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Performance and CO Production of a Non-Azide AirbagPropellant in a Pre-Pressurized Gas GeneratorRobert G. Schmitt a , P. Barry Butler a & Jon J. Freesmeier aa Department of Mechanical Engineering, The University of Iowa , Iowa City, Iowa, 52242Published online: 06 Apr 2007.

To cite this article: Robert G. Schmitt , P. Barry Butler & Jon J. Freesmeier (1997) Performance and CO Production of a Non-Azide Airbag Propellant in a Pre-Pressurized Gas Generator, Combustion Science and Technology, 122:1-6, 305-330, DOI:10.1080/00102209708935613

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Corabust. Sci. and Tech.. 1997. Vol. 122. pp.305-330Reprints available directly from the publisherPhotocopying permitted by license only

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license by Gordon lind Breach Science PublishersPrinted in India

Performance and CO Productionof a Non-Azide Airbag Propellantin a Pre-Pressurized Gas Generator

Robert G. Schmitt*, P. Barry Butler and Jon J. Freesmeier

Department of Mechanical Engineering, The University of Iowa.Iowa City, Iowa 52242

(Received 15 May 1996; In final form 30 September 1996)

This paper presents a numerical study of the transient operation of a pre-pressurized (aug­mented) airbag inflator. Augmented inflators dilute hot gaseous products of propellant com­bustion with ambient temperature, high-pressure stored gas before discharging the mixtureinto the airbag. The solid propellant selected for this study is a non-azide propellant composedof a mixture of azodicarbonirnide, potassium perchlorate, and cupric oxide, Predicted perfor­mance of the inflator is presented in terms of pressure, temperature and mass flow rate profilesin the inflator and discharge tank which is used to simulate an airbag. This work also predictsfirst-order estimates of gas-phase species exit concentrations and characteristic residence timesin the inflator. Carbon monoxide. produced as a product of combustion from the high flametemperature propellant, is partially converted to COz as it flows from an internal combustionchamber to the pressurized plenum before being discharged into the airbag, Specifically. theproduction/destruction of CO is tracked using three different gas-phase reaction models: I)chemically frozen. 2) local (shifting) equilibrium. and 3) finite-rate elementary kinetics. Resultspresented in this paper demonstrate the necessity of an airbag combustion program thatincludes finite-rate, gas-phase kinetics. Results from the finite-rate CO chemistry model arequalitatively consistent with experimental results reported by others for the same propellantformulation in a similar operating environment.

Keywords: Airbag; gas generator; solid propellant

INTRODUCTION

As discussed in a previous review article (Butler et al.,1993a), most present­day airbag gas generators can be classified as either "pyrotechnic" or

·Current address: Organization 9112, Sandia National Laboratories, Albuquerque,NM 87185.

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306 R. G. SCHMITT er al.

"augmented" devices based on the configuration used to produce and de­liver the quantity of gas required for inflating a vehicle airbag. The majordifference between the two categories (Butler et al., 1993b; Butler et al.,1993c; Berger and Butler, 1995) is pyrotechnic inflators rely solely on rapidgas production from the solid propellant, whereas augmented devices dilutehot gaseous products of propellant combustion with ambient temperature,high-pressure stored gas before discharging the mixture into the airbag. Asan historical note, the first generation airbags of the 1960's simply dis­charged a canister of pressurized gas stored at ambient temperature (Vosand Goetz, 1989). Basic choked-flow gas dynamic relations (Zucrow andHoffman, 1976) show the gas discharge rate from a stored-gas unit is in­versely dependent on ambient sound velocity. Augmented units overcomethis deficiency by mixing hot products of propellant combustion with theambient-temperature stored gas, producing a mixture less dependent onvariations in ambient temperature. This is due to the weak dependency ofpropellant flame temperature on ambient conditions. Performance specifica­tions for airbag gas generators require minimal variation in output over awide range of operating temperatures (240< T(K) < 320), and thus a consist­ent inflator temperature is desirable to achieve a consistent discharge flowrate.

Over the past decade, the most common propellants used as gas gene­rants in pyrotechnic inflators have been azide-based. This includes composi­tions such as sodium-azide (NaN3) with metal-oxides (e.g., CuO or Fe20 3)

serving as the oxidizer. Azide-based propellants are favorable in airbagapplications because of their low flame temperatures relative to other pro­pellants, and because their gas-phase products of combustion consist of alarge percentage of harmless nitrogen gas. However, along with the highlydesirable N 2 they produce, azides also generate considerable amounts ofcondensed-phase residue that must be removed from the products prior toentering the airbag. For example, a typical azide-based propellant producesnitrogen and condensed-phase slag on a 1:1 mass ratio at an adiabatic flametemperature of approximately 1,200 K. This is a considerable amount ofslag that must be filtered from the products. Furthermore, azides can ex­hibit ignition difficulties and non-homogeneous combustion characteristics.Two-examples of non-azide propellants currently used as gas generators forairbags include: nitrocellulose-based substances and carbonamide-basedpropellants (Berger and Butler, 1995; Hara et al., 1994). These propellantstypically produce less slag, but burn at much higher flame temperaturesthan azides. In addition, their product gas is generally a mixture of severalspecies including N 2, CO 2, and H20 , and also contain some less desirable

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NON·AZIDE AIRBAG PROPELLANT 307

minor species such as CO and NOx- In general, the amount of condensed­phase slag produced by non-azide airbag propellants is significantly lessthan that of azide-based propellants. In fact, some non-azide propellantscan be chemically formulated to produce 100% gas-phase products, elimi­nating the costly requirement for slag filtration.

BACKGROUND

Augmented inflators demonstrate several advantages over conventionalpyrotechnic units (Butler et al., 1993a) including: i) reduced propellant massrequirements due to the mass of pressurized gas stored in the unit, ii) alower average temperature of the gas mixture discharged into the airbag, iii)dilution of unwanted gaseous species produced by the propellant, and iv)more uniform performance under hot and cold operating conditions. Gasdilution is important since it allows the inflator designer more flexibility inselecting a propellant. For example, a fuel-rich propellant used in a corecombustion chamber can be enhanced with an oxidizing gas in an adjacenthigh-pressure chamber. This produces a two-stage combustion process; pro­pellant decomposition in the core chamber followed by combustion or thepropellant's excess gas-phase fuel species with the pre-pressurized oxidizingagent. However, one disadvantage of augmented units is the necessity tostore the augmentation gas at high pressure (e.g., 100-300 bar) for longperiods of time (> 20 years).

