays. Their heat transfer data were correlated using the Nusselt number (Nu), the Reynolds 

number (Re

2.512 2.512 , 0.76 0.56

0.76 0.56

32 32 Re Pr

k

Q D C

Nu

k

hD

f

pf f

f f

= ⎥

⎥

⎤

⎢ ⎦

⎢

⎣

⎡

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ′′ = = μ

ν

here h is the heat transfer coefficient, kf is the thermal conductivity of the liquid, νf is the liquid 

kinematic viscosity, Cpf is the liquid heat capacity, and μf is the liquid viscosity. The Reynolds nu

ased on the volumetric spray flux, Q", and the Sauter mean diameter, D32, of the spray. All the liquid 

roperties are evaluated at (Tsurf + Tf)/2 where Tf is the droplet temperature before impact. 

calculations were performed

f 32 sical properties of water and

ctic acid were obtained from the DIPPR.174

 The properties of C 5H3O, HFE-7100, and R33

ere estimated by methods described in Reid et al.,175

 the boiling point method of Sato was used f

32 culations) for

e five fluids because the atomization characteristics at the spray nozzle would be vastly different

e fluids owing to their thermophysical properties, thus resulting in different D32.

66

(4-33)

w

mber is

b

p

Figure 4-78 shows the calculated heat flux as a function of Tsurf - Tf. The

using Q′′ = 0.005 m3

/m2-s, T = 295 K, and D = 0.5 mm. The thermophy

la 3F 8mccq

w for k ,

the corresponding states method of Rowlinson for Cpf, and the method of Brule and Starling for μf. In this 

regime, water removed heat from the surface more efficiently than the other four fluids. It may not be 

feasible in practice to maintain the same D with a fixed Q′′ (the conditions used in the cal

th among

th
Surface Cooling 445

4.6.5 

Comp me in spray cooling has 

probably been the least stu surface was found to scale 

only with Tsurf-Tf in the stu ray was used. The heat

flux, q", was correlated in the form

(4-34)

In the work of Qiao and Chandra176 on nucleate boiling enhancement using a surfactant/water solution,

the heat flux measurements could also be correlated well by scaling with Tsurf - Tf in a form similar to

Equation (4-34), 

(4-35)

ssuming that the proportionality constant and the power in the scaling law do not change significantly 

a

rem

given Tsurf. 

Figure 4-78. Calculated Boiling Curves in the Convective Regime for Five Liquids. 

Nucleate Boiling Regime 

ared to other regimes on the boiling curve, the nucleate boiling regi

died. In this regime, the heat flux removal from the 

dies by Mudawar and co-workers164,165 where a water sp

 of 

1.87 10 ( ) . 5 5.55 Tsurf Tf q′′= × − −

0.56 10 ( ) . 5 6 Tsurf Tf q′′= × − −

A

mong the five chemical compounds and that the heat flux can be scaled with (Tsurf - Tf), then the heat flux

oval from the surface in the nucleate boiling regime using these five liquids should be similar for a 

Tsurf - T (K) f 

0 20 40 60 80 100 12

q" (W/m2

) 104

105

106

107

Water

Lactic acid

C3F5H3O

HFE7100

R338mccq
446 Aerosol Properties 

4.6.6 Critical Heat Flux

The dimensionless correlation developed by Mudawar and Valentine164 for water spray cooling over a 

wide range of flows was used to estimate the critical heat flux (CHF), 

, ρ

σ

ρ

ρ ( )

ρ

ρ 122.4 1 0.0118

ρ

0.198

32

2

1/ 4

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

′′ ⎪⎭

⎪

⎬

⎫

⎪⎩

⎪

⎨

⎧

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

∆

−

⎥

⎦

⎤ ⎢

⎣

⎡ = +

∆ ′′ H Q D

C T T

H Q

CHF

F

f

g vap

F pf sat f

F

g

g vap

(4-36)

where ρg is the vapor density of the fluid, Tsat is the saturation temperature of the fluid, and σf is the 

surface tension of the fluid. All the fluid properties were evaluated at the fluid saturation temperature. 

3 2 Using Q" = 0.005 m /m -s, Tf = 295 K, and D32 = 0.5 mm, the resulting CHFs for water, lactic acid,

4.6.7 Transition and Film Boiling Regimes 

Based on dimensional analysis, Deb and Yao177 statistically derived an equation for heat transfer 

effectiveness (ε) in terms of droplet Weber number (We), dimensionless wall superheat (Bw) and vapor

parameters (Kd), and surface factor (SF) for surface material effect, 

C3F5H3O, HFE-7100, and R338mccq are 0.63 × 107

 W/m2, 0.86 × 107 W/m2, 0.33 × 107 W/m2, 

0.23 × 107 W/m2, and 0.21 × 107 W/m2, respectively. In the calculations, surface tensions of water and

lactic acid were obtained from the DIPPR,174

 while those of C3F5H3O, HFE-7100, and R338mccq were 

estimated using the method of Brock and Bird.175

 Although lactic acid and water have the highest and 

second highest CHFs, respectively, under the conditions used in the calculations, it should be noted that it

is inappropriate to compare the CHFs unless they occur at the same surface temperature for the five

fluids, a situation highly unlikely given the differences in the thermophysical properties of these fluids. 

, 1

90 0.21 exp

60.5

1

35

0.08 ln

=0.027exp 1.5 ⎥

⎦

⎤ ⎢

⎣

⎡

+

+ −

⎥

⎥

⎥

⎥

⎥

⎦

⎤

⎢

⎢

⎢

⎢

⎢

⎣

⎡

⎟

⎠

⎞ ⎜

⎝

⎛ +

⎟

⎠

⎞ ⎜

⎝

⎛ +

ε

We

K B

SF B

We

d w

w

(4-37)

with 

f

F Vi Do We

σ

ρ

= (4-38)

( )

vap

pF surf sat

w H

C T T

B

∆

− = (4-39)

pF v

F

d C

k K

μ = (4-40)

and 
Surface Cooling 447

( ) −1, ρ

ρ =

steel p

s s ps

k C

k C

SF (4-41)

where Vi is the droplet impact velocity, K

ks, ρs, and C

All liquid an

of actual he

the droplets. the surface is 

d is the dimensionless vapor parameter, μv is vapor viscosity, and 

ps are the thermal conductivity, density, and heat capacity of the surface material, respectively. 

d vapor properties are evaluated at Tsat. The heat transfer effectiveness is defined as the ratio

at transfer from the hot surface to the total heat transfer required for complete evaporation of 

Therefore, the heat flux from

[ ( )], G Hvap Cpf Tsat Tf q′′=ε ∆ + − (4-42)

where G is

experiments

regime and

Figure 4-79

The calcula

illustration, ed in the

calculations, the resulting droplet We were 31, 97, 181, 297, and 339 for water, lactic acid, C3F5H3O, 

tively. The range of surface temperatures used in the calculations was

lated Boiling

Curves for Five Liquids for the 

the liquid mass flux. ε was obtained from Equation (4-37) by fitting data mainly from 

using a single stream of droplets. The experimental conditions covered the film boiling

to some extent the adjacent transition boiling regime and Leidenfrost point. 

shows the calculated heat fluxes for the five fluids as a function of the surface temperature. 

tions were performed using Tf = 295 K, Vi = 2.5 m/s, Do = 0.3 mm, and G = 0.5 kg/m2

-s. For 

a stainless steel, SS 304, was used as the target surface. Under the conditions us

HFE-7100, and R338mccq, respec

chosen so that the calculated ε did not exceed 1. The three boiling regimes (transition boiling, 

Leidenfrost condition, and film boiling) are apparent in Figure 4-79. If it is assumed that at Tsurf ≥ 800 K, 

all five fluids exhibit film boiling, and lactic acid has the highest surface heat flux. 

Figure 4-79. Calcu

Transition and Film Boiling 

Regimes. 

Tsurf (K)

300 400 500 600 700 800 900 1000

q" (W/

m2

) 103

104

105

106

Water

Lactic acid

C3F5H3O

HFE7100 

R338mccq 
448 Aerosol Properties 

Figure 4−79 also demonstrates again the difficulty in comparing the heat fluxes at a given surface 

temperature. For example, at Tsurf = 550 K, C3F5H3O, HFE-7100, and R338mccq are in the film boiling 

regime, whereas, water and lactic acid are in the transition boiling regime. As stated above, unless the 

surface temperature at which the onset of a particular boiling regime is known a priori, it is not possible to 

know which heat flux correlation pertaining to which boiling regime should be applied in the estimation. 

4.6.8 Summary 

A suppression mechanism possibility for aerosol agents (and not generally for gaseous agents) is the 

potential to cool any burning surfaces or hot surrounding surfaces. Surface cooling by liquid droplets is a 

very complicated heat transfer process because it depends on so many parameters: the thermophysical 

properties (surface tension, density, viscosity, heat of vaporization, etc.) of the liquid, droplet size, droplet 

impact velocity, spray mass flux, and surface properties. Even for non-burning, smooth surfaces, 

calculations of surface heat fluxes from first principles are not possible without resorting to experimental 

observations. Empirical heat transfer correlations under conditions encountered in fires are lacking, and 

the quantification of surface cooling as a means to extinguish the burning fuel remains elusive, especially 

when the fuel is a liquid. Nonetheless, the greatest benefit will arise from those agents that are efficient 

thermal agents, particularly aerosols that have large heats of vaporization per gram of agent. 

4.7 AEROSOLS AND BURNING SURFACES 

here are four key physical phenomena involved in the interaction of aerosols with burning surfaces: 

(1) buo

traje on 

flaming surface combustion, and (4) reduction in surface temperature caused by the particle impingement, 

tion on surface. The NGP study on the interaction of aerosol with 

 eva

d here and can also be found in References 178, 179, 180 and 181. 

onte Carlo methods are widely used to solve for the heat transfer in a wide variety of problems182

 their simplicity and accuracy. There are many methods to improve the accuracy of Monte Carlo 

ce the important rates that govern the physical processes are understood. 

he different components that describe both the droplet trajectory and the temperature distributio

e surface as a function of time are studied in detail in this project by separating the various physical 

hing a burning surface. Evaporation effects were also examined. Both monodisperse and 

polydisperse droplets were injected over PMMA flames or over heated polyurethane foam, using droplet 

T

yancy effect on the trajectory of the aerosol particles, (2) evaporation/decomposition effect on the 

ctory of the particles, (3) cessation of reaction and reduction in flame spread caused by the aerosol

spreading, and evaporation/decomposi

burning surfaces addressed the interactions of liquid droplets with surfaces. The NGP study 

experimentally determined the key water droplet properties, such as surface loading and droplet diameter, 

that affect flame spread rate, and through numerical and experimental comparison sought to determine the 

effect of droplet size and velocity on droplet buoyancy and poration as the droplets approach a hot 

surface or fire. The work is summarize

M ,183 due 

to

simulations including improved sampling,184 weighing and biasing,185 and the discrete probability 

function method.186 Therefore, the simulation of the problem utilizing Monte Carlo methods is 

straightforward, on

T n over 

th

processes and integrating them to obtain an experimentally validated model that can be used in Monte 

Carlo simulations. Experiments were conducted with non-evaporating and evaporating particles of 

controlled density injected at controlled velocities towards a variably heated surface. Particle velocity as 

a function of the impressed buoyancy forces were measured, thereby determining the likelihood of a 

particle reac
Aerosols and Burning Surfaces 449

loading, delay time after ignition, and location of the spray from the leading edge as parameters. Such 

conditions provided data for examining any correlation between spray characteristics and flame spread. 

e interactions were investigated to 

r understanding of the dominant processes and identify the key aerosol properties that impact 

ficient suppression of the fire. 

.7.1 Flame Spread Studies 

er 

of 64 and was sufficient to establish a fully developed laminar flow in the straighteners.187 A 0.16 cm 

 tunnel to reduce heat transfer to the walls. 

Figure 4-80. Wind Tunnel Used for 

Fire Spread Studies. 

