Billing Supersonic Combustion

JOURNAL OF PROPULSION AND POWERVol. 9, No. 4, July-Aug. 1993

DRYDEN LECTURESHIP ON RESEARCH

Research on Supersonic Combustion

F. S. BilligJohns Hopkins University, Laurel, Maryland 20723

Introduction

S OME qualifications to the pretentious nature of the titleof this Paper are in order. Research in this context is

applied, i.e., it is directed toward the design and developmentof devices. The devices are air breathing propulsion systems,wherein supersonic combustion is inherent or it provides anadjunct benefit. Principal applications of the former are mannedhypersonic aircraft, transatmospheric accelerators, and mis-siles. Typical examples of the latter are external burning sys-tems that sustain thrust or reduce drag when used in tandemwith primary accelerator engines.

This Paper is not at all representative of the complete bodyof the research on supersonic combustion. Instead, what ispresented is primarily based on material with which the writerhas had either direct involvement or a first-hand knowledge.It does not do justice to the exemplary works of many otherinvestigators. Hopefully, this is somewhat rectified by theextensive list of references. The reader is encouraged to con-sult these manuscripts to obtain a more balanced perspective.

This applied research must be viewed from the perspectiveof the engineer who is charged with expediting developmentand is in the need of design tools. Exact physics have fre-quently been abandoned to produce functional models. Muchof the experimental verification upon which the models arebased has been obtained in facilities which have known im-perfections with respect to duplication of flight conditions.Moreover, compromises have been made in the analysis andinterpretation of the experimental data, in part due to im-perfect or incomplete diagnostic instrumentation, and in partdue to a limited understanding of the underlying physics. Anexample that embodies all of these deficiencies would be adesign model that is based on the use of integral techniquesto describe combustion in supersonic flow from experimentsconducted in an arc-heated tunnel.

BeginningWhen a source term for heat release is included in the

energy equation and, in turn, solved simultaneously with theremaining conservation equations, the stage for supersoniccombustion is set. Tsien and Beilock1 and others,2 havingposed these equations, obtained solutions for simple diabaticflows. The mathematics certainly did not preclude supersonicinitial conditions. Indeed, Pinkel and Serafini3'4 extended themethod of characteristics to include the heat source term inirrotational supersonic flow and developed a graphical solu-

tion for shock-free flow. Having established a theoreticalfoundation, the paramount question became, can combustionbe established in steady supersonic flow? Skeptics scoffed atthe possibility. "After all, flame speeds are known to be lessthan a few hundred feet per second, even for hydrogen-airmixtures."

Was there any solid evidence to counter the mid!950s skep-tic? It was known that the speed of a traveling detonationwas supersonic, relative to a stationary observer. However,a close inspection of the wave structure reveals that a strongshock precedes the heat release, thus, the initial conditionsin the heat release structure of the wave front are subsonicand the final conditions are sonic, as explained by Chapman5

and Jouget.6 Tracer bullets fired at supersonic speeds fromaircraft had been used extensively in World War II. The de-sired luminosity of the combustion of the solid pyrotechnicattached to the base was readily observable, but, was any ofthe flow in the luminous zone supersonic? (Interestingly, the"economy" intended through selectively "fueling" only a smallfraction of a high-speed round was difficult to realize. Thecombustion reduced the base drag, thus altering the trajectoryand segregating the tracer from the principal mass of the burst.It required a skilled marksman to compensate for the offsetand obtain the desired lethality.) Following the war, Bakeret al.7 began to investigate wake combustion by injecting hy-drogen into the base of a 2Hn.-diam cone cylinder placed ina Mach 1.6 free jet. Similar tests were made by Scanland andHebrank8 who burned a solid propellant composition in thebase of 40-mm projectiles. Base pressure rise due to heatrelease was significant in both test series, thereby substanti-ating the heuristic arguments that had been made to explainthe alteration of the trajectory of the tracer bullets. Unfor-tunately, measurements to ascertain whether any part of theflame zone was supersonic were not made. The results ofDavis,9 who injected hydrogen on the base of a rearward-facing step on a flat plate at M0 = 1.7, were also inconclusive.He needed to add large amounts of oxygen and/or flameholders to stabilize the combustion, which may have beenconfined to the subsonic wake.

The pioneering experiments of Dorsch et al.,10-15 at theNACA Lewis Laboratories, laid to rest any lingering doubtsof the veracity of combustion in supersonic flow. They hadreasoned that highly reactive fuels would be needed to obtainheat release in the available residence time. Aluminum bo-rohydride was selected for most of their work. Stable com-

Dr. Billig is the Associate Supervisor and Chief Scientist of the Aeronautics Department of the Johns HopkinsUniversity Applied Physics Laboratory (APL). In 1991, Dr. Billig received the lifetime achievement award fromAPL in recognition of his pioneering analytical and experimental contributions to the understanding and de-velopment of supersonic combustion ramjet engines, his leadership on behalf of national and internationalprofessional societies, and his decades of dedication to the education and mentoring of current and futureaerospace engineers. He is a fellow, past Vice President and Director of AIAA and recipient of the 1992 DrydenLectureship on Research presented herein.