In this paper, the physical model and fundamental assumptions are pres­ented for a prepressurized airbag inflator. Inflator performance indices (But­ler et al., 1993) include: maximum combustion chamber pressure p,.ma,'maximum tank pressure p,.ma,' maximum tank temperature T,.max' and tankimpulse over the period 0 < t > 100 ms, 1100 , Another performance index ofinterest is the CO concentration in the discharge gas and its interactionwith gas-phase product species present in the inflator. Estimates of gas­phase exit concentrations and characteristic residence times in the inflatorare predicted. Here, the production/destruction of CO is tracked during theinitial inflation time period (0-100 ms) using three different gas-phase reac­tion models: 1) chemically frozen, 2) local (shifting) equilibrium, and 3)finite-rate elementary kinetics.

The propellant selected for this study is a non-azide propellant composed ofa mixture of azodicarbonimide (ADCA), potassium perchlorate (KCl04 ), andcupric oxide (CuO). The propellant was selected because it represents a rela­tively new formulation developed to compete with the traditional azide-based

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308 R. G. SCHMITT et al.

propellants (Hara et al., 1994). Research on this propellant has reported on itsdesirable performance, safety, health, and environmental aspects (Hara et al.,

1994a; Hara et al., 1994b). One drawback to the propellant is its tendencyto produce high levels of CO and NOx- Several combustion modificationswere investigated to reduce CO and NO, levels to acceptable levels (Haraet al., I994a). The design changes were primarily modifications to the pro­pellant chemistry and the use of various catalyst. Here, we will explore theeffect of a dualchamber combustion technique in reducing CO levels byconversion to CO2 as it flows from the combustion chamber to the pressur­ized plenum before being discharged.

Modeling combustion, mixing, and discharge requires the mathematicaltreatment of i) the thermochemistry of the pyrotechnic combustion, ii) vari­ous modes of heat exchange between products of combustion and hardwarecomponents of the inflator, iii) compressible flow gas dynamics of the dis­charged products, and iv) finite-rate gas-phase kinetics of select species. Toassist in describing this model, Figure 1 is provided as a simplified sketch ofa typical augmented inflator being discharged into a constant-volumevessel. The following is a brief description of the major components of theinflator package (Butler et al., 1993a).

A pyrotechnic ignitor and grains of solid propellant are located in theinterior combustion chamber. The chamber is hermetically sealed from therest of the inflator by a thin rupture film which acts as a guard against

InflatorBody

FillerScreen

Ignitor ....b~II1e-J~PI'!IiIii1~ft1~wlr- ....~ Rupture1P!i!!!!Ii~~~~~~~[J--F'tIm #1

FilterScreen

RuptureFilm#2

BodyExi/l...•l.Nozzles ,

DISCHARGETANK

FIGURE 1 Key components of typical augmented inflator (cross-sectional view) emptyinginto a constant-volume discharge tank,

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NON-AZIDE AIRBAG PROPELLANT 309

moisture leakage and as a means to achieve volumetric confinement of thepropellant product gases for the first few milliseconds following ignition.Confinement of the products increases the rate of pressurization and helpssupport a rapid flame spread throughout the numerous propellant granulestightly packed in the combustion chamber. As demonstrated later in thispaper, momentary confinement of the products results in additional COproduction due to the compression heating of the combustion productsprior to pressure release. Adjacent to the combustion chamber is a gasplenum filled with a pressurized gas. The gas plenum is sealed from thesurroundings with a second rupture film. At the exit surface of the plenum isa stainless-steel filter screen that serves to cool the gas and to capture thecondensed-phase products of combustion. As the amount of gas producedby the propellant decreases relative to the amount stored in the prepres­surized plenum, the cooling and filtering requirements of the filter screendiminish. In most inflator qualification experiments, the product gases areexhausted into a constant-volume discharge tank as Figure 1 Illustrates.The discharge tank is fixed by industry requirements, and is generally 20,60, or 120 L in volume to mimic a fully deployed airbag. It should beemphasized that there are obvious deficiencies in using the tank test as anairbag simulator. In an airbag system, the product gases are discharged intoan inflatable bag made of tightly woven polymer fibers. The bag unfoldsand allows gas leakage through the fibrous mesh as it fills. For comparison,in the tank test configuration the products are usually discharged into arigid reservoir initially filled with air at ambient temperature. Mixing thehot products of combustion with ambient temperature tank air does notprovide the same thermodynamic process as filling a porous, initially col­lapsed airbag.

A typical sequence of events that occurs during airbag inflation can bedescribed as follows. Initially, a collision sensor system electronically firesan ignitor located in the combustion chamber of the inflator. The ignitorimmediately releases high temperature gas and condensed-phase productsfrom combustion of a small pyrotechnic charge (e.g., BKN03) . The result isa rapid increase in pressure and temperature within the combustion cham­ber. Heat flux from the high temperature ignitor gas and hot particles theninitiates the main propellant grains adjacent to the ignitor, and the deflagra­tion front rapidly spreads throughout the remaining grains. Decompositionof the main propellant grains rapidly increase the pressure and temperaturein the combustion chamber. When the combustion chamber pressurereaches .a design value, the internal rupture film opens and the high tem­perature gas/condensed-phase products are discharged into the gas plenum.

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310 R.G. SCHMIlT er aJ.

In the plenum they are mixed with a pressurized gas (argon in the presentstudy) that is at ambient temperature.