MMA sheet was used as the fuel. PMMA supports surface co bustion, and the flame spread rate

is steady and well characterized. The properties of PMMA, such as density, specific heat at constant 

ressure, and heat of combustion, are consistent and well known. PMMA test specimens were 30 cm x 

0 cm sheets of thickness 0.32 cm or 0.64 cm. The PMMA sheets were ignited with a small methanol 

trough before placing them into the wind tunnel. The experiments were carried out at four opposin

/s, 84 cm/s, 121 cm/s, and 153 cm/s. Three air flows, ranging from 9.5 cm3

/min to 

cm3

/min, and three inlet pressures (880 kPa, 1375 kPa, and 2115 kPa) were used to vary the mass 

loading and the droplet size of the water spray used for suppression. 

Surface temperature changes as a function of aerosol impingement were determined using a Two Color 

Planar Pyrometer. Finally, Monte Carlo simulations of the mist/surfac

develop fulle

ef

4

Experimental 

Flame spread experiments were carried out in a wind tunnel, whose photograph is shown in Figure 4-80. 

The wind tunnel measured 45 cm high, 76 cm wide, and 175 cm long. The roof had three rectangular 

slots 7.6 cm wide x 28 cm long for observation of the flame by means of a video camera. The slots were 

covered with 1.25 cm thick glass windows during the experiments. The velocity inside the tunnel was 

controlled by an exhaust fan. 

Honeycomb flow straighteners of polycarbonate, placed at the mouth of the duct, were used to remove 

large scale turbulent eddies and create uniform flow conditions in the test section. Each flow straightener 

had a diameter of 0.64 cm and is 25 cm in length. This L/D ratio, yielded a maximum Reynolds numb

thick ceramic paper was applied to the interior of the

P m over it 

p

2

g wind 

speeds of 69 cm

15.5 
450 Aerosol Properties 

Initially, the fire suppression experiments were carried out with one surface in contact with an insulating 

ceramic material. This is the conventional method, called here “with backing” mode or on

combustion mode. When fire suppression experiments were begun, it was realized that the ceramic 

acking was quick to absorb water and was instrumental in extinguishing the flame almost as soon as the 

 Mist. 

he aluminum platform 

has a maximum value of about 1.4 cm. A 6 cm gap between the bottom surface of the PMMA specimens 

and the support platform ensured that free stream conditions existed on both the top and bottom and that 

the flame spread was unaffected by the boundary layer on the suppo platform. 

A hollow cone spray was generated using a pressure atomizer which delivered good atomization quality 

over a large range of flows at the same inlet pressure. The nozzle was mounted on a bracket at a distance 

of about 60 cm upstream of the sample and at a height of about 17.5 cm above the surface of the PMMA. 

The droplets were carried downstream and onto the flame by the airflow present in the wind tunnel. 

To obtain local surface loading of the water spray, pieces of absor

surface of a PMMA sample in a 4 x 4 grid. The spray was established at a given test condition and the 

difference in weight of the tissues was determined after a suitable interval of time. A Phase Doppler 

Particle Analyzer was used to obtain the droplet size and droplet ean velocity data. The PDPA was 

operated in forward scatter mode (θ =30"). The probe volume was 14 cm downstream of the nozzle exit. 

A black-and-white 8-bit CCD camera was used to track the motion of the flame. The camera was 

mounted on a bracket above the observation window and looked down perpendicularly onto the burning 

samples. It was observed that the edges of the PMMA samples burned faster than the center. In order to 

avoid erroneous estimates of flame spread rate on account of edge effects, the camera focused on the 

central 23 cm of the flame. 

e-sided 

b

water spray was turned on. Therefore, the present studies were carried out with the PMMA sample 

supported along its sides with air flowing above and below the specimen. A photograph of the flame 

suppression experiment with the water mist applied is shown in Figure 4-81.

Figure 4-81. Flame Spread 

Experiment with Water

Calculations showed that the Blasius boundary layer thickness over the surface of t

rt

bent tissue paper were taped onto the

m
Aerosols and Burning Surfaces 451

During a typical experiment, the flame spread images were obtained 8 s apart. The location of the flame

was assigned to be the pixel of maximum intensity. The difference in location of the flame in a particular

row of pixels between two frames (that were separated by a known time interval) yielded the flame spread

rate for that row. To reduce the discretization errors, the flame spread was calculated from images

separated by 160 s. The mean velocity of the flame calculated for all the rows was designated to be the

velocity of the flame at that time. 

perature due to droplet loading was obtained using a two-color planar

Figure 4-82. Schematic of the 

Two-color Planar Pyrometer. 

The reduction in flame tem

pyrometer (TCPP). A schematic diagram of the two color planar pyrometer is shown in Figure 4-82. The 

TCPP had an imaging lens that was used to collect broad band radiation from the fire. The radiation was 

split into two beams that were spectrally differentiated using band pass filters. An array imaging camera

was used to obtain the spectral image of the flame at two wavelengths. The band pass filters were 

centered at approximately 900 nm and 1000 nm, and each had a band width of 80 nm. The TCPP was 

calibrated using a standard laboratory black body source. 

Results and Discussion 

The first issue addressed was the measurement uncertainty. If x is the position of the flame and t is the 

time interval between measurements, the velocity is given by v = x/t. The uncertainties in the

measurement of flame position (wx) and time (wt) were used to estimate the uncertainty in the flame 

spread rate (wV) 188

2

1

2 2

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡ ⎟

⎠

⎞ ⎜

⎝

⎛

∂

∂ ⎟ + ⎠

⎞ ⎜

⎝

⎛

∂

∂

V = V wx x

v

w

T

v

w (4-43)

For the imaging s stem, the uncertainty in position was ± 1 pixel, and that in time (provided by the 

camera timer) was 1 μs. Therefore, for a typical flame event lasting 160 s, with the flame m

y

oving about

12 pixels, the uncertainty in velocity was: 

= = 8.3%

X

w

V

wV X (4-44)

Typical flame velocity data obtained using the imaging system is shown in Figure 4-83. The imaging

system started recording at a rate of 1 image every 8 s immediately following flame ignition. During the 

first few minutes, the flame spread rate rapidly decreased until a steady state was reached. At steady 
452 Aerosol Properties 

state, the flame spread velocity changed approximately 10 % due to the small-scale fluctuations in the 

flame front and the uncertainty of the measurement. 

Figure 4-83. Typical flame 

(approximately 60 points per run). The experimental uncerta

pread Rates

without Water Mist. 

ceramic paper (on which the PMMA is burned) was not a perfect insulator.

spread time trace obtained 

using Focal Plane Array

Imaging. 

The flame spread rate as a function of opposing wind speed is shown in Figure 4-84. The vertical bars 

on of the data on each side of the mean 

inty is less than 10 % of the mean. 

represent one standard deviation in the statistical variati

Figure 4-84. Flame S

As expected, the flame spread rate without backing was higher than that with backing for the same sample

thickness, since burning was allowed to take place on both sides in the 'without backing' mode. Also in

'with backing' mode, some amount of heat loss occurred to the supporting aluminum plate since the 

0.06

0 100 200 300 400 500 600 700

Time [s]

0.07

0.08

0.09

0.1

Mean fla

me s

0.11

read 0.12

1

0.14

p rate min 0. 3 [cm/ 0.15

0.16

]
Aerosols and Burning Surfaces 453

Another observation is that the flame spread rate decreased as the opposing wind speed increased. This 

can be explained by realizing that the forward heat transfer by gas-phase conduction from the flame to the 

ned fuel surface also increased. This 

sed flow velocity. At flow

velocities of 85 cm/s and 120 cm/s, the results in the present work were lower than the findings of 

e results at the two higher flows were within the

 the experimental uncertainty and can be explained 

stand-off height was more than at the higher 

s ter than that

in Reference 48, there was greater radiation heating of the un Hence the 

measured flame spread rate was greater.

At the higher opposing wind speeds the flame stand-off height was much less. At these high wind speeds, 

radiation from the flame was much lower, due to a small vie factor. It is believed that it was the size of

the experimental facility that was responsible for differences in the measured flame spread velocities 

between this study and the work of Reference 48. In particular, the wind tunnel duct used by Reference

48 was 5 cm x 4 cm in cross-section and 100 cm long, while the duct used in the present work had a

cross-section of 45 cm x 76 cm and was 175 cm long. The greater re-radiation from the walls of the duct

ed over previously ruled

n in

unburned fuel decreased. Convective heat transfer from the unbur

result is therefore expected. It was also found that the flame spread rate was higher for the thinner fuel 

than the thicker fuel. Since the heat transfer required to raise the temperature of the unburned fuel to its 

vaporization temperature is less for a thin fuel, this finding was also expected. These three results were

consistent with the observations reported by previous investigators. 

The flame spread rate results found in this work were approximately 20 % greater than the findings of 

Fernandez-Pello et al.48 for the 0.64 cmm thickness PMMA at 70 cm/s oppo

Reference 48 by about 9 % and 6 %, respectively. Th

limits of experimental uncertainty. 

Differences at the lower flow velocity were outside of

as follows. At the lower opposed wind speeds, the flame 

opposed wind speeds. As the width of the sample in the pre ent work (30 cm) was much grea

burned fuel surface by the flame.

w

in the work of Reference 48 was believed to be the reason for the difference in the findings. 

Another reason for the differences between the current measurements and those of Reference 48 could be 

different diagnostic and analytical techniques used to determine the flame spread rate. The progress of 

the flame in the experiments of Reference 48 was measured visually as it pass

marks on the PMMA surface, whereas in the present work a CCD camera tracked the flame motion. The

data analysis method used by Reference 48 is not known, but in the present work the estimate of flame 

spread rate is over a rather large length of time (160 s) rather than being “instantaneous.” 

The flame spread rate as a function of surface loading for the 0.32 cm thickness sample is show

Figure 4-85. Similar results for the 0.64 cm thick PMMA are shown in Figure 4-86. The systematic 

decrease in flame spread rate with increasing surface loading on the PMMA can be clearly seen in all

cases. 

The effect of surface loading was greatest in the case of low opposing wind speeds, viz. 69 cm/s and

84 cm/s. The reason may be that the flame spread rate was low to begin with at the higher opposing wind

speeds and with increasing loading there was a practical limit to the flame spread rate that can exist and

still be measurable. It was conjectured that a threshold value of flame spread rate exists, between 

0.15 cm/min and 0.20 cm/min, below which the flame would be rapidly extinguished. 
454 Aerosol Properties 

The role of surface loading on the flame spread rate was explained in terms of heat transfer effects. On 

account of surface cooling, there was a reduction in pyrolysis of the fuel ahead of the flame front. As the 

rate of fuel vapor production decreased, less heat was liberated in the exothermic chemical reaction 

d velocity data are presented in Figure 4-87. 

Figure 4-85. Flame 

Thick PMMA. 

between the fuel vapor and oxygen. This resulted in a net decrease in heat transfer to the unburned 

PMMA ahead of the flame. As a result, the flame spread rate dropped. 

In order to determine the effect of spray operating parameters on the droplet size, namely the atomizer 

inlet pressure and flow, measurements of droplet size and velocity were performed using a Phase Doppler 

Particle Analyzer. Droplet size an

Spread Rate as a 

Function of Surface 

Loading for 3.2 mm 
Aerosols and Burning Surfaces 455

From Figure 4-87, it is seen that both the droplet size and the mean velocity increased as the mass flow

rate through the nozzle increased for a constant inlet pressure. At the same flow through the nozzle, the 

droplet size was smaller and the mean velocity greater at the higher pressure. The difference in the mean

droplet diameter at 10 cc/min flow at the three different pressures was approximately 2.5 μm; at

15 cc/min flow rate it was approximately 1 μm. By the d of droplet evaporation the ratio of 2

-law

evaporation times of droplets 27.5 μm and 25 μm in diameter is 1.2. Therefore, a notable impact due to 

evaporative cooling in the flame vicinity cannot be expected. Consequently, the droplet size variation 

should not be a significant factor in flame spread rate variations for the conditions considered here. 