Presented as the Dryden Lectureship on Research (Paper 92-0001) at the AIAA 30th Aerospace Sciences Meeting, Jan. 6-9, 1992, Reno, NV;received April 8, 1992; revision received Nov. 1, 1992; accepted for publication Nov. 13, 1992. Copyright © 1993 by the American Institute ofAeronautics and Astronautics, Inc. Under the copyright claimed herein, the U.S. Government has a royalty-free license to exercise all rights forGovernmental purpose. JHU/APL reserves all proprietary rights other than copyright; the author(s) retain the right of use in future works oftheir own; and JHU/APL reserves the right to make copies for its own use, but not for sale. All other rights are reserved by the copyrightowner.

499

500 BILLIG: RESEARCH ON SUPERSONIC COMBUSTION

bustion, without resort to the use of physical flame holders,was demonstrated adjacent to the external surfaces on a va-riety of models at Mach numbers of 1.5-4.0. Computed Machnumbers based on static and pitot pressures taken within theluminous flame zone showed that the Mach number was sub-sonic through most of the flame zone, and became sonic andlow supersonic as the hot-cold interface was approached.

These exciting results prompted the two leading institutionsin ramjet development, the Johns Hopkins University Ap-plied Physics Laboratory (JHU/APL) and the MarquardtCompany, to focus attention on ramjet cycles using supersoniccombustion. Nearly all of the significant research in the firstseveral years by these organizations was not made availableto the general technical community due to security classifi-cation. In recent years, permission has been granted for re-lease of the information but, unfortunately, little has beenrepublished. Reference 16 presents an excellent summary ofthe early work at Marquardt. This Paper will highlight thework at JHU/APL.

Three outstanding classified conferences on the researchand application of supersonic combustion were held in theperiod 1959-1964. References 17-19 identify the papers thatwere presented, most of which should now be available. Manyother authors and organizations20"47 were also contributing tothe growing body of information on the subject.

ApplicationsA requisite for guidance to provide focus for applied re-

search is an appreciation of the probable applications thatcould exploit the results. Interestingly, the more viable con-cepts for the application of supersonic combustion in contem-porary air breathing systems were contemplated by the pi-oneers in the mid-to-late 1950s. Consequently, it is appropriateto frame the discussion of the research in terms of their clair-voyant concepts. The intervening years have only providedminor modifications to the original concepts.

Applications for supersonic combustion group are dividedinto four major categories: 1) external burning devices forthrust production (or drag reduction) and/or lateral control;2) primary propulsion for missiles; 3) primary propulsion forhypersonic airplanes and transatmospheric accelerators; and4) thrust augmentation for fuel-rich rockets.

Figure I9 is a composite of the concepts for external burning(EB), wherein the flow in the combustion zone is supersonic.As shown, the flame zones are diffusive, whereas the originalsketches of EB combustion zones48'49 depicted a thin planarflame. The planar flame concept was a convenience to ex-pedite simple calculations to obtain estimates of performance.From a practical perspective, it is quite doubtful that fuelcould be injected in such a manner that it would mix with theair, but not ignite until reaching a prescribed plane, and theninstantaneously react. Nonetheless, calculations of the re-quired flame height in terms of model chord and the deter-

c) Thrust and lift generator Axisymmetric Two-dimensionalFront views

Fig. 1 External burning configurations.

mination of engine performance as measured by specific im-pulse were not too different from those that would be obtainedfrom models based on diffusion flames. The subsequent dis-cussion will develop the arguments which show that for thesame amount of heat release, performance is primarily de-pendent on the airflow conditions entering the combustionzone, and is only weakly affected by the character of thecombustion process.