As the gas plenum pressure increases, the external rupture film isbreached and the mixture of combustion products and pressurized gas arethen exhausted into the airbag. As the inflator effluent exits the plenum, thegas and particles pass through the filter screen. The filter screen captures thecondensed-phase slag and cools the gas as it permeates through the high­surface-area wire mesh. As the process continues, the main propellant in thecombustion chamber burns for approximately 50-100 ms before beingcompletely consumed. For comparison, side impact units operate about fivetimes faster. During the entire event, heat transfer takes places between thegas phase, condensed phase and inflator hardware components in eachchamber.

In the combustion model presented here, the pre-pressurized inflator de­picted in. Figure I is divided into three discrete control volumes: i) thecombustion chamber including propellant grains and ignitor, ii) the gasplenum containing the pressurized inert gas, and iii) the discharge tank. Forthe dynamic thermochemical event occurring in the inflator and dischargetank, the basic assumptions listed in Table I were imposed in developingthe governing equations. Based on these assumptions, the governing conser­vation equations for the inflator model were derived by expressing conser­vation conditions for species mass and energy contained in each of the threecontrol volumes. The result is a system of stiff ordinary differential equa­tions that express mass conservation for each gas/condensed phase species,separate energy equations for gas phase and condensed phase, respectively,and energy equations for the filter screen and inflator hardware. Figure 2

TABLE I Model Assumptions

• Gas- and condensed-phases are composed of muitiple species with temperature-depend­ent specific heats,

• The gas phase is a well-mixed.ideal gas and condensed-phase species are incompressible.

• The solid propellant decomposes according to an experimentally verified reaction equa­tion, producing equilibrium product species at the adiabatic flame temperature.

• Heat transfer occurs between gas/solid phases and hardware components of the inflatorand tank.

• The fiiter does not accumulate gas-phase species, but it can accumulate solid/liquidparticles.

• The instantaneous collection of condensed-phase product species by the filter is in thesame proportions (mass fractions) as the condensed-phase product mixture flowingthrough it.

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NON-AZIDE AIRBAG PROPELLANT 311

illustrates the computational network assembled for the inflator shown inFigure 1.

GOVERNING EQUATIONS

Conservation Equations

The current model keeps track of multiple gas and condensed-phase chemi­cal species present in all computational cells (combustion chamber, gasplenum and discharge tank). In the results presented here, a total of 20different chemical species are considered, where the species can exist as gas,liquid, or solid, depending on the local state of the system. All species aremodeled with temperature-dependent specific heats Cp(T) expressed aspolynomial functions (Chase et al., 1985). Condensed-phase constituents areassumed incompressible, and all gas-phase thermodynamic properties arebased on the ideal gas equation of state. Standard-state species enthalpiesare determined by integrating the standard-state specific heat functions,

H. = IT Cp.dT + H U 9 8298

(1)

Species internal energies are determined from the fundamental ther­modynamic relation U. = H. - RT, and mixture properties Hand U arebased on standard mixture summing rules.

The governing equations for the airbag inflator are derived by applyingconservation principles of species mass and energy of each phase for eachcomputational cell shown in Figure 2. In this zero-dimensional, multi-zonedmodeling approach, the properties of each cell are volume averaged (i.e., nospatial variation within the cells). Combustion of the propellant m',g,. andignitor m.,;g. to produce the k-th species is restricted to the combustionchamber. This assumption is consistent with the present inflator designwhich meters the flow from combustion chamber to plenum through anumber o( small orifices. Finite-rate, gas-phase chemistry is modeled in allcomputational cells. The resulting species equations are,

Combustion chamber

(2)

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312 R.G. SCHMITf er al.

VOLUMES HARDWARE

/ Propellant Grains

High-PressureGas Plenum (V=6Occ)

DischargeTank(V=30,OOOCC)

- :1"0bastion0=1llich3iiie Wall

...Heai

Exchange

FIGURE 2 Computational network used in present model to simulate augmented inflator.

Gas plenum

dm::.:.:..:!d'.=w -Y. riJ +Y. m k=l,KKde k,p k,p out,p k,c out.c

Discharge tank

dmk •1 =w + Y. riJ k = I, KKdt Ie,f k,p out,p

(3)

(4)

Separate conservation of energy expressions are written for both the gasand condensed phases present in each of the three cells to differentiategas-phase and condensed-phase temperatures. The resulting gas-phase en­ergy equations for each cell arc written in terms of temperature as,

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NON·AZIDE AIRBAG PROPELLANT

Combustion chamber

dTg., 1 [ KK d"-'- - U W g" - Hg.,m P vg., _'1'_,dt - mgas ega! L k,e Ie.c + Qc C out,c + C C dtc v.c 1e=1

Gas plenum

313

(5)

d Tg., 1 [ KK d" ]:...::....l?.. • gas gas' gas 'f'p gas .dt = m&"Cg" L ti, .•W k•• + Q. - H. mout •• + p.V. dt + H, mout.,

p v.p Ie= 1

(6)

Discharge tank

dP"' 1 [ KK ]I _ • gas gasrh----;It - mg" Cga> L Uk.,Wk., + Q. + H. out.p

I v.t k= 1

(7)

In Eqs. (5-7) Qr' represents the net heat flux to the gas in each cell (j=c, p, t).This includes the summation of all heat transfer paths illustrated inFigure 2 in addition to particle-gas heat exchange. The enthalpy flux termsappearing in Eqs. (5- 7) are consistent with the assumption of volume-aver­aged gas-phase properties in each cell. The condensed-phase products ex­change heat with the surrounding gas and with the filter hardware whenthey are captured by it.