Figure 4-86. Flame 

Spread Rate as a 

Function of 

Surface Loading 

for 0.64 cm Thick 

PMMA. 
456 Aerosol Properties 

The droplet mean velocity measurements provided an explanation as to why the flame spread rate 

increased with increasing nozzle discharge rate at low opposing wind speeds. The reason is that the 

different injection pressures 

Velocity as a 

droplet mean velocity increased as the flow rate increased. This reduced drift due to the cross-wind and 

led to deposition of the spray in the collection pan (placed below the spray nozzle) upstream of the target. 

The surface loading on the PMMA was thus reduced and, consequently, the flame spread rate increased. 

The data obtained from all tests were used to develop an empirical correlation. The normalized flame 

spread rates for all conditions are shown in Figure 4-88. Three 

(corresponding to different droplet diameters), two different PMMA thicknesses, and four different 

opposed flow velocities were used in developing the empirical correlation. 

Data represented by open symbols shown in Figure 4–87 were obtained with 0.32 cm thick PMMA, and 

those represented by closed symbols were obtained with 0.64 cm thick PMMA sheets. 

Figure 4-87. Mean 

Droplet Size and 

Function of Spray 

Nozzle Discharge 

Rate. 
Aerosols and Burning Surfaces 457

Liquid loading (g/min/m2)

0 20 40 60 80

Nor

maliz

p

0.00

0.25

0.50

0.75

1.00

0.60 m/s

0.94 m/s

1.21 m/s

1.50 m/s

ed Fla

me 

S read

100

The regression analysis for the normalized flame spread in the thinner samples provided an approximately

85 % correlation with liquid mass loading. For the thicker samples, the correlation of the normalized 

flame spread with the liquid mass loading was approximately 75 %. The correlation obtained for all 

twenty-four conditions showed that the normalized flame spread rate was roughly a function of the local 

liquid mass loading alone. Correlations for flame spread rates with zero liquid mass loadings are readily 

available in the literature. 

Normalized 

Flame Spread 

with Liquid 

Loading. 

Therefore, the current results can be used to estimate (within a factor of two) what the local liquid mass 

flux should be to achieve extinction. This information is useful, since it provides a direct estimate of the

optimal suppressant mass required to extinguish the fires. The effect of water in

the mean flame temperature is shown in Figure 4-89. The mean temperature t any

water loading varied between 1200 K and 1400 K depending on the oppo e

temperature dropped rapidly with increasing water loading, and the flame was extinguished at

approximately 1000 K. 

Figure 4-89. Mean Flame 

Temperature Obtained Using 

the TCPP. (880 kPa, 0.64 cm 

thick PMMA) 

Figure 4-88. 

Variation in 

loading on the decrease

of the flame withou

sing wind velocity. Th
458 Aerosol Properties 

nd the

duced into a thick-

 the reservoir and passed 

mizer. The filter had 

let pressure ranged from 

 Distilled water was used

was used for the buoyancy 

) was used to produce 

monodisperse droplets with flows up to 4 g/s.189 A schematic of the atomizer is sh e 4-91. A

2 MHz function generator (Wavetek, Model 19) was used with an amplifier to sup

(approximately 30 V peak to peak) to the atomizer. The output frequencies ranged .

Three round stainless steel plates with orifices of different diameters made by laser drilling were used as 

nozzles in the atomizer. The plates were 50 μm thick, and were supported by a 1 m e

with a 1 mm hole in the center. The orifice diameters were 100 μm, 50 μm, and The atomizer

was placed above the center of the heater in a fixture that allowed 360 o rotation. In addition, the atomizer 

was capable of being moved in both the vertical and horizontal direction. 

Figure 4-90. Experimental 

Arrangement for the 

Measurements. 

4.7.2 Droplet Interaction with a Hot Surface 

Experimental 

Schematics of the experimental arrangement are shown in Figure 4-90 through Figure 4-93. The

experimental apparatus consisted of six main parts: the liquid supply system, the piezoelectric ceramic 

atomizer and its power supply system, the heater assembly and its power supply system, the temperature 

data acquisition system, the Phase Doppler Particle Sizer (PDPA) spray characterization system, a

spray imaging system. 

Compressed shop air was reduced to an appropriate pressure by a regulator then intro

walled plastic reservoir in which the liquid was stored. Liquid flowed out of

through a filter and a rotameter, before finally reaching the piezoelectric ceramic ato

a pore size of 10 μm. The mass flow ranged from 0.1 g/s to l.0 g/s, and the in

27.6 kPa to 82.7 kPa. They were adjusted simultaneously using the rotameter.

for evaporation experiments. Dow Corning 510, a non-evaporative fluid, 

experiments. 

A piezoelectric atomizer (purchased from Universitat Erlangen-Nurnberg, Germany

own in Figur

ply square wave signals 

from 3 kHz to 2 MHz

m thick stainless plat

20 μm.
Aerosols and Burning Surfaces 459

Figure 4-91. 

Configuration of the 

A heat source was necessary to provide a surface temperature as uniform as possible. A modified design 

of a hot plate was used in the experiments.190 A schematic of the heater assembly is shown in 

Figure 4-92. The main components were an exterior aluminum ring with top and bottom cover plates, a 

copper block with inside cartridge heaters, and insulation material. To reduce oxidization of the heated 

surface and clogging in the orifices and pipes, distilled water was used. The plate used was taken out of 

the atomizer after each experiment and cleaned using an ultrasonic cleaner before the next test. 

The arrangement of the oxygen-free copper block and the thermocouples is show

block was 10 cm in diameter. The exposed surface was nickel-plated to a thickn -

three cartridge heaters were obtained from OMEGA, each rated at 100 W maximu ed 

to heat the block as uniformly as possible. Twelve type K subminiature thermoco

the copper block at 4 radial locations. There were 3 in each group, located at different depths to 

extrapolate the surface temperature. Holes for the cartridge heaters were 0.94 cm eter, and a total 

of 23 were arranged in three concentric circles as shown in Figure 4-93. The holes for the thermocouples 

were 0.16 cm in diameter, and a total of 12 were located as shown in Figure 4-93. 

The copper block was surrounded by Cotronics Rescor 360 Ceramic Board insulation. This insulation 

was selected for low thermal conductivity (0.055 W/m-K at 260 oC, 0.108 W/m-K at 600 oC), resistance 

to high temperature (continuous operating temperature limit of 1260 oC), rigidity, and machinability. 

There was a center hole in each insulation layer below the copper block to allow the thermocouples and 

cartridge heater leads to pass through. 

A square, 1.59 mm thick Macor machinable ceramic cover was placed above the uppermost piece of the 

insulation layer, with a hole 10.16 cm in diameter in it for the exposed nickel-plated copper surface. The 

Macor sat on a ridge machined into the top aluminum cover plate. The exposed heater surface was 

maintained flush with the surrounding Macor plate by adjusting three screws beneath the insulation at the 

very bottom of the assembly. High temperature sealant 906 (Megnesia Adhesive from Cotronics) was 

used to fill the gap between the Macor and the aluminum cover plate. Excess water flowed down along 

the exterior surface of the aluminum ring and was drained out through a hole in the slot, which was 

formed by the cylinder and the short ring mounted on the bottom plate. 

Piezoelectric Atomizer. 

n in Figure 4-93. The 

ess of 50 μm. Twenty

m. They were arrang

uples were inserted into

 in diam
460 Aerosol Properties 

Figure 4-92. 

Schematic 

Diagram of the 

Heater Assembly. 

Figure 4-93. Temperature 

Measurement Locations 

on the Heater. 

Insulation 

Copper 

Macor Ceramic

Aluminum 

Wire leads from heaters and 

thermocouples
Aerosols and Burning Surfaces 461

A PDPA was used to measure the sizes and velocities of the droplets (See Section 4.3.4). 

transmitter (Model 11 00) contained a 10 mW, polarized Helium-Neon laser. The colli

 The PDPA 

mating lens was 

receiver lens was 1000 mm to permit the largest range for measuring droplet sizes. The 

mode was used to obtain data. The position of the Phase Doppler Particle Analyzer 

from the thermocouples were received by the EXP-16 board (Computer Boards)

and multiplexed into a single signal for amplification. A gain of 100 was used to achieve the largest 

e operating frequencies were changed to get different droplet sizes and velocities. A 

frequency was chosen by visually observing the jet shape and ensuring that there were no satellite 

on Lund et al.191 that accounts

for coalescence and evaporation of droplets was evaluated using the data. 

The influence of evaporation was evaluated through use of the d2

 law as described by 192

300 mm, and the

forward scattering 

(PDPA) was fixed. The heater and the atomizer were moved together to measure droplets at different

axial and radial locations. A three-dimensional motor controller was used to provide accurate relative

positioning of the laser probe volume. 

The 12 voltage signals

voltage measurable and programmable by the CIO-DAS1402/12 board (Computer Boards). The EXP-I6

also contained a CJC (cold junction compensation) circuit for correcting thermocouple readings. 

Experiments were performed for different liquids, nozzle plates, flows, and operating frequencies. The 

initial surface temperature of the copper was chosen to be 600 oC for water and 250 oC for Fluid 510. In 

addition, a base line data set with the heater at 25 oC was also obtained. For a given a nozzle plate and a 

given flow, th

droplets breaking away from the main stream. A plastic curtain was used in the experiments to reduce the

ambient disturbances. Measurements at different axial and radial locations within the spray were

obtained by moving the heater and the nozzle assembly using a slide. 

Modeling 

For designing suppression systems, models are routinely used to calculate droplet trajectories. Since 

droplets arriving at a fire location have to pass through a hot environment, the models also have to 

account for buoyancy and evaporation. The data obtained using the PDPA can be used for assessing the 

accuracy of these models. During the course of the project, a model based

ShD

RT

p D

dt

dm v = 2π V (4-45)

where m is droplet mass, t is time, DV is the mass diffusion coefficient, p

ideal gas constant for the vapor, T is the temperature, Sh is the Sherwoo

diameter. 

A change of variables converted the change in mass with respect to time in Equation (4-45) to a change in

diameter with respect to axial location. The resulting equation was: 

v is the vapor pressure, R is the

d number, and D is the droplet

DRTV

D p Sh

dz

dD

L

V v

ρ

4 = (4-46)

The remaining unknown variable, the Sherwood number, was calculated using the correlation found in

Ranz and Marshall:193
462 Aerosol Properties 

(4-47)

where Re is the Reynolds number and Sc is the Schmidt number. 

Calculations were made using Equations (4-46) and (4-47), plus the definitions of Re and Sc, to determine 

the change in droplet diameter with axial location for a particular size class. Since droplet diameter was 

found to decrease by only 0.015 μm for every 1 cm of axial distance, it wa ncluded that the influence 

of evaporation was negligible for the hydrocarbon oil used in this investigation.

The extent to which droplet coalescence played a role in increasing SMD etic

theory-based model. The model neglected viscous, gravitational and pressu e

coalescence of a droplet of size-velocity class i with a droplet of size-velocity class j to yield a droplet of

size-velocity class k: 

0.5 0.33 Sh = 2 + 0.6Re Sc

s co

was determined via a kin

re forces. It was based on th

Di + Dj → Dk (4-48)

From this relationship the rate equation for change in number of droplets of class i was: 

p

i

ij i j ij

i

ij i j i V

V

k

V

k

dz

dn

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ = −∑ ∑+

η η η η α

ε (4-49)

where ni, the counts per unit volume for droplet class i, was defined as 

p run i

i

i A t V

n

η = (4-50)

The collision frequency for droplets of class i colliding with those of class j, was defined as: 

( ) ij Di Dj Vij k 2

4 = +

π (4-51)

The probability of a collision leading to coalescence, ε, was determined by extrapolating the work of

Jiang et al.194 to the fluids studied here, and then approximating the impact parameter vs. Weber number

curves as straight lines. Such a procedure overestimated the probability of a collision leading to

coalescence. 

In Equations (4-49) through (4-51), ni represents the number of droplets in class i, Ap is the PDPA probe 

area normal to the flow, trun is the run time, Vi is the velocity of class i, Vp is the PDPA probe volume, and 

aij represents the diameter Dk that results from coalescence in Equation (4-48). 