In effect, EB represents a volume source which causesstreamlines about the body to be deflected, giving a pressurerise similar to an aerodynamic flap, but with a significantlylesser expansion effect and no drag penalty. There is also areaction force caused by the injection, which has componentsin the thrust and/or lateral directions, depending on the angleof injection. The attitude controller for an axisymmetric ve-hicle (Fig. la) has injection aft of the center-of-gravity (e.g.)in any one of four quadrants. (The periphery of the vehiclecould alternatively be subdivided into any number of desiredsegments.) Longitudinal "fences" separate the quadrants toreduce the dissipation of the positive pressure field throughcircumferential spillover. The downward force due to EB leadsto positive pitch (a), and therefore, puts the EB in the leewardzone, which could, at large a, produce adverse conditions forcombustion. However, if EB is being used solely to trim thebody, then an aerodynamically unstable vehicle could be de-signed, in which case the EB will always occur in the windwardzone. Attitude control systems based on EB ahead of the e.g.are conceivable, but appear to be less attractive due to thedifficulty of confining the positive pressure field to producean effective pitching moment. The thrust generating device(Fig. Ib) could be either the total vehicle or a podded orairfoil engine. At the "knee," fuel is added to the air andcombustion maintains a positive pressure field on the aft bodywhich is greater than that on the compression surface, thusproducing net thrust. It turns out, however, that the specificimpulse decreases drastically as the EB changes from a drag-reducing to a thrust-producing device. Consequently, EB isimpractical as an accelerator, and in plausible applicationsmust be used in tandem with a primary thrust-producing de-vice. Figure Ic bifurcates the vehicles shown in Fig. Ib, whichtherefore provides a combined "axial" and lateral force ca-pability. Here, EB can provide its maximum potential. Thelateral force generated by combustion negates the need fordeflection of aerodynamic surfaces, and thereby eliminatesinduced drag with a corresponding reduction in engine fuelflow. The axial component of force due to the rise in pressureeither reduces or cancels drag, and if sufficiently high levelscan be obtained, then net thrust is produced.

It is advantageous to establish external combustion im-mediately following the forebody compression to enhance ig-nition in the zone of higher static temperature and pressure.Even with this assist, it is far more difficult to burn externallythan within the confines of an internally ducted engine. Notonly are the pressures and temperatures considerably lowerin the external burner, but the residence time for completionof heat release is shorter. Moreover, if flame stabilizationdevices are needed, the drag penalties are much greater dueto the higher dynamic pressure.

Figure 2 is taken from U.S. patent 4,291,533 awarded toG. L. Dugger and the author in 1981, following release froma long-standing "order of secrecy." This concept for a super-sonic ramjet missile had been disclosed in 1959 and was re-ported in Ref. 50. Many interesting features are embodied inthe concept. The inlet innerbody is comprised of three sec-tions. The forward and aft sections translate, thereby adjust-ing the internal shock structure to obtain wave cancellationat the corners. The objective is to minimize internal lossesand produce a tailored inlet compression ratio to maximizeengine performance at all flight speeds. Fuel injectors arelocated at several axial stations to provide the variable arearatio combustion process that is desired in a scram jet engine.Many of the potential applications of vehicles powered by

BILLIG: RESEARCH ON SUPERSONIC COMBUSTION 501

Augmented rocket contour

Fig. 2 Scramjet missile, U.S. patent 4,291,533, Dugger and Billig.

supersonic combustion ramjets were posed in this patent ap-plication.

Theoretical studies have shown the possibility of pro-pelling ramjet vehicles with hydrogen fuel to orbitalspeeds and with storable fuels to Mach 15 or above. . . .The margin of superiority of ramjet missiles to rocketsincreases rapidly as missile speed increases. To provideequivalent performance to a Mach 7 supersonic com-bustion ramjet vehicle at sea level, a rocket would haveto have roughly three times as much weight. As animportant object, therefore, the present invention pro-vides a missile, comparable with length and volumerequirements of the Terrier rocket missile system, whichwould accelerate from Mach 4.0 at an average of 32 g'sto a cruise speed of between Mach 6.5 (7257 ft/s) andMach 7 (7810 ft/s) at sea level, and at speeds betweenMach 8.5 (8296 ft/s) and Mach 10 (9676 ft/s) at altitude.

Figure 351 is one of the original concepts for augmentationof the thrust of a rocket with afterburning of the fuel-richexhaust. Air, captured in multishock inlets, is ducted into thesupersonic portion of the rocket nozzle and establishes asupersonic mixing and combustion zone. The higher pressuregenerated in the expansion nozzle more than compensates forthe drag of the air inlets, and thereby increases the net thrustof the entire vehicle. As initially proposed, the thrust aug-mentation would more than double the range of the solid-fueled Polaris missile. The concept was demonstrated in aliquid bipropellant (UDMH-N2O4) system at JHU/APL in1963.

The fourth application is the transatmospheric accelerator,remarkably similar in concept to that of the National Aero-space Plane (NASP). In this 1959 design study,18 the questionregarding the selection of a low-speed (Mach 0-3) propulsionsystem was begged. The engine operated as a subsonic com-bustion ramjet from Mach 3 to 5, and as a scram jet fromMach 5 to 27.3. The high terminal speed permitted coastingto a low Earth orbit, with slowdown due to drag through theremaining atmosphere.

Figure 418b is a schematic drawing of a sectional view of thevehicle concept. Compression was provided by a shock fromthe leading edge of the inlet, the convex isentropic turningsurface, the cowl reflected shock, and a flame-induced shock.Properties of the inlet flowfield were obtained assuming su-perposition of the solutions for the inviscid and viscous por-tions of the flowfield. In the concept shown, a portion of theair ingested in the inlet did not participate in the combustionprocess, and thereby provided a film barrier to mitigate theextremely high heat transfer rates in the combustor

Half-viewwith

I augmentation

Fig. 3 Schematic illustration of air-augmented rocket.