Propellant decomposition is modeled as a function of pressure and ambi­ent temperature. The surface regression rate takes the common form,

dr_=aea~Tp'

dt(8)

where r is the burn depth measured from the initial surface of the grains and6 T = T - 298 K. The coefficients a and n appearing in Eq. (8) normallycome from closed-bomb, burn-rate experiments where the regression rate ofthe propellant is measured as a function of pressure. The temperature­sensitivity coefficient (J comes from experiments conducted over a widerange of propellant ambient temperatures To in order to account for thebehavior of inflator propellants under extreme operating conditions.Equation (8) does not include a term to account for effects of pressure

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314 R.G. SCHMITT et al.

transients on burn rate. Given the initial pressure rise-rates common to mostsealed airbag inflators, this term may be significant to accurately modelingthe performance of the propellant over the initial pressurization period (Kuoand Summerfield, 1984). This term was not included in Eq. (8) for the currentstudy due to limited information on the dynamic pressure coefficient. Thetotal mass production rate resulting from propellant combustion is thus,

mgon = N i pA

and the production of species k is given as,

(9)

(10)

where 1;.,gon is the mass-fraction of species k produced by surface reaction ofthe propellant. In Eq. (9), N represents the total number of propellantgrains in the combustion chamber, A is the instantaneous surface area of agrain, and p is the propellant true density adjusted for the presence of voidsin the grains. The variation of propellant burn area with burn depth A isreferred to as the form function of the grain and is dependent on the graingeometry and flame-spreading characteristics. For brisant ignitors in tightlypacked combustion chambers some grain fracture usually occurs duringignition. This increases the actual burn area above the theoretical value.The grains are assumed to burn uniformly on all exposed surfaces and donot fracture on ignition. Thus, the form function is dependent only ontheoretical grain geometry. The current analysis uses a cylindrical grainshape with initial dimensions h = 5.44 mm and d = 12.8 mm. The ignitordecomposition mign is modeled in a similar fashion to Eqs. (9) and (10).

In this analysis a set of 33 elementary, reversible gas-phase chemicalreactions involving KK chemical species are modeled,

(i = 1, ... ,II) (II)

where Vk i are the stoichiometric coefficients and Xk represent chemical sym­bols. Mass production rate of the k-th species in each computational cellcan be written as a summation of the rate of progress variables for allreactions involving the k-th species,

(k=I, ... ,KK)

(12)

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NON-AZIDE AIR BAG PROPELLANT 315

The forward rate constants kjt for the reactions are assumed to have thefollowing Arrenhius temperature dependence,

kl i = AiTP;exp( - EJ RT) (13)

Reverse rates are computed from the forward rates and the concentration­based equilibrium constants,

k.=~rI K; (14)

Gas flow exitmg the combustion chamber and plenum cells is eitherlocally sonic or subsonic depending on the mixture specific heat ratio of thehigh-pressure cell and the difference in pressure between the adjacent cell towhich it is transporting mass. From basic laws of gas dynamics, the gas­phase mass flow rate through the discharge nozzles is given by the followingexpressions,

Subsonic

Sonic

~ p .p .(p2/Y_ ply+ 1)/y)Y-1 hi hI hi 10

(15a)

(2 )y+ 1/2(y-1)

mou• = ACdJyPhiPhi Y+ I (ISb)

where either Eqs. (ISa) or (15b) is used depending on the pressure ratiobetween adjacent cells. The parameter Cd is a nozzle discharge coefficient,and y is the specific heat ratio of the gas, defined as y = Cp(T, cornposi­tion)/Cu(T, composition), is determined from the gas-phase compositionand temperature. The condensed-phase mass-flow rate between the combus­tion chamber and plenum is assumed to be proportional to that of the gasphase, and all condensed-phase products are assumed to be captured in thefilter screen between the plenum and discharge tank.

The governing equations presented here were solved numerically for aseries of test cases to demonstrate the effect of pre-pressurized gas on theperformance of the inflator. The results are highlighted in the followingsections. The computer program developed for this analysis incorporates a

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316 R. G. SCHMITT er 01.

detailed thermochemical package (Kee, 1980) to evaluate thermodynamicproperties and gas-phase, finite-rate kinetics for some species. A stiff ordi­nary differential equation solver (Hindmarsh, 1983) was used to solve thegoverning equations.

PROPELLANT EQUILIBRIUM ANALYSES

Table II summarizes the propellant formulation used in all results presentedin this study. Using this formulation, a series of chemical equilibrium calcu­lations were made using PEP (Cruise, 1973) to assess the variability in flametemperature and chemical composition of the propellant's products over arange of typical operating conditions. Figure 3 shows the calculatedadiabatic flame temperature as a function of pressure for cold (To= 245 K),ambient (To = 300 K), and hot (To= 325 K) propellant temperatures. Therange in ambient temperature examined here (t. To = 80 K) is consistent withindustry requirements. The results shown in Figure 3 demonstrate thechange in flame temperature is relatively small over the range of pressurestypically encountered during the transient operation of an inflator unit(0.1 < P(MPa) <25). For example, Figure 3 shows the change in calculatedflame temperature is approximately 30 K over the range in combustionpressure of 10 to 20 MPa. Figure 3 also illustrates the change in flametemperature is relatively small when the initial propellant temperature isvaried for a fixed operating pressure.

Figure 4 shows the variation of equilibrium product mass fractions as afunction of temperature for a typical inflator operating pressure (P = 18M Pal. At this pressure the adiabatic flame temperature is approximately2,400 K (see, Fig. 3 for To = 300 K). The equilibrium products of combus­tion at this state (P = 18 MPa, T = adiabatic flame temperature) werechosen as the species produced (Yk.g,n, Eq. 10) by the surface reaction of thepropellant for the entire ballistic period. Consequently, the results inFigure 4 show the equilibrium product composition as the products arecooled from the flame temperature to 1,800 K, values typical of an inflator

TABLE II Propellant Composition

Mass Component W p L1H'98(%j (glmol) (glcm' ) (kJIKmol)

40.5 C,H.N.O, (ADCA) 116.08 1.651 -292,59649.5 KCIO. 138.54 2.510 -430,74510.0 CuO 79.54 6.400 -146,306

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NON-AZIDE AIRBAG PROPELLANT

2450

2400.....:lid......... 2350=.........l:l.8 2300..~..8 2250~

2200

21500 5 10 15 20 25

Pressure [MPa]

317

. FIGURE 3 Adiabatic flame temperature predicted for ADCA propellant (Tab. II) at hot(325 Kl, ambient (300 K), and cold (245 K) test conditions.

combustion chamber. Figure 4 illustrates that most of the major equilib­rium product species (C0 2, N 2, H20, O 2 and CuO(L)) are relatively inde­pendent of temperature. However, the major product species KCI andK2C1 2, and the minor species CuCl, KOH, CO and Cu a.e dependent onthe temperature. The variation of the equilibrium mixture mass fractions asa function of pressure for a typical operating temperature of 1,800 K isshown in Figure 5. This graph also illustrates that most of the major equi­librium product species (C02, N 2, H20, O 2 and CuO(L)) are independentof the operating pressures greater than approximately 2 MPa. However,once again the major product species KCl and K 2C1 2, and minor speciesCuCl, KOH, CO and Cu show a pressure dependency.