Equations (4-49) through (4-51) were used to calculate the change in number of droplets for a particular 

class i with respect to axial location. Care was taken when distributing the product mass and momentum

over the appropriate size-velocity bins. The coarse-grained PDF technique of Sivathanu and Gore186

 was

employed to keep computational evaluation of Equation (4-49) tractable while ensuring spray mass and

momentum conservation. Having calculated dni/dz in this fashion, dSMD/dz was determined by taking 

the derivative of SMD with respect to z. 
Aerosols and Burning Surfaces 463

( )

( )2 2

2

∑

∑ ⎟

⎠ ⎜⎝

⎛

= −

i i

i i i

n D

dz

dn D

dz

dSMD (4-52)

Equation (4-52) was then used to calcula SMD due to coalescence of droplets with axial 

location. The results were compared with those obtained using the PDPA instrument. 

The temperature field above the hot plate is shown in Figure 4-94. The tempe just

above the hot plate and rapidly decreased due to b y induced entrainment. f

100 mm, the temperature was less than 200 oC ter line. The buoyancy induced velocity field 

was unsteady. (The time-resolved velocity field was not quantified.) The injector was located 250 mm 

above the hot plate in the centerline. 

Figure 4-94. Steady State 

Temperature Field above the Hot 

to temperature 

field. However, even under steady state conditions, the local velocity field had both positive and negative 

values. The model did not have a mechanism to account for flow reversals. Therefore, the model 

predictions showed very large divergence from the experimental observations and are not provided here. 

A 3-D large eddy simulation code is needed to improve the model predictions. 

The change in droplet diameter with distance from the injector is shown in Figure 4-95. The PDPA 

results obtained using a cold plate and a hot plate are also shown. The effects of coalescence are readily 

apparent. For both cases, the mean diameter of the droplets increased by approximately a factor of three

by the time they arrived at the hot plate. For the hot surface, the droplet diameter increased only by a

factor of two. The smaller droplet diameters in the hot environment are likely due to evaporation. They

may also be due to decreased coalescence. If the buoyancy induced flow field affects the local velocities 

3

∑i ⎞

n D

te a change in the 

Results and Discussion 

rature was 600 oC

uoyanc At an axial distance o

at the cen

Plate. 

A steady state k-ε-g model, assuming a boundary layer flow, was used try to predict the
464 Aerosol Properties 

Distance from injector (mm)

50 100 150 200 250

SM

D / S

M

D

O

0

1

2

3

4

298 K 873 K

0.014 g/s

0.018 g/s

0.028 g/s

Water, 50 micron injector

Distance from injector (mm)

0 50 100 150 200 250

SM

D

D/S

M

O

0

1

2

3

298 K 523 K

Fluid 510, 50 micron injector, 0.014 g/s

8.29 KHz

12.9 KHz

significantly, this could lead to lower coalescence and reduce the SMD. These two effects can be

separated using non-evaporating droplets. 

size history. When the fluid was

sprayed onto the surface at 600 oC, a thick smoke was formed. This smoke prevented the PDPA from 

oplet velocity and size data. In addition, the fluid formed a very hard coating on the copper 

fore, the plate was cleaned, and the temperature reduced to 250 oC (below the flash point of

did not significantly affect the 

en the plate was

n Diameter of the Droplets with 

Axial Distance for Fluid 510. 

Figure 4-95. Change in Sauter 

Mean Diameter of the Water 

Droplets with Axial Distance. 

Fluid 510 was used to determine evaporation effects on the droplet

obtaining dr

plate. There

Fluid 510). 

The change in droplet diameters with axial locations for Fluid 510 is shown in Figure 4-96. Two different 

excitation frequencies, representing two different droplet diameter distributions were used for the

experiments. For both cases, it was found that the variation in SMD was identical, irrespective of the 

temperature of the plate. Therefore, the buoyancy induced flow field

coalescence of droplets. It was evaporation that was responsible for the lower SMD wh

hot. 

Figure 4-96. Change in Sauter 

Mea
Aerosols and Burning Surfaces 465

T

Distance from injector (mm)

50 100 150 200 250

V

elo

city (

m/s)

0

5

10

15

Fluid 510

298 K 523 K

0.020 g/s

0.014 g/s (50 μm, 8.29 KHz)

0.014 g/s (50 μm, 12.9 KHz)

(100 μm, 7.1 KHz) Figure 4-98. V

o evaluate the coalescence model, the velocity and number density of the droplets was required. The 

rly evident that the velocity of droplets 

uch lower when the plate was hot. This 

Figure 4-97. Variation in Velocity

with Axial Distance for Water 

Spray.

Figure 4-98.

Due to the lower temperatures attained, the change in velocity due to the buoyancy induced upward

elocity of air was minimal. Similar results were obtained for the 20 μm nozzle. In all cases, the droplet 

diameter increased a factor of two or higher with downstream distance. Coalescence was the p

ariation in Velocity

with Axial Distance for Fluid 510. 

velocity of the water droplets is shown in Figure 4-97. It is clea

was affected significantly by buoyancy. The velocities are m

usually increased the droplet SMD since the lower velocities provided a longer time interval for

coalescence. 

The change in velocity with axial distance for three conditions using Fluid 510 is shown in

Distance from injector (mm)

50 100 150 200 250

V

elo

city (

m/s)

0

5

10

15

298 K 873 K

0.014 g/s

0.018 g/s

0.028 g/s

Water, 50 micron injector

v

rimary

reason for the increase in droplet diameter. 
466 Aerosol Properties 

odeling results of the change in SMD with diameter using Equation (4-46) for coalescence and 

Equation (4-52) for droplet evaporation differed significantly from the experimental data

coalescence model predicted approximately two orders of magnitude lower changes in SMD wi

distance. The evaporation model suggested that the SMD of droplets would change by less then 2 % in all 

Improvements in the model using a Monte Carlo-based approach are needed to delineate the effects of 

Also needed is a more rigorous method of predicting the temperature field 

ement the Monte Carlo simulations. Finally, wake effects are important in 

f water mist on the reduction in the flame spread velocity into an 

opposing wind. Flame spread velocities with and without water mist were obtained for a solid fuel, 

polymethylmethacrylate (PMMA), using Focal Plane Array Imaging. PMMA sheets of different

thickness were burned in an opposed flow configuration, with varying wind velocities. The flame spread 

rate for the 'without backing' tests was found to be greater than for the 'with backing' tests due to flame 

spread on both sides in the ‘without backing’ case. At ambient oxygen concentration, the heat transfer 

effect on the fuel vapor evolution rate was the dominant factor controlling the flame spread rate. 

A spray nozzle, suspended over the PMMA sheets, was used to vary the water loading at the flame front. 

Flame spread in the presence of a water mist indicated a strong educing

flame spread rate. The effect was more pronounced at low valu

spread rate for a given liquid loading decreased slightly for smal

the droplet size, is the more important parameter for reducing flame The reduction in flame 

spread rate in the presence of the mist can be modeled empirically sing the experimental parameters. 

Droplet diameter and velocity measurements from a monodisperse spray onto a hot plate showed that 

d diameters. A non-evaporating spray 

ncy and an evaporating spray (water) 

was used to study the combined effects of buoyancy and evaporation on the droplet trajectories. For both

fluids, coalescence was the dominant mechanism by which drop eir flight to a 

hot surface. For the temperature range achieved during the study, the change in water droplet diameters

M

. The

th axial

cases. 

coalescence and evaporation.

above the hot surface to impl

this study and should be incorporated into the models. 

4.7.3 Summary and Conclusions 

This research examined the effect o

influence of surface loading in r

es of opposing wind speeds. The flame

ler droplets. Water loading, rather than

spread. 

u

buoyancy had a significant impact on both droplet velocities an

(Dow Thermal Fluid 510) was used to assess the effects of buoya

lets changed size during th

due to evaporation was much smaller than due to coalescence. Two coalescence and evaporation models

developed were not successful in predicting droplet coalescence and evaporation. The models need to be

improved significantly before they can be used in fire suppression studies. Temperature reduction of 

burning surfaces inhibited with water mist has to be measured for more conditions to obtain a more 

accurate assessment of the optimal water loading required for fire extinguishments. 

The results of the study are applicable to design considerations for aerosol-based fire suppression devices

to ensure optimum operation. 
The New State of Knowledge of the Mechanism of Flame Extinguishment by Aerosols 467

4.8 THE NEW STATE OF KNOWLEDGE OF THE MECHANISM OF FLAME 

EXTINGUISHMENT BY AEROSOLS 

NGP research has greatly advanced the understanding of the processes by which aerosols in the form of 

solid particles or liquid droplets suppress flames, the aerosol properties that control these processes, and 

the values of the properties that promote efficient extinguish ent. The interrelated dependencies of 

particle size and flame conditions explain why seemingly contradictory suppression results appeared in 

the earlier literature exploring the effectiveness of aerosols in differing flame systems. Experimental 

quantification of the aerosol particle size distributions, number densities, and particle velocities actually 

delivered to the flames was key to establishing valid corr ion 

effectivenesses with modeling predictions. Comparison of n with 

modeling predictions permitted the exploration of suppression in have 

been much more difficult to explore solely with either approach alone. 

• Rosser and co-workers summarized the extinction mechanism of powders of metal 

compounds based on experiments in the 1960s.13 GP assessment of subsequent works 

combined with detailed NGP experimental and modeling studies concluded that with some 

modification these conclusions can be extended to qualify the suppression mechanism of 

aerosols in general. 

• To achieve maximum suppression effectiveness, individual components of an aerosol must 

pass through a succession of steps: I - Heating; II - Evaporation/decomposition; II - Generation 

• Any generated chemical inhibitor species must directly participate in the flame radical 

• Suppression effectiveness is controlled by the time required for the aerosol particles to 

progress through each step AND the amount of tim allowed, i.e. the residence time. 

• For premixed flames, higher flame temperatures imply higher burning velocities and shorter 

particle residence times. 

• For non-premixed counterflow flames, the particle esidence time is related to the inverse of 

the strain rate; particles can oscillate near the stagnation plane and lengthen the effective 

residence time for certain particle sizes and strain r

• For non-premixed co-flow flames with air-stream 

flames), the residence times are dictated by the air 

• Residence time requirements are greater for liquid aerosols with a chemical inhibitor additive 

than for the pure liquid; liquid with additive aerosols must complete steps I and II for the 

liquid followed by steps I and II again for the additive residue. 

m

elations between observed suppress

experimental data in combinatio

 aerosol parameter space that would

 N

of the chemical inhibitor specie(s) for chemical agents; IV - Flame inhibition. 

• Evaporation/decomposition leads to dilution of the available oxygen and lowers the 

temperature which alters the flame chemistry (dominates for those aerosols with no chemical 

inhibitor). 

chemistry. 

• Small particles are favored; evaporation/decomposition must be close to or in the flame. 

• The transition diameter from effective, small particles to larger, less effective larger ones 

depends on the time to complete steps I and II (i.e., complete evaporation of the aerosol 

material). 

e 

 r

ates. 

entrainment of the aerosol (e.g. cup burner 

stream velocity through the flame. 
468 Aerosol Properties 

me fuel.

eness there is a tradeoff between flame 

dence time, and particle physical and chemical 

c effectiveness behavior with particle size. 

 properties ultimately determine the magnitude 

IR

The conclusions in this chapter are based for the most part on tions in laboratory-scale flame 

suppression studies. Aerosol particles were typically injected directly into the flame. For real-scale fires, 

the conclusions outlined above will include a critical additional first step; the aerosol must first get to the 

fire to begin the process. The type of aerosol generation system ens in aerosol transport to

the fire will also impact suppression performance. Thus, in addition to answering what aerosol properties

are optimum at the fire, one must also address what properties should the aerosol exhibit when generated 

in order to achieve the optimum suppression properties at the fire. The required set of aerosol 

characteristics can change as the fire changes, especially as the ire grows in size involving more of its 

surroundings. Transport and delivery topics are presented and discussed in Chapter 8. 

a need to consider the effective dispersion throughout the a

densities need to be generated in order to fill the space ade his often impacts the resulting 

particle size distributions. 