===== Shock wave Fuel in— -— Streamline— — — Mach line/2v2viv> Constant pressure

heat addition zone

Fig. 4 Schematic of hypersonic ramjet.

and nozzle. For the cycle calculations made at that time, theflow properties were mass-averaged to obtain conditions en-tering the combustor. Nozzle expansions were computed forthe limiting cases of frozen and equilibrium chemistry, andan assumed value of 0.33 was adopted for the nonequilibriumindex. (See Ref. 52 for a discussion of nozzle loss coefficients.)The author can attest to the tedious nature of the calculationsmade on a mechanical desk calculator and the accompanyingeyestrain from the extensive use of Mollier diagrams. None-theless, these calculations provided the guidance for an in-tensive research program at JHU/APL sponsored by NASAthat began in 1961.

Interestingly, the performance estimates from 1959 are nearlyidentical to those that are being calculated more than 30 yrlater. Moreover, several important design guidelines wererecommended, as shown below.

1) Heat transferred to the vehicle must be conserved toraise the fuel to the highest allowable temperature. Therefore,convectively cooled panels are preferred over radiating sur-faces, as long as the weight penalty is not too great.

2) At high hypersonic Mach numbers, the engine shouldbe operated at a fuel-rich equivalence ratio (ER) that is largerthan that required to provide adequate cooling. Typical valuesare ER = 1.3 at M0 = 12, ER = 4 at M0 = 20, and ER =6 at M0 = 25. High ER limits the maximum temperature inthe combustion zone and thereby reduces the fraction of dis-sociated species in the nozzle expansion, a large-loss mech-anism.

3) The fuel injection angle can be normal to the airstreamat M0 < 10 to enhance penetration, but must approach coaxialat high M0. At M0 > 10, the momentum of the fuel becomesan increasingly important element in the thrust potential ofthe engine.


4) At M0 > 20, stored oxygen can be added to the hydrogenfuel to increase engine thrust and vehicle net force specificimpulse.

5) The flight path should be suppressed to increase enginepressure to the structural limit in order to minimize nozzlenonequilibrium losses. (Dynamic pressure levels of 2000 \bf/ft2 and higher were recommended.)

Initial Conditions for Supersonic CombustionFor practical external burning devices, the conditions en-

tering the combustion zone are approximately the same asthey would be in the absence of combustion. In a free bound-ary flow about a body having a region of expansion, it is verydifficult to produce additional compression by generating aflame. For a wedge airfoil such as that shown in Fig. 1, arealizable objective is to have the heat release delay the turn-ing of the flow at the knee so that the expansion waves donot strike the airfoil. For base flows, such as behind a pro-jectile, effective combustion would displace the expansionwaves emanating from the base corner so that they fall down-stream of the point of wake closure. However, for bodies thatdo not have an expansion region, flame compression can begenerated, but the zone of higher pressure is of limited extent.A wedge-like flame zone can be established where the effec-tive "wedge angle" corresponds to the turning angle that wouldcause separation of the upstream boundary layer. An examplewould be combustion on a flat plate aligned with the flow.53

In summary, the initial conditions for EB lie between thosein the undisturbed freestream and those that correspond to afew degrees of supersonic turning.

Conversely, in the other cited applications, the combustionis confined in a duct and the heat release can produce anupstream shock compression and thereby significantly changethe initial conditions. To explain this effect, consider the flow-field in the ram jet-scram jet engine shown in Fig. 5. In theinlet the pressure increases from the freestream, P0 to Plyacross the forebody shock and to P3 through further turningon the external compression surface. The cowl reflected wavesraise the pressure to P4 at the entrance to the through-duct.The zone from 4 to 5 contains the precombustion shock struc-ture, wherein the pressure rises from P4 to Ps. The section ofduct which confines the shock train to prevent undesirablecombustor-inlet interactions is called an isolator. In the ab-sence of combustion this shock train is not present. The strengthof the shock train is determined by the amount of heat releaseand the effective area ratio A5/A4 of the combustor and, inturn, whether or not the flow at station 5 is thermally "choked."At M0 > 8 the heat release is not sufficient to choke the flow,precombustion shock train disappears, and the pressure dis-tribution in the combustor is similar to a free boundary flow.When initial conditions are considered, both those corre-sponding to 4 and s are pertinent. For ignition or enginerelight, the shock structure would not be established, thus theinitial condition is 4. Once burning is established, 5 is the

initial condition. The pressure decreases from s to 5 in thecombustion chamber and from 5 to 6 in the exhaust nozzle.