Conclusions drawn from the equilibrium calculations presented inFigures 4 and 5 are as follows: i) equilibrium mass fractions of the majorproducts of combustion vary little over a wide range of combustion cham­ber pressure and temperature, ii) equilibrium mass fractions of potassiumand chlorine species show significant variations with pressure and tempera­ture, and iii) approximately 3,000 ppm of CO is produced at the adiabaticflame state due to the high flame temperature of the propellant. Thus, in

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0.35 0.022

0.0200.30

0.0181:1 1:1

~ 0.25 0.016 ~... ...III 0.014 III.. ..'"' '"''"

0.200.012 '"1a 1a

~ 0.010 ~.. 0.15 .... ..:I 0.008 :I- -~ 0.10 0.006 ~

0.0040.05

0.002

0.00 0.0001800 1900 2000 2100 2200 2300 2400

Temperature [K]

FIGURE 4 Equilibrium product mass functions for ADCA propellant (Tab. II) at P = 18 MPaand range of temperatures.

0.35 0.040Tempemture =1800 K

-CO2 0.Q350.30

1:1 0.030 1:1:e 0.25 0;:...

III0.025 III.. ..

'"' 0.20 '"''" '"1a '"0.020 III~ ~

t0.15 _1120 ..

I 0.QI5 ..:I :I- -~ 0.10

.c

0.010 ~

0.05 0.005

0.00 0.0000 5 10 15 20 25

Pressure [MPa]

FIGURE 5 Equilibrium product mass functions for ADCA propellant (Tab, II) at T= 1,800 Kand range of pressures.

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NON-AZIDE AIRBAG PROPELLANT 319

attempting to model the production, condensed-phase filtering, and subsequentdischarge of the propellant products of combustion, the major products areassumed to be chemically frozen at the adiabatic flame state composition. Thechlorine-potassium kinetics are also assumed frozen in the present study. Topredict the CO concentration in the airbag for this inflator requires a finite-ratechemical kinetic reaction scheme. This scheme models the gas-phase chemistryas the temperature and pressure evolve during the operation of the inflatorunit. Forward rate constants for the CO elementary chemical kinetic modelused in this study are presented in Table III (Alkam, 1996). The reverse reac­tion rates k'i are computed using the concentration-based equilibrium con­stants and the forward rates. In addition to finite-rate chemistry, two otherlimiting conditions can be modeled using this reaction scheme. The first, a

TABLE III Gas-phase Kinetics Model [Alkam et al., 1996]

Gas Phase Reaction Ai {Ji E,/R

H+O, .... OH+O 1.91 x 1014 0 8.27 X 103

H,+O .... H+OH 5.13x 104 2.7 3.16xI0'OH+H, .... H+H'O 2.14x108 1.5 1.73x 103

20H -ee-O+H,O 5.63x 10" 0.3 7.17x1O'H,+M .... 2H+M 8.51 x 1010 -1.1 5.25x 104

H+OH+M .... H,O+M 1.38x 10" -2 0.00 x 10°°H+O, .... HO, 4.79x 1013 0 -1.94 x 10'H+HO, .... 20H 1.70x 1014 0 4.38 X 10'H+HO, -ee- H,+O, 6.61 x 10" 0 1.07X 10'HO,+O .... O,+OH 1.74x 10" 0 -2.01 X 10'OH+HO, .... H,O+O, 1.45 x 1Q16 -I 0.00 x 10°H,O,+OH .... H,O+HO, 7.08x 1012 0 7.19X 10'2HO, .... H,O,+O, 8.51 x 101' 0 2.12x 103

20H .... H,O, 7.59X 10" -0.4 0.00 x 10°H,O,+H .... HO,+H, 4.79x 1013 0 1.81 X 103

H,O,+H .... H,O+OH 1.00x 1013 0 1.81 X 103O+H+M .... OH+M 4.68x 1018 -I 0.00 x 10°20+M .... O,+M 3.98x 1014 0 -9.01 X 10'H,O,+O .... OH+HO, 9.55 x 106 2 2.00 X 103

HCO+OH .... CO+H,O 3.02x 10" 0 0.00 x 10°OH+CO .... H+CO, 2.85x 10' 1.3 -3.85 X 10'HCO+M .... H+CO+M 3.47x 10'8 -I 8.56 x 103

CO+HO, .... CO,+OH 6.03x 10" 0 1.16X 104

CO+O .... cO, 1.82x 10'° 0 1.22X 103

CO+O, .... CO,+O 2.51x 1012 0 2.40 X 10'HCO+H -ee-CO+H, 7.24x 10" 0 0.00 x 10°O+HCO ee- CO+OH 3.02x 10" 0 0.00x 10°HCO+O, ee- CO+HO, 7.59x 1012 0 2.07X 10'O+N, .... NO+N 1.84x 1014 0 3.84 X 104

N+O, -ee- NO+O 6.40x 100 I 3.16x103

H+NO .... N+OH 2.22x 1014 0 2.54 X 104

N,+M .... 2N+M 3.71 x 1021 -1.6 1.13x 10'CO+O+M .... CO,+M 3.20x 10" 0 -2.11 X 103

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320 R. G. SCHMITT er aJ.

shifting equilibrium analysis, is obtained by artificially increasing the reactionrates by several orders of magnitude. This effectively forces the species concen­trations to their equilibrium values at the local (P,T) state. The second, frozenchemistry, is obtained by setting the gas-phase reaction rates to zero. Thus,with the frozen chemistry assumption the only factor influencing change inchemical composition after production at the propellant surface is mixing withthe plenum and tank gases.