Small fires in large spaces, especially small, obstructed fires, are particularly challenging for gaseous 

agents. These fires are even more challenging for aerosol suppression agents. Smaller fires generate less 

heat, resulting in lower overall tempeatures and lower aerosol evaporation or decomposition rates. 

Entrainment currents around smaller fires are less intense and less likely to draw large amounts of aerosol 

into the fire. For a given aerosol, the smaller particles are more likely to follow the weaker airstream 

trajectories than the larger particles, but smaller particles eva ol

distribution (or other combination) may address this issue. A he

aerosol requirements for protecting a wider range of fire threats and protecting against fires in varying

stages of development. 

• For chemical inhibitor species, thermodynamic limitations can impose restrictions on the gas 

phase chemical inhibitor specie(s) (see Chapter 3); complete evaporation/decomposition 

and/or longer residence times may offer no advantage. 

• Aerosols are generally less effective in premixed flames than in non-premixed flames of the

sa

• For achievable aerosol suppression effectiv

temperature, flow field velocity, particle resi

properties; this tradeoff can lead to non-monotoni

• Specific flame type, nature of fuel, and aerosol

of the aerosol suppression effectiveness. 

4.9 RELATION OF FINDINGS TO PRACTICAL F ES

observa

 and what happ

f

Spaces where real fires occur are generally distinctly larger than the dimensions of a laboratory flame 

apparatus. Aerosol travel distances are longer; changes will occur in the size distribution and particle 

velocities from generation point to the fire location. This location is not known in advance so there is also

rea requiring protection. Higher aerosol 

quately. T

porate more quickly. A bi-modal aeros

bi-modal distribution may also address t
References 469

4.10 REFERENCES 

1. Braidech, M.M., Neale, J.A., Matson, A.F., and Dufour, R.E., The Mechanism of Extinguishment of 

Fire by

Underwriters,

 Finely Divided Water, Underwriters Laboratories lnc. for the National Board of Fire 

 N.Y., 1955.

MD, 1980. 

B., “Fire Protection: Water Curtains,” Fire Safety Journal 20, 

ent Properties of Dry 

17. Fischer, G., and Leonard, J.T., Effectiveness of Fire Extinguishing Powders Based on Small Scale 

ire Suppression Tests, Naval Research Laboratory, NRL/MR/6180-95-7778, Washington, D.C., 1995. 

18. Chattaway, A., Dunster, R.G., Gall, R., and Spring, D.J., “The Evaluation of Non-Pyrotechnically 

Generated Aerosols as Fire Suppressants,” 1995, in Gann, R.G., Burgess, S.R., Whisner, K.C., and 

2. Rasbash, D.J., Rogowski, Z.W., and Stark, G.W.V., “Mechanisms of Extinction of Liquid Fires with 

Water Sprays,” Combustion and Flame 4, 223–234, (1960). 

3. Lugar, J.R., “Preliminary Test Results of Fine Water Mist Fire Protection Systems Study,” David W. 

Taylor Naval Ship Research and Development Center, 

4. Coppalle, A,, Nedelka, D., and Bauer, 

241-255, (1993). 

5. Mawhinney, J.R., Dlugogorski, BZ., and Kim, A.K., “A Closer Look at the Fire Extinguishing 

Properties of Water Mist,” Proceedings of the Fourth International Symposium on Fire Safety Science, 

The International Association for Fire Safety Science, pp. 47–60, 1994. 

6. Seshadri, K., “Structure and Extinction of Laminar Diffusion Flames Above Condensed Fuels with 

Water and Nitrogen,” Combustion and Flame 33, 197-215, (1978). 

7. McCaffrey, B.J., “Momentum Diffusion Flame Characteristics and the Effects of Water Spray,” 

Combustion Science and Technology 63, 315-335, (1989). 

8. Fielding, G.H., Williams, F.W., and Carhart, H.W., Fire Suppression – Why Not Water?, NRL 

Memorandum Report No. 3435, Naval Research Laboratory, Washington, DC, February 1977. 

9. Friedman, R., “On the Mechanisms of Flame Extinguishment by Dry Powders,” Journal of Fire 

Protection Engineering 5, 290-318, (1993). 

10. Dolan, J.E., and Dempster, P.B., “The Suppression of Methane-Air Ignitions by Fine Powders,” 

Journal of Applied Chemistry 5, 510-517, (1955). 

11. DeWitte, M., Verbosch, J., and van Tiggelen, A., “Inhibition and Extinction of Premixed Flames by 

Dust Particles,” Combustion and Flame 8, 257-266, (1964). 

12. Birchall, J.D., “On the Mechanism of Flame Inhibition by Alkali Metal Salts,” Combustion and Flame

14, 85-96, (1970). 

13. Rosser, W.A., Jr., Inami, S.H., and Wise, H., “The Effect of Metal Salts on Premixed HydrocarbonAir Flames,” Combustion and Flame 7, 107-119, (1963). 

14. G. Fischer and J. T. Leonard, The Effectiveness of Fire Extinguishing Powders Based on Small Scale 

Fire Suppressant Tests, Internal NRL Progress Report, 1984. 

15. Ewing, C.T., Hughes, J.T., and Carhart, H.W., “The Extinction of Hydrocarbon Flames Based on the 

Heat-Absorption Processes which Occur in them,” Fire and Materials 8, 148-156, (1984). 

16. Ewing, C.T., Faith, F.R., Hughes, J.T., and Carhart, H.W., “Flame Extinguishm

Chemicals: Extinction Concentrations for Small Diffusion Pan Fires,” Fire Technology 25, 134-149, 

(1989). 

F
470 Aerosol Properties 

Reneke, P.A., eds., Papers from 1991-2006 Halon Options Technical Working Conferences (HO

IST SP 984-4, National Institute of Standards

TWC), 

CD-ROM, N and Technology, Gaithersburg, MD, 2006. 

sociation, Quincy, MA, 1997. 

23. Monson

Additives fo

Laboratorie

24. Monson

Additives for Fire Fighting Report No. 4, U.S. Army Engineer Research and Development 

Laborat e

25. Grove, 

Abstracts and Reviews

26. Aidu , 

the Fire-Fig

27. Wa r

Characte t

28. Wa r ting 

Charact t

29. Aid A

Fire Resear

30. Aidun, A.R. “Additives to Im

Abstrac n

31. Aidun, A.R.,”Additives to Im

Abstrac n

32. Grove, C.S., Jr., Grove, idun, A.R., “Improving the Effectiveness of Water for Fire 

Fighting,” F

33. Aidun, A

Fire Research Abstracts and Reviews

34. Dav , 

Suppression Crew,” 

35. Dav 

Control Ma

19. Knuth, E.L., Ni, W.F., and Seeger, C., “Molecular-Beam Sampling Study of Extinguishment of 

Methane-Air Flames by Dry Chemicals,” Combustion Science and Technology 28, 247-262, (1982). 

20. Hamins, A., Gmurczyk, G., Grosshandler, W., Rehwoldt, R.G., Vazquez, I., Cleary, T., Presser, C., 

and Seshadri, K., “Flame Suppression Effectiveness,” in Evaluation of Alternative In-Flight Fire 

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

Institute of Standards and Testing, No. NIST SP 861, pp. 345-465, April 1994. 

21. Hamins, A., “Flame Extinction by Sodium Bicarbonate Powder in a Cup Burner,” Proceedings of the 

Combustion Institute 27, 2857-2864, (1998). 

22. Wahl, A.M., “Water and Water Additives for Fire Fighting,” Section 6, Chapter 1, in Fire Protection 

Handbook, 18th edition, A.E. Cote (editor), National Fire Protection As

, W.L., Reid, L.J., Schreiner, M., Jr., Bonney, D.T., and Huff, W.J., Research on Water 

r Fire Fighting, Quarterly Report No. 3, U.S. Army Engineer Research and Development 

s, Fort Belvoir, Virginia, December 15, 1953. 

, W.L., Reid, L.J., Schreiner, M., Jr., Bonney, D.T., and Huff, W.J., Research on Water 

, Quarterly

ori s, Fort Belvoir, Virginia, March 15, 1954. 

C.S. et al., “Additives to Improve the Fire-Fighting Characteristics of Water,” Fire Research 

1, 18, (1958). 

n A.R., Walker, E.J., Gabriel, S., Buckman, W.R., and Grove, C.S., Jr., “Additives to Improve 

hting Characteristics of Water,” Fire Research Abstracts and Reviews 1, 58, (1958). 

lke , E.J., Aidun, A.R., and Grove, C.S., Jr., “Additives to Improve the Fire-Fighting 

ris ics of Water,” Fire Research Abstracts and Reviews 1, 129, (1958). 

lke , E.J., Aidun, A.R., and Grove, C.S., Jr., “Additives to Improve the Fire-Figh

eris ics of Water,” Fire Research Abstracts and Reviews 2, 71, (1960). 

un, .R. and Grove, C.S., Jr., “Additives to Improve the Fire-Fighting Characteristics of Water,” 

ch Abstracts and Reviews 3, 91, (1961). 

prove the Fire-Fighting Characteristics of Water,” Fire Research 

ts a d Reviews 3, 93, (1961). 

prove the Fire-Fighting Characteristics of Water,” Fire Research 

ts a d Reviews 3, 213, (1961). 

 S.T., and A

ire Research Abstracts and Reviews 4, 54, (1962). 

.R. and Grove, C.S., Jr., “Additives to Improve the Fire-Fighting Characteristics of Water,” 

5, 34, (1963). 

is J.B., Dibble, D.L., and Phillips, C.B., “Firefighting Chemicals–New Weapon for the Fire 

Fire Research Abstracts and Reviews 5, 37, (1963). 

is, J.B., Dibble, D.L., Phillips, C.B., and McBride, R.S., “Viscous Water and Algin as Fire 

terials,” Fire Research Abstracts and Reviews 5, 37, (1963). 
References 471

36. Friedric

Abstracts an

37. Elki G

38. Kida, H n of Fires of Liquid Fuel with Sprays of Salt Solutions,” Fire Research Abstracts 

and Reviews

39. Ball, D.

June 21, 199

40. Finnerty

1138, Army Ground, July, 1996. 

Dispersed Liquid Agent Fire 

te of Standards and Technology, 

0, 29 June-1 July, 

1999. 

logy 19 (1-2), 19-30, (1978). 

52. Kwok, K.C., A Fundamental Study of Air-Spray Painting, Ph. D. Thesis, Mechanical Engineering 

epartment, University of Minnesota, 1991. 

3. Hicks, P.G., and Senser, D.W., “Simulation of Paint Transfer in an Air Spray Process,” Journal of 

luids Engineering 117, 713-719, (1995). 

h, M., “Fire Extinguishing Experiments with Aqueous Salt Solution Sprays,” Fire Research 

d Reviews 6, 44, (1964). 

ns, .H.J., “Use of Anti-Freeze in Water Extinguishers,” Fire 57, 448, (1965). 

., “Extinctio

11, 212, (1969). 

N., Spring, D.J., and Davis, S.J., “Fire Extinguishing Substance,” UK Patent GB 2,233,226A, 

0. 

 A.E., McGill, R.L., and Slack, W.A., Water-Based Halon Replacement Sprays, ARL-TR-

 Research Laboratory, Aberdeen Proving 

41. Yang, J.C., Donnelly, M.K., Prive, N.C., and Grosshandler, W.L., 

Suppression Screen Apparatus, NISTIR 6319, National Institu

Gaithersburg, MD, July 1999. 

42. Zheng, R., Bray, K.N.C., and Rogg, B., “Effect of Sprays of Water and NaCl-Water Solution on the 

Extinction of Laminar Premixed Methane-Air Counterflow Flames,” Combustion Science and 

Technology 126, 389–401, (1997). 