To limit the discussion, climb conditions along the flightpath of a transatmospheric accelerator suffice for applications1,3, and 4. For air breathing missiles, the operating envelopeof a Mach 3-8 surface-launched vehicle will be taken. Forthe former, the trajectories from Ref. 54 which temper thebenefits of optimal energy management with the limitationsimposed by active cooling, flutter, and sonic boom are used.For the latter, a lower altitude bound is based on heating andload limits for a passively cooled structure. The upper boundis set by minimum pressure in the combustor.

Figure 6 shows the trajectories for the transatmosphericaccelerator. Takeoff is at 500 ft/s and the first segment to u= 1200 ft/s, Z = 24,218 ft is modeled as Z = 2.035 x 10~2

(u2 - 5002); the next segment to w = 6500 ft/s, Z = 77,938ft is modeled as u = 500 + 7.23 x 10~3 Z + 8.85 x 10~7

Z2. From u = 6500 ft/s to u = 14,000 ft/s, the vehicle fliesat a constant dynamic pressure q = 2000 Ib/ft3. Above u =14,000 ft/s, the constant q trajectory would produce excessiveheat transfer, so in this segment to u = 28,865 ft/s, q is reducedand the trajectory follows u = 39,256 + 0.6647 Z - 1.648x 10~6 Z2. The velocity at the terminal point for poweredflight at Z - 175,000 ft is sufficient to overcome drag andreach a 100-nm circular orbit. A typical shuttle ascent trajec-tory is shown in Fig. 6 for reference.

To compute the mean flow conditions at station 4 in Fig.5, modeling for the inlet compression process is introduced54:Inlet contraction ratio

(A0/A4) = -3.5 + 2.17M0 - 0.017M2

Inlet compression ratio

(P4AP0) - -8.4 + 3.5M0 0.63M2

(1)

(2)

[Those familiar with the literature regarding performanceestimates for scram jet-powered systems are aware that manyauthors have defined the inlet compression process by mod-eling the efficiency and one other parameter, usually M4, h4,or M4. In Ref. 55, arguments are presented that the contractionratio and the compression ratio are the more fundamentalparameters. The experimental data bases and flowfield cal-culations which provided the formulation of Eqs. (1) and (2)have yet to be published in the open literature.]

Radiation to space from the inlet surfaces is assumed toreduce the total enthalpy by 1%. Table 1 lists the mean flowconditions in the freestream and the isolator entrance for theducted supersonic combustion devices. Although the pressurerise in the inlet P4/P0 increases by a factor of 60 over therange of M0, the static pressures at the isolator entrance onlyvary by a factor of 6. The static temperature in the core flowentering the combustor increases by about 200°R for a unit

Inlet Profile Precompresston H ,sh

Length on inner surface

Fig. 5 Flowfield and axial-pressure distributions in a dual-mode ramjet/scramjet engine.

320

280

240

200

160

120

80

40

010 12 14 16 0 200 22 24 26

Velocity x 10-3 ft/s

Fig. 6 Typical trajectories for transatmospheric vehicle.


Table 1 Conditions in the freestream and the combustor inlet in a typical + ram-scramjet engine3

Freestream conditions Isolator entrance conditions

Z0 ,kft r0,°R <?0, Ib/ft2 ht()* Btu/lb,, P4, psia

Reference air breathing trajectory (Fig. 6). bh = 0 (a 536.7°R.

, ft/s3456710152026.8

47,95057,48065,72073,30080,07795,500114,250137,760178,210

1.8681.1837.978 -1

5.569- '4.049- '1.984-1.S.577-23.194-26.759~3

390390390394397.7406.1424.8460.8480.5

2,9043,8724,8405,8396,8449,87915,15521,04028,865

1,6941,9102,0112,0202,0002,0001,9451,287425.8

133.3264.3432.8646.7902.3

1,918.24,561.08,824.616,629.8

1.5291.9452.3632.7673.1434.1435.5026.6508.049

2.864.916.928.9110.8516.4925.2333.1142.45

7.815.724.935.34789.6185.9313.6537.9

14.5118.5719.8619.6519.0317.7815.9410.023.63

744930

1,1021,2791,4511,9582,8804,0745,450

2,0342,8853,7994,7705,7578,74413,90819,64827,070

increase in M0, and reaches a level (T4 > 4000°R) wheredissociation of oxygen becomes significant at about M0 = 19to 20. The values of M4 increase monotonically with M0, andthe ratio of M4/M0 decreases from about 0.5 at M0 = 3 to0.3 at M0 = 26.8. The velocity decrements u0/u4 increase fromabout 870 ft/s at M0 = 3 to about 1785 ft/s at M0< = 26.8. Thevelocities in the combustor remain very high at high M0, whichmeans that the mixing and combustion must be extremelyrapid. For example, at M0 = 20, the residence time in a 2-ft-long duct would be only 100 JLCS.