RESULTS-INFLATOR SIMULATION

The results for three different simulations of an augmented inflator unit arediscussed in this section. The input parameters used in these simulationsare summarized in Table IV. In each case the plenum was pressurized to6.8 MPa with argon, the combustion chamber was initialized to 0.1 MPa withargon, and the discharge tank was initialized to 0.1 MPa with nitrogen. Thethree cases analyzed in the following sections consist of: i) a baseline casewith frozen gas-phase chemistry (i.e., Case I), ii) a repeat of Case I withequilibrium CO chemistry (i.e., Case 2), iii) and a repeat of Case I withfinite-rate CO chemistry (i.e., Case 3).

Pressures in the combustion chamber, high-pressure plenum, and dis­charge tank for the baseline numerical simulation (Case I) are illustrated inFigure 6. Here, 0 ms corresponds to the time when the ignitor fires. Afterpropellant ignition occurs 0.2 ms later, the pressure in the combustion

Alii teT

Anozzlc

Inpmigna[]"oTo'1Itiltctr;J-;po.PVp

PO,I

V.x;.. (N,)X o.P (Ar)X o., (Ar)

TABLE IV Model Parameters

165.28 em'0.0962 em'0.025 kg0.0005 kg0.03340 mm/(s- Pa")0.595350.0298 K0.055 kg0.101 MPa0.085 L6.86 MPa0.060 L0.101 MPa30 L1.01.01.0

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NON-AZIDE AIRBAG PROPELLANT

20.,.------------__....,.. 0.2

a::s=..