43. Mitani, T., and Niioka, T., “Extinction Phenomenon of Premixed Flames with Alkali Metal 

Compounds,” Combustion and Flame 55, 13-21, (1984). 

44. Edwards, M., Watkins, S., and Glockling, J., “Low Pressure Water Mist, Fine Water Spray, Water 

Source and Additives: Evaluation for the Royal Navy,” Proceedings of the 8th International Fire Science 

and Engineering Conference, Interflam ’99, Edinburgh, Scotland, Vol. 1, pp. 639-65

45. De Ris, J.N., “Spread Of A Laminar Diffusion Flame,” Proceedings of the Combustion Institute 12, 

241-252, (1969). 

46. Lastrina F.A., Magee R.S., and McAlevy III, R.F., “Flame Spread Over Fuel Beds: Solid-Phase 

Energy Considerations,” Proceedings of the Combustion Instisute 13, 935-948, (1971). 

47. Fernandez-Pello, A.C., Ray, S.R., and Glassman, I. “Downward Flame Spread In An Opposed Forced 

Flow,” Combustion Science and Techno

48. Fernandez-Pello, A.C., Ray, S.R., and Glassman, I., “Flame Spread in an Opposed Forced Flow: The 

Effect of Ambient Oxygen Concentration,” Proceedings of the Combustion Institute 18, 579-589, (1981). 

49. Apte, V.B., Bilger, R.W., Green, A.R. and Quintiere, J.G., “Wind-Aided Turbulent Flame Spread and 

Burning Over Large-Scale Horizontal PMMA Surfaces,” Combustion and Flame 85, 169-184, (1991). 

50. Loh, H.-T. and Fernandez-Pello, A.C., “A Study Of The Controlling Mechanisms Of Flow Assisted 

Flame Spread” Proceeding of the Combustion Institute 20, 1575-1582, (1984). 

51. Perzak, F.J. and Lazzara, C.P., “Flame Spread Over Horizontal Surfaces of Polymethylmethacrylate,” 

Proceedings of the Combustion Institute 24, 1661-1668, (1992). 

D

5

F
472 Aerosol Properties 

54. DiMarzo, M.. Kavoosi, and Klassen, M., Transient Cooling of a Hot Surface by Droplets 

Evaporation, Report prepared for the National Institute of Standards and Technology, NISTGCR- 89-559, 

ce, NBSIR 86-3384, National Insitute of Standards and Technology, 

D., Black, B., Friderichs, T., Peatross, M., 

 1991-2006 Halon Options Technical 

ire Protection Engineering 4, 35-52, 

, Ouellette, R.J., Hughes, J.T., and Carhart, 

ant Testing in a 56 m (2000 ft ) Compartment,” 1993, in 

6. 

 Pyrogenic Solid Aerosol,” 1994, in Gann, 

 Conferences (HOTWC), CD-ROM, NIST SP 984-4, National Institute of Standards 

 April, 1994. 

ixed Counterflow Flames,” Proceedings of 

 Flames with 

 Mist Suppression of Methane-Air Diffusion Flames,” Combustion Science and Technology 132, 

University of Maryland, College Park, MD, 1989. 

55. DiMarzo, M., and Evans, D.D., Evaporation of Water Droplet Deposited on a Hot High Thermal 

Conductivity Solid Surfa

Gaithersburg, MD, 1986. 

56. Sheinson, R.S., Maranghides, A., Eaton, H., Barylski, 

Salmon, G., and Smith, W., “Total Flooding Agent Distribution Considerations,” 1995, in Gann, R.G., 

Burgess, S.R., Whisner, K.C., and Reneke, P.A., eds., Papers from

Working Conferences (HOTWC), CD-ROM, NIST SP 984-4, National Institute of Standards and 

Technology, Gaithersburg, MD, 2006. 

57. Carhart, H.W., Sheinson, R.S., Tatem, P.A., and Lugar, J.R., “Fire Suppression Research in the U.S. 

Navy, “Proceedings - First International Conference on Fire Suppression Research, Stockholm and Boras, 

Sweden, May 5-8, 1992. 

58. Ewing, C.T., Faith, F.R., Romans, J.B. Hughes, J.T. and Carhart, H.W., “Flame Extinguishment 

Properties of Dry Chemicals: Extinction Weights for Small Diffusion Pan Fires and Additional Evidence 

for Flame Extinguishment by Thermal Mechanisms,” Journal of F

(1992). 

59. Ewing, C.T., Faith, F.R., Romans, J.B., Siegmann, C.W

H.W., “Extinguishing Class B Fires with Dry Chemicals: Scaling Studies,” Fire Technology 31, 1-16, 

(1995). 

60. Sheinson, R.S., Eaton, H.G., Black, B.H., Brown, R., Burchell, H., Salmon, G., St. Aubin, J., and 

Smith, W.D., “Total Flooding Fire Suppress 3 3

Gann, R.G., Burgess, S.R., Whisner, K.C., and Reneke, P.A., eds., Papers from 1991-2006 Halon 

Options Technical Working Conferences (HOTWC), CD-ROM, NIST SP 984-4, National Institute of 

Standards and Technology, Gaithersburg, MD, 200

61. Sheinson, R.S., Eaton, H.G., Zalosh, R.G., Black, B.H., Brown, R., Burchell, H., Salmon, G.,and 

Smith, W.D. ,”Intermediate Scale Fire Extinguishment By

R.G., Burgess, S.R., Whisner, K.C., and Reneke, P.A., eds., Papers from 1991-2006 Halon Options 

Technical Working

and Technology, Gaithersburg, MD, 2006. 

62. Discussions between R.S. Sheinson and A. Baratov, 12

63. Zegers, E.J.P., Williams, B.A., Sheinson, R.S., and Fleming, J.W., “Dynamics and Suppression 

Effectiveness of Monodisperse Water Droplets in Non-Prem

the Combustion Institute 28, 2931-2937, (2000). 

64. Chen, N.H., Rogg, B., and Bray, K.N.C., “Modelling Laminar Two-Phase Counterflow

Detailed Chemistry and Transport,” Proceedings of the Combustion Institute 24, 1513-1521, (1992). 

65. Prasad, K., Li, C., Kailasanath, K., Ndubizu, C., Ananth, R., and Tatem, P.A., “Numerical Modeling 

of Water

325-364, (1998). 

66. Lefebvre, A.H., Atomization and Sprays, Hemisphere Publishing Corporation, pp. 91-92, 1988. 

67. Sirignano, W.A., Fluid Dynamics and Transport of Droplets and Sprays, Cambridge University Press, 

Cambridge, United Kingdom, pp. 5-6, 1999. 
References 473

nd 

lanar 

 Practical Guide, Spring-

: Concepts and Applications, 1st ed., McGraw Hill, Inc., 

ook 

97. 

pan, Paper 23, 17-19 July 2002. 

C2 Organic Flames,” Proceedings of the Combustion Institute 26, 

Review,” 

deling”, Combustion and 

uence of Droplet Number 

nal Journal of Heat and Mass Transfer 21, 251-253, (1978). 

on and Flame 115, 158-179, (1998). 

ty Journal 36, 55-72, (2001). 

o-flow Non-Premixed Flames”, 

g, J.W., “Inhibition of Premixed 

Methane/Air Flames by Water Mist,” Proceedings of the Combustion Institute 29, 361-368, (2002). 

68. Black, D.L., McQuayt, M.Q., and Bonint, M.P., “Laser-Based Techniques for Particle-Size 

Measurement: A Review of Sizing Methods and their Industrial Applications,” Progress in Energy a

Combustion Science 22, 267-306, (1996). 

69 Durst, F., Brenn, G., and Xu, T.H., “A Review of the Development and Characteristics of P

Phase-Doppler Anemometry,” Measurement Science and Technololgy 8, 1203–1221, (1997). 

70. Chigier, N., “Optical Imaging of Sprays,” Progress in Energy and Combustion Science 17, 221-262, 

(1991). 

71 Rafel, M., Willert, C.E., Kompenhans, J., Particle Image Velocimetry: A

Verlag Berlin, 1998. 

72. Turns. S, An Introduction to Combustion

New York, p. 314, 1996. 

73. Kays, W.M. and Crawford, M.E., Convective Heat and Mass Transfer, 2nd ed., McGraw-Hill B

Company, New York, p 388, 1980. 

74. Perry, R.H., Green., D.W., Maloney, J.O., Perry’s Chemical Engineers’ Handbook, Seventh Edition, 

The McGraw-Hill Companies, Inc., New York, p. 6-51, 19

75. Fleming, J.W., Williams, B.A., Sheinson, R.S., Yang, W., Kee, R.J., “Water Mist Fire Suppression 

Research: Laboratory Studies”, Proceedings of the Second National Research Institute for Fire Disaster 

Symposium: Technology Standards for Fire, Tokyo, Ja

76. Noto, T., Babushok, V, Burgess Jr., D.R., Hamins, A., Tsang, W., and Miziolek, A., “Effect of 

Halogenated Flame Inhibitors on C1-

1377-1383, (1996). 

77. Andrews, G.E., and Bradley, D., “Determination of Burning Velocities: A Critical 

Combustion and Flame 18, 133-153, (1972). 

78. Continillo, G., and Sirignano, W.A., “Counterflow Spay Combustion Mo

Flame 81, 325-340, (1990). 

79. Lacas, F., Darabiha, N., Versaevel, P., Rolon, J. C., and Candel, S., “Infl

Density on the Structure of Strained Laminar Spray Flames,” Proceedings of the Combustion Institute 24, 

1523- 1529, (1992). 

80. Seshadri, K. and Williams. F.A., “Laminar Flow Between Parallel Plates with Injection of a Reactant 

at High Reynolds Number,” Internatio

81. Lentati, A.M. and Chelliah, H.K., “Dynamics of Water Droplets in a Counterflow Field and their 

Effect on Flame Extinction,” Combusti

82. Yang, J.C., Donnelly, M.K., Prive, N.C., Grosshandler, W.L., An Apparatus For Screening Fire 

Suppression Efficiency Of Dispersed Liquid Agents,” Fire Safe

83. Hirst, R. and Booth, K., “Measurement of Flame-Extinguishing Concentrations,” Fire Technology 13, 

296-315, (1977). 

84. Fisher, B.T., Awtry, A.R., Sheinson, R.S., and Fleming, J.W., “Flow Behavior Impact on the 

Suppression Effectiveness of sub-10-μm Water Drops in Propane/Air C

Proceedings of the Combustion Institute 31, 2731-2739, (2007). 

85. Fuss, S.P., Chen, E.F., Yang, W., Kee, R.J., Williams, B.A., Flemin
474 Aerosol Properties 

rays: Influences of 

mooke, M.D., and Miller, J.A., A Fortran Program for Modeling Steady 

and Kee, R.J., “The Effect of Monodispersed Water Mists on the Structure, Burning 

, F.M. and Rupley, F.M., OPPDIF: A Fortran Program for Computing 

, S., Gardiner, Jr., W.C., Lissianski, V.V., and Qin, Z., GRI-Mech 3.0, 

s, B.A., Fleming, J.W., “Inhibition of Premixed 

ase Jr., M.W., Davies, C.A., Downey Jr., J.R., Frurip, D.J., McDonald, R.A., and Syverud, A.N., 

, 1985. 

he Combustion Institute, Pittsburgh, 

eedings Of The Eastern States Section, 

Propene Pyrolysis and Oxidation Kinetics in a Flow Reactor 

unication with H.K. Chelliah. 

86. Mitani,T., “A Flame Inhibition Theory by Inert Dust and Spray,” Combustion and Flame 43, 

243-253, (1981). 

87. Blouquin, R. and Joulin, G., “On the Quenching of Premixed Flames by Water Sp

Radiation and Polydispersity”, Proceedings of the Combustion Institute 27, 2829-2837, (1998). 

88. Kee, R.J., Grcar, J.F., S

Laminar One-Dimensional Premixed Flames, Sandia National Laboratories, SAND85-8240, 1987. 

89. Yang, W. 

Velocity, and Extinction Behavior of Freely Propagating, Stoichioetric, Premixed, Methane-Air Flames,” 

Combustion and Flame 130, 322-335, (2002). 