To obtain the conditions at s, modeling for ER and thecombustor area ratio A5/A4 is needed. Those used herein werebased on design studies of a single-stage-to-orbit vehicle.54

The engine equivalence ratio for hydrogen fuel was 1 for M0< 10, 2 at M0 = 15, 4 at M0 = 20, and 6 at M0 = 26.8.Below M0 = 5, the shock strengths in the isolator were equalto that of normal shocks. The corresponding A5/A4 were 4.18at M0 = 3.0 and 2.60 at M0 = 4.9. For 4.9 < M0 < 8, thecombustor area ratios were varied from A5/A4 = 1.8 to A5IA4 = 1, to limit Ps to 125 Ify/in.2 for structural considerations.Pressure rises across the shock train for this range of M0correspond to those across a single oblique wave. At M0 =8, the combustor area ratio was held at unity, and PJP4 de-creases from about 6 at M0 = 8 to about 3 at M0 > 20.

Figure 7 shows pertinent temperatures of flows in super-sonic combustors over the entire range of flight Mach numbersalong the reference trajectory of Fig. 6. Static temperaturesin the core flow downstream of the precombustion shock trainTs are considerably higher than the corresponding T4 values.Nonetheless, M0 > 5 is required to exceed the autoignitiontemperature of hydrogen. In the absence of heat release, theshock train would not be present and the autoignition con-dition would not be reached at M0 < 9. Consequently, toassure engine relight in the event of flameout or engine un-start, an ancillary ignition source will probably have to beprovided up to M0 = 9. Maximum temperatures in the bound-ary layer can be considerably higher than in the core flowand, therefore, could provide an alternate means for ignition,but the amount of air at these elevated temperatures is smalland the design of a fuel injection system that could exploitthis hotter region of the flow and provide ignition would bearduous.

It should be noted that the existence of shock train struc-tures having pressure rises PJP4 greater than that requiredto separate an incoming boundary layer Psep or, alternatively,the wisdom of designing an engine that would be operated inthis mode, has been the subject of heated arguments at severaltechnical conferences. This has been a continued puzzle tothis author in that the very first tests of supersonic combustorsprovided conclusive evidence of the shock train structure.Moreover, the dual mode scram jet, wherein operation at Ps> Psep is fundamental to the concept, has been successfullydemonstrated by several organizations. Aside from the con-troversy, failure to account for the presence of strong shocktrains in analysis of engine test data leads to spurious results

20000

100008000

: 6000

3 4000\>I" 2000

1000800600

400

Referencetrajectory

ignitionofH2

ionizationregion

\N dissociationregion

O dissociationregion

Vibrationalexcitation

region

Idealgas

region

250 5 10 15 20Flight Mach number

Fig. 7 Temperatures of flows entering supersonic combustors.

and erroneous solutions. In particular, unrealistically fast-mixing and heat release rates are generally deduced.

These comments regarding ignition of hydrogen are per-tinent for hypersonic airplanes at M0 > 8 and transatmo-spheric accelerators. Hydrogen, perhaps augmented by storedoxygen, is the only fuel that can provide the required cooling.Additives to the fuel may prove useful to either enhanceignition, accelerate the recombination reactions, and/or in-crease the density impulse of the fuel. For the thrust aug-mentation application (Fig. 3), the governing conditions re-garding ignition in this "afterburning" mode of operation arethose in the fuel-rich exhaust. Typical "fuel jet" Mach num-bers are 2-3 with static temperatures of 2400-3200°R. Pro-pellants are chosen that have large amounts of excess un-burned hydrogen and ignition is readily obtained. For externalburning systems, temperatures on a 10-deg wedge airfoil typ-ify the initial conditions for supersonic combustion. Theserelatively low temperatures are a fundamental impediment tothe design of a viable EB device. Hydrogen is a candidatefuel, but only with the use of flame holders whose high dragcan easily negate the beneficial effects of EB. The other can-didates are highly reactive liquids, e.g., aluminum, borohy-dride, and the aluminum alkyls or fuel-rich rocket exhausts.

Total temperatures Tt4 are shown in Fig. 7 for referenceand to provide insight to the difficult problem in producingsimulated flows in ground test facilities. An even greater im-pediment is the corresponding total pressures which reach5000 Ib/in.2 at M0 = 10, and rapidly grow to >470,000 \bf/in.2 at M0 = 26.8. Note also, that if the entire inlet compres-sion field needed to be simulated in the ground test, thecorresponding total pressures in the freestream would be>19,000 Ib/in.2 at M0 = 10 and 2.4 x 107 at M0 = 26.8.

Figure 8 shows the operating envelope and initial conditionsfor a supersonic combustion missile, such as suggested in


«JC

1

HN

100

80

60

40

20

o

P4 = 0.33 ATM ^ _ — — — """" *-*— • "*" *"*

x^^"-

—

q = 50,000 Ibs/ft 2

1 1 L — ̂ -^T3 4 5 6 7 8

Flight mach number

4000

3000

2000

1000

2000

4 5 6Flight mach number

3 4 5 6 7Flight mach number

Fig. 8 Operating envelope and initial conditions for a supersoniccombustion mission.