...... E: 15 0.15ClI '"

~~~ol:l.... ..::s ..

10 0.1"'.0~ E.. ClIll..c.. uQI)=i Qe,,: 0.05

'"::s.0aQ

oo

10 20 30 40 50 60 10 80 90 100

Time [ms]

321

FIGURE 6 Combustion chamber. gas plenum. and discharge tank pressure profiles. Profilesare identical for Cases 1,2 and 3.

chamber rises rapidly 10 the burst pressure of the foil separating it from thegas plenum. The burst foil between the combustion chamber and plenumruptures at a prescribed pressure approximately 4 ms causing the plenumpressure to rapidly increase from its initial value of 6.8 MPa. The burst foilbetween the plenum and discharge tank also ruptures at a prescribed press­ure at approximately 5 ms, and the discharge tank begins to pressurize. Thepressure in the combustion chamber is unaffected by the rupturing of theburst foil between the plenum and discharge tank (i.e., exit flow remainschoked), whereas the high-pressure plenum shows a slight drop in pressure.The pressure drops because the exit flow exceeds the inflow coming fromthe combustion chamber. Continued mass generation from propellant de­composition causes the combustion chamber and plenum pressures to con­tinue to increase until they approach maximum pressures at approximately20 ms and 24 ms, respectively. The pressure in each chamber decreasesthereafter and at approximately 94 ms, the propellant burns out causing thepressure in both the combustion chamber and plenum to decrease in amanner similar to a simple polytropic expansion process. During the in­flator discharge, the tank pressure increases to a maximum pressure of

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322 R. G. SCHMITT er al.

approximately 0.18 M Pa gauge at approximately 83 ms. While the resultspresented here are for only one set of initial conditions, these profiles aretypical of the tank-test firing of a driver or passenger-side airbag inflatorunit.

The temperature of the combustion chamber gas, plenum gas, tank gas,condensed-phase products captured by the filter screen, and the filter screenare shown for Case I as a function of time in Figure 7. The hot product gasesfrom the igniter and propellant cause a rapid increase in the combustionchamber gas-phase temperature. During this time, the combustion chambertemperature increases to 2,747 K, approximately 12% above the adiabaticflame temperature. This temperature is representative of the constant-volumecombustion process occurring in the combustion chamber. When the foilbetween the combustion chamber and plenum bursts open at 2 ms, thetemperature in the combustion chamber begins to decrease towards theadiabatic flame temperature. Additional heat losses to the combustion cham­ber wall forces the temperature to continue to drop as time evolves. Thetemperature of the plenum gas during this time period increases due tomixing of the hot combustion gas with the pre-pressurized argon. The tem­perature of the plenum gas is also influenced by several other factors which

PropeUanI

~bums out '~'..

:lid 2000~.....:I- 1500ell.....g"

a.. 1000Eo<

500

o+-......-..,..-.......-..,..-......--r-.......--r-......,...&,..~o W w m ~ ~ ro m W 00 100

Time [ms]

FIGURE 7 Temperature profiles for various components of inflator system (Case 1).

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NON-AZIDE AIRBAG PROPELLANT 323

include: interaction with hot, condensed-phase products, interaction with coolfilter screen, and heat transfer to the chamber walls.

For this propellant formulation, CuO(L) is the sole condensed-phaseproduct of the propellant. It is initially formed in the liquid phase (2,445 K)in the combustion chamber, passes through the high-pressure plenum and iscollected by the filter screen. The CuO particles transfer heat by convectionto the plenum gas and by conduction to the filter screen after they arecaptured. Throughout the process, the temperature of the filter screen in­creases from ambient towards an equilibrium temperature between thescreen, gas and condensed-phase products. As the liquid copper oxide,CuO(L), is cooled in the filter it eventually reaches a point at which itundergoes a change of phase to solid copper oxide, CuO(S). During thesolidification process, the temperature of the copper oxide remains constantat the melt temperature. The solidification process of the CuO(L) occursbetween 18 ms and 40 ms. After solidification is complete, t > 40 ms, thetemperature of the solid copper oxide decreases toward that of the plenumgas and the filter screen. At approximately 94 ms the propellant burns outand the temperature of the combustion chamber and plenum gases begin todecrease in a polytropic manner. The temperature of the gas within thedischarge tank slowly approaches a maximum at 66 ms and then begin toslowly decrease due to heat transfer to the tank walls.

Figure 8 shows the propellant mass production rate and mass flow rateinto the discharge tank as a function of time. The propellant used for thistest case is comprised of 25 g with a cylindrical grain geometry (d = 12.8mm and h = 5.44 mm). The propellant ignites at 0.2 ms and the massproduction rate rises sharply to a maximum point of approximately 550 g/s.Mass flow from the plenum into the discharge tank begins when the plenumnozzles rupture at 4 ms. The mass flow rate at this time rises rapidly to apeak of 500 glsand then decreases gradually. This initial burst of mass fromthe plenum provides the necessary impulse required for deployment of theairbag. There is a slight influence of the propellant burn-out at 94 ms on theflow rate into the discharge tank.

In the baseline simulation (Case I), the propellant products of combus­tion were assumed to be chemically "frozen" at their adiabatic flame tem­perature equilibrium composition and allowed to mix with the plenum andtank gas as they exited the combustion chamber, but were restricted fromundergoing reactions in the gas phase. In an actual inflator system, how­ever, the product composition can change as the gas and solids work theirway through the various filters and chambers and eventually exit the in­flator. The initial conditions for Case 2 are identical to Case 1, except the

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324 R.G. SCHMITT et al.

600

500

~

~400.

'".....~ 300

~ propellant ":!1 200 mass genemIion "

100

00 10 20 30 40 50 60 70 80 90 100

Time [ms]

FIGURE 8 Propellant product mass generation and mass flow rate into discharge tank(identical [or Cases 1,2, and 3).

model was extended to include very fast (i.e., equilibrium) CO gas-phasekinetics. Due to the small proportion of CO produced by the propellant,this change to the model has a minimal influence on the pressure andtemperature. Consequently, Figures 6-8 used to describe Case I can also beused to accurately describe Case 2. The only observable difference is in thecombustion chamber temperature. During the first 5 ms, the combustionchamber gas temperature is lower than the "frozen" calculation. This is theexpected result because the CO2 contained in the combustion productsdissociates at temperatures exceeding the adiabatic flame temperature. Thedissociation absorbs energy resulting in lower temperatures. After 5 ms, theCO present in the combustion chamber undergoes recombination reactions.The exothermicity of the recombination reaction results in the slightly high­er combustion chamber gas temperature for Case 2. As expected, Case 3(finite rate CO chemistry) also showed nearly identical pressure and tem­perature profiles to Case 2.

While the pressure profiles for Cases I, 2 and 3 are nearly identical, theassumption of frozen chemistry in Case I results in the model predicting ahigher level of CO present in the tank at any time. In Case I, the model

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NON-AZIDE AIRBAG PROPELLANT 325

predicted a CO mass fraction of 1,280 ppm at 100 ms. The results of Case 1can be treated as a theoretical limit of infinitesimally slow CO chemistry,On'the other hand, Case 2 can be considered as the other limit of infinitelyfast CO chemistry. Figure 9 shows the CO concentration in the tank forCases 1 and 3. Case 2 is not shown because the relatively low temperaturein the tank (see Fig. 7) drives the equilibrium quantity of CO to nearly zero.The reduction in the CO mass fraction to approximately 5 ppm is account­ed for by the recombination reaction for CO as it cools in the plenum. Thisis illustrated in Figure 10, a plot of the CO destruction rates in each of thethree chambers for r> 5 ms. The CO production/destruction rate in thecombustion chamber is illustrated in Figure II. The initial production ofCO at time < 5 ms is associated with CO 2 dissociation and the CO destruc­tion after 5 ms is associated with recombination reactions as the gas cools.As discussed previously, the CO production/destruction is responsible forthe temperature differences between the finite-rate kinetic and the frozencalculations. Table V summarizes the key performance parameters for thethree cases. While the final tank CO levels are considerably different for thethree cases, Table V shows the three predict nearly identical performance in