90. Kee, R.J., Grcar, J.F., Smooke, M.D., and Miller, J.A., A Fortran Program for Modeling Steady 

Laminar One-Dimensional Premixed Flames, Sandia National Laboratories, SAND85-8240, 1987. 

91. Lutz, A.E., Kee, R.J., Grcar

Opposed-Flow Diffusion Flames, Sandia National Laboratories, SAND96-8243, 1997. 

92. Smith, G.P., Golden, D.M., Frenklach, M., Moriarty, N.W., Eiteneer, B., Mikhail Goldenberg, M., 

Bowman, C.T., Hanson, R.K., Song

http://www.me.berkeley.edu/gri_mech/, 2001. 

93. Fuss, S.P., Chen, E.F., Yang, W., Kee, R.J., William

Methane/Air Flames by Water Mist,” Proceedings of the Combustion Institute 29, 361-368, (2002). 

94. Parks, D.J., Alvares, N.J., and Beason, D.G., “Fundamental Flame-Speed Measurements in 

Combustion Gases Containing CF3Br,” Fire Safety Journal 2, 237-247, (1979/80). 

95. Ch

JANAF Thermochemical Tables, 3rd Edition, Journal of Physical And Chemical Reference Data 14, 

927-1856, Suppl. 1

96. Noto, T., Babushok, V., Hamins, A., and Tsang, W., “Inhibition Effectiveness of Halogenated 

Compounds,” Combustion and Flame 112, 147-160, (1998). 

97. Davis, M.P., Fleming, J.W., Williams, B.A., and Ladouceur, H.D., “Flow Field Considerations for 

Counter Flow Burners”, Proceedings of the Eastern States Section, T

PA, pp. 200-203, Oct 1999. 

98. Lazzarini, A.K., Krauss, R.H., Chelliah, H.K., and Linteris, G.T, “Extinction Conditions of NonPremixed Flames with Fine Droplets of Water and Water/NaOH Solutions,” Proceedings of the 

Combustion Institute 28, 2939-2945, (2000). 

99. Chelliah, H.K., Lazzarini, A.K., Wanigarathne, P.C., and Linteris, G.T., “Inhibition of Premixed and 

Non-Premixed Flames with Fine Droplets of Water and Solutions,” Proceedings of the Combustion 

Institute 29, 369-376, (2002). 

100. Chelliah, H.K.,Williams, B.A., Mowrey, R.C., and Fleming, J.W., “Water Mist Suppression Of 

Propane-Air Flames: Modeling And Experimental Study,” Proc

The Combustion Institute, Pitsburgh, PA, 26-29 October 2003. 

101. Davis, S.G., Law, C.K., and Wang, H., “

and Laminar Flames,” Combustion and Flame 119, 375-399, (1999). 

102. H. Wang, U. Delaware, private comm
References 475

ings of the Combustion Institute 23, 503-511, (1990). 

eters, N., “Asymptotic Structure and Extinction of Methane-Air Diffusion 

e to Flame Stretch and Non-Premixed Turbulent Combustion,” 

mall Stoichiometric Mixture Fractions,” SIAM Journal of Applied Mathematics

tion Institute 28, 2931–2937, (2000). 

l Working Conference, 403-414, (1998). 

nts in a Counterflow 

 for Evaluating the Extinguishing Efficiency of Dry 

999, 

onate 

ann, R.G., Burgess, S.R., Whisner, K.C., and Reneke, P.A., 

103. Chelliah, H.K., Law, C.K., Ueda, T., Smooke, M.D., and Williams, F.A., “An Experimental and 

Theoretical Investigation of the Dilution, Pressure and Flow-Field Effects on the Extinction Condition of 

Methane-Air-Nitrogen Diffusion Flames,” Proceed

104. Li, S.C., Libby, P.A., and Williams, F.A., “Spray Structure in Counterflowing Streams with and 

without a Flame,” Combustion and Flame 94, 161-177, (1993). 

105. Seshadri, K., and P

Flames,” Combustion and Flame 73, 23-44, (1988). 

106. Chelliah, H.K., and Williams, F.A., “Aspects of the Structure and Extinction of Diffusion Flames in 

Methane-Oxygen-Nitrogen Systems,” Combustion and Flame 80, 17-48, (1990). 

107. Peters, N., “Local Quenching du

Combustion Science and Technology 30(1-6), 1-17, (1983). 

108. Kim, J.S., and Williams, F.A., “Structures of Flow and Mixture-Fraction Fields For Counterflow 

Diffusion Flames With S

53(6), 1551-1566, (1993). 

109. Zegers, E.J.P., Williams, B.A., Sheinson, R.S. and Fleming, J.W., “Dynamics and Suppression 

Effectiveness of Monodisperse Water Droplets in Non-Premixed Flames,” Proceedings of the 

Combus

110. Zegers, E.J.P., Williams, B.A., Fisher, E.M., Fleming, J.W., and Sheinson, R.S., “Suppression of 

Nonpremixed Flames by Fluorinated Ethanes and Propanes,” Combustion and Flame 121, 471-487, 

(2000). 

111. Berglund, R.N. and Liu, B.Y.H., “Generation Of Monodisperse Aerosol Standards”, Environmental 

Science and Technology 7(2), 147-153, (1973). 

112. Chelliah, H.K, Williams, B.A., and Fleming, J.W., “Water Mist Suppression of Propane-Air Flames: 

Interpretation of Experimental Results,” NRL/MR/6180--05-8853, March 10, 2005. 

113. Chin and Lefebvre, “Steady-State Evaporation Characteristics of Hydrocarbon Fuel Drops,” AIAA

21(10), 1437- 1443, (1983). 

114. Fleming, J.W., Reed, M.D., Zegers, E.J.P., Williams, B.A., and Ronald S. Sheinson, R.S., 

“Extinction Studies of Propane/Air Counterflow Diffusion Flames: The Effectiveness of Aerosols,” 

Proceedings of the Halon Options Technica

115. Reed, Mark D., A Study of the Behavior of Bicarbonate Powder Extinguisha

Propane/Air Diffusion Flame, M.S. thesis, Worcester Polytechnic Institute, Fire Protection Engineering, 

Worcester, MA, (work carried out at the NRL), 1997. 

116. Seeger, P.G., “A Laboratory Test Method

Powder,” AGARD Conference Proceedings Vol. 166, Paper 24, 1975. 

117. Williams, B.A. and Fleming, J.W., “Suppression Mechanisms of Alkali Metal Compounds,”' 1

in Gann, R.G., Burgess, S.R., Whisner, K.C., and Reneke, P.A., eds., Papers from 1991-2006 Halon 

Options Technical Working Conferences (HOTWC), CD-ROM, NIST SP 984-4, National Institute of 

Standards and Technology, Gaithersburg, MD, 2006. 

118. Trees, D. and Seshadri, K., “Experimental Studies of Flame Extinction by Sodium Bicarb

(NaHCO3) Powder,” Combustion Science and Technology 122, 215-230, (1997). 

119. Wanigarathne, P.A., Krauss, R.H., Chelliah, H.K., and Davis, R.J., “Fire Suppression by Particulates 

Containing Metallic Compounds,” 2000, in G
476 Aerosol Properties 

on of Opposed-Jet Methane Air Diffusion Flames - the Effects 

e Chemical 

Hale, E.T., “Flame Inhibition by Potassium Compounds,’’ Combustion and Flame 24, 277-279, 

chmoll, W. J., and Dressler, J. L., “Characterization of a Velocity-Modulation 

Handbook of Chemistry and Physics, 81st edition., CRC Press, Boca Raton, FL, 

Burner-Stabilized Premixed Laminar Flames by Boundary-Value 

e Origins of Spray Sectional Conservation 

rn States Section, 

ty of Virginia, Charlottesville, VA, 2001. 

f Water and Solutions,” Proceedings of the Combustion 

is, G.T., “Experimental Study of the Inhibition of Premixed and Diffusion 

ompounds,” Combustion and Flame 29, 221-238, (2002). 

mings, Menlo Park, CA, 1985. 

ygen 

and Nitrogen,” Proceedings of the Combustion Institute 24, 1083-1090, (1992). 

eds., Papers from 1991-2006 Halon Options Technical Working Conferences (HOTWC), CD-ROM, 

NIST SP 984-4, National Institute of Standards and Technology, Gaithersburg, MD, 2006. 

120. Friedman, R. and Levy, J.B., Inhibiti

of Alkali Metal Vapors and Organic Halides, Combustion and Flame 7, 195-201, (1963). 

121. Jensen, D.E., and Jones, G.A., “Kinetics of Flame Inhibition by Sodium,” Journal of th

Society, Faraday Transactions I 78, Part 9, 2843-2850, (1982). 

122. Mc

(1975). 

123. Takahashi, F., S

Atomizer,” Review of Scientific Instruments 65, 3563–3569, (1994). 

124. Lide, D.R., ed, CRC 

p. 8-108, 2000. 

125. Smooke, M.D., “Solution of 

Methods,” Journal of Computational Physics 4, 72-105, (1982). 

126. Greenberg, J.B., Silverman, I., and Tambour, Y., “On th

Equations,” Combustion and Flame 93, 90–96, (1993). 

127. D’Angelo, Y., Silverman, I., Gao, L. P., Gomez, A., and Smooke, M. D., “Computational and 

Experimental Study of Counterflow Spray Diffusion Flames,” Proceedings of the Easte

The Combustion Institute, Pttsburgh, PA, 1996. 

128. Wanigarathne, P.C., Fire Suppression in Studies with Particulates Containing Metallic Compounds,’ 

M.S. thesis, Universi

129. Chelliah, H.K., Lazzarini, A.K., Wanigarathne, P.C., and Linteris, G.Y., “Inhibition of Premixed and 

Non-premixed Flames with Fine Droplets o

Institute 29, 369-376, (2002). 

130. Reinelt, D., and Linter

Flame by Iron Pentacarbonyl,” Proceedings of the Combustion Institute 26, 1421-1428, (1996). 

131. Linteris, G.T., Knyazev, V., and Babushok, V., “Inhibition of Premixed Methane Flames by 

Manganese and Tin C

132. NIST Chemistry WebBook, no. 69, http://webbook.nist.gov/chemistry/, release July 2001. 

133. Vanpee, M., and Shirodkar, P.P., “A Study of Flame Inhibition by Metal Compounds,” Proceedings 

of the Combustion Institute 17, 787-795 , (1978). 

134. Patrick, R., and Golden, D.M., “Termolecular Reactions of Alkali Metal Atoms with O2 and OH,” 

International Journal of Chemical Kinetics 16, 1567-1574, (1984). 

135. Lask, G., and Wagner, H.G., “Influence Of Additives On The Velocity Of Laminar Flames,” 

Proceedings of the Combustion Institute 8, 432-438, (1962). 

136. da Cruz, F.N., Vandooren, J., and Van Tiggelen, P., “Comparison Of The Inhibiting Effect Of Some 

Halogenomethane Compounds On Flame Propagation,” Bull. Soc. Chim. Belg. 97, 1011–1030, (1988). 

137. Williams, F.A., Combustion Theory, 2nd ed., Benjamin/Cum

138. Chelliah, H.K., Yu, G., Hahn, T.O., and Law, C.K., “An Experimental and Numerical Study on the 

Global and Detailed Kinetics of Premixed and Nonpremixed Flames of Chloromethane, Methane, Ox
References 477

., Naylor, A.M., and Goddard, W.A., “Starburst Dendrimers - Molecular Level Control 

alogh, L., Dendritic Polymers as Fire Suppressants, ARLoving 

or Screening Fire 

al Institute of Standards and Technology, Gaithersburg, 

Fire 

yldedecyclammonium 

wa, “Experiments on the Atomization of Liquids in Air-Stream,” Transactions 

agnauer, G., “The Properties of Dendritic 

ndbook of Chemistry and Physics, 81st edition, CRC Press, Inc., Boca Raton, 

reen,” 1999, in Gann, R.G., Burgess, S.R., Whisner, K.C., 

iew,” Proceedings of the First 

International Symposium of Fire Safety Science, C. E. Grant And P.J. Pagni, Editors, Hemisphere, 

Washington, pp. 1145-1163, 1986. 