Fig. 2. The lower altitude boundary has a limiting dynamicpressure of 50,000 \bf/in.2 which corresponds to a reasonabledesign limit of known structural materials, passive insulators,radomes, etc. Missiles require storable fuels, which limits theavailable cooling capacity to flight speeds of about M0 = 8.The upper bound is set by a minimum isolator inlet pressureof one-third atmosphere. Both ignition and combustion ef-ficiency become marginal at this pressure level. For mostmissile applications a broad operating envelope is required,and the vehicle must be capable of both acceleration andcruise. Typically, the air breathing engine must be able tooperate at flight speeds as low as M0 = 3 to 4, where thetemperatures T4 are quite low. Moreover, movable parts withinthe supersonic combustor must either be avoided or held toa minimum. This generally prevents the use of retractableflame holders which could provide flame stabilization at lowM0, and could then be removed at high M0, where drag andstructural loads would be excessive. Indeed, these problemshave been so formidable that the only successful missile con-cepts based on the simple scramjet cycle have required theuse of highly reactive fuels.56 Unfortunately, these fuels arelogistically unsuitable for most applications. This realizationprompted the invention of a hybrid scramjet at JHU/APL in1977, known as the dual combustor ramjet, or DCR.57 58

It is appropriate to add a description of the DCR to com-plement the discussion of Figs. 1-4. A schematic illustrationof the DCR concept is shown in Fig. 9. In this sketch anaxisymmetric forebody serves as the initial compression sur-face of the supersonic inlet. In the plane of the cowl lip, the

Supersoniccombustor " 7 V~

air inlet (7YP) Fuel injection \Subsonic

combustor

Fig. 9 Schematic illustration of dual combustor engine.

flow is subdivided into four small sectors and four large sec-tors. The smaller sectors direct flow to a dump-type subsonic -combustor through a duct whose cross-sectional area increaseswith streamwise direction. This provides stable operation overa range of Mach numbers by controlling the position of thenormal shock structure. These inlets are operated super crit-ically, i.e., the normal shock is swallowed because expulsionof the normal shock would produce detrimental interactionswith the flow entering the larger flow passages.

The major portion of the air is turned supersonically towardthe engine axis by the outer cowl compression surface. In thisdesign, the flow in the four ducts is spread circumferentiallyto form an annulus of flow in the dump plane of the subsoniccombustor. The aft portion of these supply ducts is shapedso as to provide a constant or slightly increasing cross-sectionalarea in the streamwise direction, which serves as a combustor-inlet isolator. The shock-train structure in these ducts supportsa pressure rise equivalent to a normal shock when the vehicleis operating at a high ER and low flight Mach number. Forlower ER and/or higher M0, the shock-train pressure risecorresponds to an oblique wave structure. In the "normalshock" operating mode, the mean Mach number at the com-bustor entrance is subsonic, and the mean Mach number inthe combustor exit is either sonic or supersonic. During op-eration in the "oblique shock" mode, the mean flow at allstations throughout the combustor is supersonic. This givesrise to the term "dual-mode" operation, which has been dis-cussed in considerable detail in the literature (e.g., Refs. 59-61).

The dump combustor can act as either a pilot or a gasgenerator to assure that heat can be efficiently released in thesupersonic combustor, even when M0 is low. If it is operatedas a pilot, fuel is added to both streams; if it operated as agas generator, all, or nearly all, of the fuel is added within it,and the main combustor becomes a supersonic "afterburner."

Research ActivitiesThe applied research on supersonic combustion is com-

prised of experimental and analytical studies of "unit pro-cesses," complemented by tests and analysis of the isolator-combustor as a component or as part of an entire engine.Unit processes in this context refer to studies of shock trains,penetration of underexpanded jets, atomization of liquid jets,etc. An entire journal would be required to adequately discussthe body of material on these subjects. Instead, a selectedgroup of the studies at JHU/APL will be described herein,and the remaining, and many of those done elsewhere, willbe cited by reference. The items that are discussed are in-tended to 1) provide the tools to arrive at a conceptual designof a scramjet combustor; 2) present a method for simplifyingand expediting computational fluid dynamics CFD analysesof flows in supersonic combustors; and 3) review experimentalresults to support the veracity of the modeling.

Table 2 lists the subjects that will be discussed and thosethat will be referenced. To keep the total list manageable,only a few representative papers from the large body of lit-erature on computational methods (i.e., CFD) are included.