tOOOO

......6Cl. tOOOCl.~

c.:o:o c•• <II1;jE-<.. .."'OilC .. 100~ <IIc.co ...U·!!l.,Q[(l c:; ..

to• Fini~ Rale

0 ,U

t0 to 20 30 40 50 60 70 80 90 100

Time [ms]

FIGURE 9 CO mass concentration in discharge tank for finite-rate kinetics case (Case 3) andchemically frozen case (Case I).

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10..---- ----.

..... 0.1~CI0 0.01~:s:e~t\ 0 0.001... uCiS~ .

0.0001 l>i.o:harBe TlIIlk:s0u

IE-05

IE-060 10 20 30 40 50 60 70 80 90 100

Time [ms)

FIGURE 10 Mass destruction rate of CO in each chamber for Case 3 resulting from finite­rate reactions in gas phase.

IO~--------------__,

8

6

4

2-

°tT---==::::::====="i-2-

I I' I I I

10 W ~ 40 ~ 60 m W 90 ~

Time [ms)

·6+-~.................~......~.......~.................~.......~......._.._~o

FIGURE II Mass production rate of CO in combustion chamber for Case 3 resulting fromreactions in gas phase. .

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NON-AZIDE AIRBAG PROPELLANT 327

TABLE V Summary of Performance Parameters for 3 Test Cases

Case CO Kinetics Pc,m31 P"mlll 1'c.mal 'T,.mu 10 - 10 0 Yeo.... 100

# Model Used [MPa] [MPa] [K] [K] [MPa-ms] [ppm]

I Frozen 19.4 0.18 2854 536 1726 12502 Finite-Rate 19.6 0.18 2747 540 1748 4.393 Equilibrium 19.6 0.18 2747 540 1747 «1

the following categories: maximum combustion chamber pressure P,,,,,,,maximum tank pressure p,.m" combustion chamber temperature I;.m,,'maximum tank temperature I;.max' and tank impulse over the periodo< t < 100 ms, I, 00'

Previous experiments (Hara, 1994) used a 10 g sample of a similarADCA-based propellant. The products of combustion were discharged intoa 7,500 cc tank initially filled with either air or pure N 2' When no coolingfilter was used their results showed CO levels of 400 ppm in the air tankand 950 ppm in the N 2 tests. When the products were. cooled in a filterbefore entering the tank, the levels dropped to 3.3 and 3.7, respectively.When their results are scaled to the operating conditions reported herein(25 g of propellant and 30,000 cc tank), the CO levels are 2.6 ppm and 2.9ppm for the air and N 2 environments, respectively. These results cannot becompared directly with the present simulations because of incomplete infor­mation about experimental test conditions. However, these experimentsdemonstrated the effects of I) initial tank oxygen levels and 2) the dischargetemperature on CO levels, thus emphasizing the importance of includingfinite-rate chemistry in the simulations.

CONCLUSIONS

Results presented in this paper demonstrate the necessity of including finite­rate gas-phase kinetics in any computer program used to model an airbaginflator combustion. The program used here was developed as a researchtool to study the role of pyrotechnic and gas properties on the performanceof an automotive airbag inflator, specifically an augmented inflator thatburns a high CO-producing propellant. Properties of the pre-pressurizedinert gas system of interest include: specific heat of the gas, initial pressure,burning-rate, grain geometries, hardware configurations, propellants, heattransfer characteristics, and initial temperature. Performance of the inflatorwas presented in terms of pressure, temperature and mass flow rate profiles

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328 R. G. SCHMITT et al.

in the inflator and discharge tank. The pre-pressuized pyrotechnic inflatorhas certain advantages over conventional pyrotechnic units, including sig­nificantly lower requirements for solid propellant mass, lower operatingtemperature, more uniform performance at hot and .cold ambient condi­tions, and higher thermal efficiency. Less propellant mass and lower operat­ing temperatures reduce the amount of unwanted chemical species, such asCO and NO x present in the inflator products. Results from the finite-rateCO chemistry model are qualitatively consistent with experimental resultsreported by others for the same propellant formulation in a similar opera­ting environment.

LIST OF SYMBOLS

Variable Definition Units (cgs)

'A area cm

Ai pre-exponent factor in gas-phase reaction depends on reactiona propellant burning rate prefactor cm/(s-MPa")

Cd discharge coefficientCp constant pressure specific heat ergs/g-KC, constant volume specific heat ergs/g-Kd diameter of filter wires cmE activation energy ergs/gH enthalpy ergs/gI impulseII number of reactionsKK number of speciesk reaction rate constant depends on reactionIII mass gm mass production rate g/sN number of propellant grainsp pressure dynes/em?Q energy transfer ergsr burn depth cmi propellant burning rate cm/sR gas constant ergs/rnole-KT temperature KU internal energy ergs/g

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NON-AZIDE AIRBAG PROPELLANT 329

V volume cm 3

W molecular weight gjmole[Xk ] concentration of k-th species molejccy mass fraction

Greek

/iH298 heat or formation at T= 298 K ergsjg

X chemical symboly specific heat ratioy' reactant stoichiometric coefficienty" product stoichiometric coefficientp density gjccOJ gas-phase mass production gjsljl volume fraction

Subscripts

c combustion chamber

f forwardfilter filter screengen generant (propellant)

reaction indexign ignitork species index0 initialout exiting a computational cellp plenumr reverset tank298 reference temperature

Superscripts

f3 gas-phase reaction-rate temperatureexponent

a propellant temperature sensitivitycoefficient

n propellant burning-rate pressure indexgas gas phase

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330

References

R. G. SCHMITI et 01.

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Berger, J. M. and Butler, P. B. (1995) Equilibrium Analysis of Three Classes of AutomotiveAirbag Inflator Propellants. Combustion Science and Technology, 104(1-3), pp. 93-114.

Butler, P. 8., Kang, J. and Krier, H. (1993) Modeling and Numerical Simulation of the Inter­nal Thermochemistry of an Automotive Airbag Inflator, Progress in Energy and Combus­tion Science, 19, pp. 365-382.

Butler, P. B., Kang, J. and Krier, H. (1993) Modeling Pyrotechnic Combustion in an Automo­tive Airbag Inflator. 5th International Congress of the Groupe de Travail de Pyrotechnie,June, France.

Butler, P. 8., Kang, J. and Krier, H. (1993) Numerical Simulation of a Pre-Pressurized Pyro­technic Automotive Airbag Inflator. 5th International Congress of the Groupe de Travail dePyrotechnie, June, France.

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Hara, K., Misawa, T, Maekawa, T., Iyoshi, S., Yokoyama, T., Kazumi, T, Hayashi, M. andYoshida, T. (1994) Concept and Performance of a Non-Azide Propellant For AutomotiveAirbag Inflators. Proceedings of 19th International Pyrotechnics Seminar, Christchurch,New Zealand, 20-25 Feb.

Hara, K., Shimizu, Y., Higashi. K.• Nojima, T., Hasegawa, T.and Yoshida, T. (1994) HazardAnalysis of Manufacture of a Non-Azide Propellant For Automotive Airbag Inflators.Proceedings of 19th International Pyrotechnics Seminar, Christchurch, New Zealand,20-25 Feb.

Hindmarsh, A. C. (1983) ODEPACK: A Systemized Collection of ODE Solvers. ScientificComputing, edited by R. S. Stepleman, M. Carver, R. Peskin, W. F. Ames andR. Vichnevestsky, North-Holland. Amsterdam, pp. 55-64.

Kee, R. J., Miller. J. A. and Jefferson, T. H. (1980) CHEMKIN: A General-Purpose, Problem­Independent, Transportable, FORTRAN Chemical Kinetics Code Package. Sandia Na­tional Laboratories, SAND 80-8003. Livermore, CA.

Kuo, K. K. and Summerfield, M. (1984) Fundamentals of Solid-Propellant Combustion (Prog­ress is Astronautics and Aeronautics), A1AA, 90, pp. 622-623.

Vos, T. H. and Goetz, G. W. (1989) Inflatable Restraint Systems: Helping Save Lives on theRoad TR W Space and Defense Quest. Winter Issue, pp. 1-6.

Zucrow, M. J. and Hoffman, J. D. (1976) Gas Dynamics, 1, John Wiley and Sons.

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