139. Tomalia, D.A., “Dendrimer Molecules,” Scientific American 272, 62-66, (1995). 

140. Tomalia, D.A

of Size, Shape, Surface Chemistry, Topology, and Flexibility from Atoms to Macrocopic Matter,”

Angewandte Chemie. International Edition 29, 138-175, (1990). 

141. Tapscott, R.E., “Current Topics on New Halon Replacements,” National Industrial Research 

Institute, Agency of Industrial Science and Technology, Nagoya, Japan, 10 April 1997. 

142. Babushok, V.I. and Tsang, W., “Inhibitor Rankings for Hydrocarbon Combustion,” Combustion and 

Flame 123, 488-506 (2000). 

143. Beck Tan, N.C., DeSchepper, D.C., and B

TR-2071, Army Research Laboratory, Aberdeen Proving Ground, October, 1999. 

144. Finnerty, A.E., McGill, R. L., and Slack, W.A., Technical Report ARL-TR-1138, Aberdeen Pr

Ground, Maryland, 1996. 

145. Yang, J.C., Donnelly, M.K., Prive, N.C., and Grosshandler, W.L., “An Apparatus f

Suppression Efficiency of Dispersed Liquid Agents,” 1998, in Gann, R.G., Burgess, S.R., Whisner, K.C., 

and Reneke, P.A., eds., Papers from 1991-2006 Halon Options Technical Working Conferences 

(HOTWC), CD-ROM, NIST SP 984-4, Nation

MD, 2006. 

146. Yang, J.C., Donnelly, M. K., Prive, N.C., and Grosshandler, W.L., “Dispersed Liquid 

Suppression Screen Apparatus,” NISTIR-6319, Gaithersburg, MD, 1999. 

147. Chen, C.Z., Beck Tan, N.C., and Cooper, S.L., “Incorporation of Dimeth

Chloride Functionalities onto Poly(propylene) Dendrimers Significantly Enhances Their Anti-bacterial 

Properties,” Chemical Communications 16, 1585-1586, (1999). 

148. Nukiyama and Tanasa

of the Society of Mechanical Engineers 4,5,6, (1938-1940). 

149. Uppuluri, S. , Dvornic, P. R., Beck Tan, N. C., and H

Polymers II: Generation Dependence of the Physical Properties of Poly(amidoamine) Dendrimers,” 

ARL-TR-1774, Aberdeen Proving Ground, Maryland, 1999. 

150. Lide D.R., ed., CRC Ha

FL, p.6-3, 2000. 

151. Yang, J.C., Prive, N.C., Donnelly, M.K., and Grosshandler, W.L., “Recent Results from the 

Dispersed Liquid Agent Fire Suppressant Sc

and Reneke, P.A., eds., Papers from 1991-2006 Halon Options Technical Working Conferences 

(HOTWC), CD-ROM, NIST SP 984-4, National Institute of Standards and Technology, Gaithersburg, 

MD, 2006. 

152. Pitts,W.M., Yang, J.C., Bryant, R.A., Blevins, L.G., and Huber, M., Characteristics and 

Identification of Super-Effective Thermal Fire-Extinguishing Agents, Final Report, NGP Project 

4C/1/890, NIST Technical Note 1440, National Institue of Standards and Technology, Gaithersburg, MD, 

June 2001. 

153. Downie, B., Polymeropoulos, C., and Gogos, G., “Interaction of a Water Mist with a Buoyant 

Methane Diffusion Flame,” Fire Safety Journal 24, 359-381, (1995). 

154. Rasbash, D.J., “The Extinction Of Fire With Plain Water: A Rev
478 Aerosol Properties 

ansfer 35, 2377-2388, (1992). 

51-173, (1995). 

Journal of Fluid Mechanics 283, 141-173, 

h Spray Quenching,” Inernational Journal of Heat and Mass Transfer 39, 

nternational Journal of Heat and Mass Transfer 40, 73-88, 

rops Impinging on a Polished Surface,” International Journal of Heat and Mass Transfer 40, 

Boiling Heat Transfer of Droplet Streams and Sprays,” 

e For Spray Cooled 

, J.C., and Mudawar, I., “Film And Transition Boiling Correlations For 

ansfer 37, 347-362, 

International Journal of Heat and Mass Transfer

ter and Critical Heat Flux for Spray 

 On Solid And Liquid Surfaces,” Fluid Dynamic 

or Explosions,” ASME 

155. Engel, O. G., “Waterdrop Collisions With Solid Surfaces,” Journal of Research of the National 

Bureau of Standards 54, 281-298, (1955). 

156. Chandra, S. and Avedisian, C.T., “On The Collision Of A Droplet With A Solid-Surface,”

Proceedings of the Royal Society of London A-Mathematical Physical and Engineering Sciences 432, 

13-41, (1991). 

157. Chandra, S. and Avedisian, C.T., “Observations of Droplet Impingement on a Ceramic Porous 

Surface”, International Journal of Heat and Mass Tr

158. Mundo, C., Sommerfeld, M., and Tropea, C., “Droplet-Wall Collisions - Experimental Studies of the 

Deformation and Breakup Process,” International Journal of Multiphase Flow 21, 1

159. Yarin, A.L. and Weiss, D.A., “Impact of Drops on Solid-Surfaces - Self-Similar Capillary Waves, 

and Splashing as a New-Type of Kinematic Discontinuity,” 

(1995). 

160. Bernardin, J.D. and Mudawar, I., “Experimental and Statistical Investigation of Changes in Surface 

Roughness Associated wit

2023-2037, (1996). 

161. Bernardin, J.D., Stebbins, C.J., and Mudawar, I., “Effects of Surface Roughness on Water Droplet 

Impact History and Heat Transfer Regimes,” I

(1997). 

162. Bernardin, J.D., Stebbins, C.J., and Mudawar, I., “Mapping of Impact and Heat Transfer Regimes of 

Water D

247-267, (1997). 

163. Bernardin, J.D. and Mudawar, I., “Film 

International journal of Heat and Mass Transfer 40, 2579-2593, (1997). 

164. Mudawar, I. and Valentine, W.S., “Determination Of The Local Quench Curv

Metallic Surfaces,” ASME Journal of Heat Treating 7, 107-121, (1989). 

165. Klinzing, W.P., Rozzi

Quenching Of Hot Surfaces With Water Sprays,”ASME Journal of Heat Treating 9, 91-103, (1992). 

166. Mudawar, I. and Deiters, T.A., “A Universal Approach to Predicting Temperature Response of 

Metallic Parts to Spray Quenching”, International Journal of Heat and Mass Tr

(1994). 

167. Hall, D.D. and Mudawar, I., “Experimental and Numerical Study of Quenching Complex-Shaped 

Metallic Alloys with Multiple, Overlapping Sprays,” 

38, 1201-1216, (1995). 

168. Estes, K.A. and Mudawar, I., “Correlation of Sauter Mean Diame

Cooling of Small Surfaces,” International journal of Heat and Mass Transfer 38, 2985-2996, (1995). 

169. Mudawar, I. and Estes, K.A., “Optimizing And Predicting Chf In Spray Cooling Of A Square 

Surface,” ASME Journal of Heat Transfer 118, 672-679, (1996). 

170. Rein, M., “Phenomena Of Liquid-Drop Impact

Research 12, 61-93, (1993). 

171. Henry, R.E. and Fauske, H.K., “Nucleation Processes In Large-Scale Vap

Journal of Heat Transfer 101, 280-287, (1979). 
References 479

NIST Standard 

es of Gases & Liquids, 4th ed., McGrawo, Y.M., and Chandra, S., “Spray Cooling Enhancement by Addition of a Surfactant,” ASME 

t Transfer of Impacting Sprays,” 

jka, P.E., and Sivathanu, Y.R., “Water Mist Impingement onto a Heated Surface,” 

ction Between Evaporating Sprays and 

 a Hot Surface, M.S. Thesis, School of Mechanical 

fer Problems Advances in Heat Transfer,” 

Numerical Heat Transfer, John Wiley & Sons Inc., NY, 1988. 

quation of 

n to Fluid Mechanics, John Wiley & Sons (SEA) Pte. 

 Single-Sample Experiments,” ASME 

 for the Production of Monodisperse Sprays 

g, 

172. Spalding, D.B., Some Fundamentals of Combustion, Butterworths Scientific Publications, London, 

1955. 

173. Kanury, A.M., Introduction to Combustion Phenomena, Gordon & Breach, New York, 1975. 

174. DIPPR Data Compilation of Pure Compound Properties Database, V9.02, 

Reference Data Program #11, NIST, Gaithersburg MD, 1995. 

175. Reid, R.C., Prausnitz, J.M., and Poling, B.E., The Properti

Hill, New York, 1987. 

176. Qia

Joournal of Heat Transfer 120, 92-98, (1998). 

177. Deb, S. and Yao, S.-C., “Analysis on Film Boiling Hea

International Journal of Heat and Mass Transfer 32, 2099-2112, (1989). 

178. Fu, C., So

Proceedings of the 5th Joint ASMElJSME Joint Thermal Engineering Conference, 1999. 

179. Fu, C., Sojka, P.E., and Sivathanu, Y.R., “On the Intera

Heated Surfaces,” Proceedings of the 12th Annual Conference on Liquid Atomization and Spray Systems, 

1999. 

180. Oke, H.P., An Experimental Study of Flame Spread over PMMA Subject to a Water Mist, M.S. 

Thesis, School of Mechanical Engineering, Purdue University, November, 1999. 

181. Chunming Fu., Droplet Interactions with

Engineering, Purdue University, March 2000. 

182. Howell, J.R, “Application Of Monte Carlo To Heat Trans

in Advances In Heat Transfer, Vol. 5, J. P. Hartnett and T. Irvine, Eds., Academic Press, New York, p. 1, 

1968. 

183. Haji-Sheik, Handbook of 

184. Kahn, H., and Marshall, A. W., “Methods Of Reducing Sample Size In Monte-Carlo Computations,” 

Journal of the Operations Research Society of America 1, 263-278, (1953). 

185. Lanore, J.M., “Weighting and Biasing of a Monte Carlo Calculation for Very Deep Penetration Of 

Radiation,” Nuclear Sciene and Engineering 45, 66-72, (1971). 

186. Sivathanu, Y.R., and Gore, J.P., “A Discrete Probability Function-Method for the E

Radiative-Transfer,” Journal of Quantitative Spectroscopy and Radiative Transfer 49, 269-280, (1993). 

187. Fox, R.W., and McDonald, A.T., Introductio

Ltd., Singapore, 1995. 

188. Kline, S.J., and McClintock, F.A., “Describing Uncertainties In

Mechanical Engineering 75, 3-8, (1953). 

189. Brenn, G., Helpio, T., and Durst, F. , “A New Apparatus

at High Flow Rates,” Chemical Engineering Science 5, 237-244 , (1997). 

190. Dingel, B.E., An Experimental and Theoretical Study of Air and Water Mist Impingement Coolin

M.S. thesis, Mechanical Engineering, Purdue University, 1997. 

191. Lund, M.T., Sojka, P.E., Lefebvre, A.H., Gosselin, P.G., “Effervescent Atomization at Low Mass 

Flow Rates. Part I: The Influence Of Surface Tension,” Atomization and Sprays 3, 77-89, (1993). 
480 Aerosol Properties 

l Investigation on the Collision Behavior 

192. Masters, K., Spray Drying, 2nd Edition, John Wiley and Sons, New York, 1976. 

193. Ranz, W.E., and Marshall, W.R, “Evaporation from Drops,” Chemical Engineering Progress 48, 

173-180, (1952). 

194. Jiang, Y.J., Umemura, A., and Law, C.K., “An Experimenta

of Hydrocarbon Droplets,” Journal of Fluid Mechanics 234, 171-190, (1992).