Table 2 Research items in supersonic combustion

Unit processes ReferencesTests and analysis combustor or

complete engine ReferencesShock trains3

Jet penetrationGaseous injector designShear layer mixingLiquid injection fuels

KineticsWall cooling

62-6869-8324-9019b, 24e, 29, 31, 41, 91-1039, 104-1129, 113-11526, 27, 52, 115-121122-125

Modeling of flow structure3

Coupling of integral and finite-difference techniques3

Instrumentation and data analysisDirect connect isolator-combustor

tests and analysis3

Free-jet engine tests3

17b, 52, 54, 58, 63,130-134

135-13716, 17c, 18b, 35, 51

138-145

58, 60, 61, 146-150

126-129

, 56, 60,

altems discussed in text.

The design of an isolator combustor for a given applicationbegins with a parametric sensitivity study using a cycle analysiscode. The codes usually solve the integral form of the con-servation equations for a range of initial conditions such asthose given in Table 1 and Figs. 7 and 8. Engine ER andcombustor area ratio A5/A4 are the principal variables. Tosolve the integral equations, the forces due to pressure andshear and the heat flux on the lateral surfaces need to bedefined. The key to quantifying the pressure force is the spec-ification of the character of the shock train. Consequently,this subject is chosen for the example of research on a unitprocess.

Modeling of Shock TrainsIn the early days of the supersonic combustion research

program, it was surmised that the shock-train structure, whichhad been observed in tests with fuel injection and combustion,could be duplicated in an underexpanded or throttled non-reacting flow. Moreover, it was felt that Reynolds and Machnumbers would be the fundamental correlating parameters,not pressure and temperature. Consequently, an enormouslysimplified test apparatus could be used. Figures lOb and lOcshow two structures of shock trains in the "cold flow" testapparatus. The duct geometry duplicates the constant cross-sectional area isolator and the upstream portion of the com-bustor. A throttling valve is placed downstream of the step-engine combustor, shown in Fig. lOa, to simulate the blockageeffects of heat release. An alternative method for generatingthe shock train is to lower the air supply pressure to the pointwhere the supersonic flow at 4 is highly over expanded. Asthe throttling valve is closed, a shock wave structure formsin the flow to produce the required pressure rise. At very lowpressure rises, the shock structure is a series of intersectingweak waves (Fig. lOa). As the required pressure rise in-creases, the shocks become stronger and reach the level wherethe initial wave locally separates the boundary layer. If M4 islow and the boundary layer is thick, compatibility in flowdirection can only be attained with the Lambda shock struc-ture shown in Fig. lOb. For practical engine geometries, thepressure rise on the duct walls increases monotonically overthe distance St, but can have quite a different character in theinterior of the flowfield. At high PS/P4 with very thick initialboundary layers or overly long isolators, the effects of vis-cosity change the character of the downstream portion of theinteraction zone. Here, the mean flow conditions can be sub-sonic and the wall pressure reaches a maximum and decreasesdownstream.62

The length scales of the shock-train structure will be thebasis for the engineering design tools. The overall length isdenoted 5r, which is then subdivided into two parts: 1) S0 isthe length between the origin of the shock-train pressure riseand the combustor entrance, and 2) the remaining length Sdis the distance that the shock train extends into the combustor.Figure 10 has intentionally been kept simple. Actual com-bustors generally have far more complex injector configura-tions and wall geometries.

Figure 11 shows wall static pressures for a test in a cylin-drical duct with an entrance Mach number M4 = 2.6. Data

7.Precombustionshock structure injection

b) Overexpanded orthrottled shock structure

Overexpanded orthrottled shock structure

•///////////A

'//S&&&& * '

Y/////////S/,

Fig. 10 Schematics of flow structure: a) isolator combustor with anoblique shock train; b) Overexpanded or throttled nonreacting flowwith an oblique shock train; and c) Overexpanded, throttled nonreact-ing flow with a lambda shock train.

(ps/p4)r=3.05

-15 -10

(Ps/P4)/*= 6.28

(ps/p4)r=9.58

-15 -10 10 15-5 0 I-St//7

Fig. 11 Axial distributions of wall static pressure normalized to thetotal pressure at M4 = 2.6, D4 = 2.75 in., (Pw/Pto)r = 0.387.


from other cases are given in Ref. 63. In all cases, the staticpressure point plotted at the duct exit station PS9 is the max-imum shock-train pressure, whether it be atmospheric or thatgenerated by throttling. The data points for the shock struc-tures originating close to the duct exit represent the lowestvalue of Ps IP4, in which some definition of the shape of thepressure rise could be made (i.e., over a 1-3-in. length). Thehighest data points correspond to the highest value of PS/P4,in which the shock train could be stabilized in the availableduct length. The value of Pw/Pt0 = 0.34 is somewhat lowerthan that corresponding to a simple normal shock, i.e., PJPt0 = 0.42 at M4 = 2.6.

Pitot pressure measurements at radii of 0.125 and 0.375 in.from the duct center line are shown as the solid lines in Figs.12a and 12b. An approximate representation of the flowfieldis shown in 12c, wherein the expansion fans are representedas single waves and curvature of the waves is neglected, exceptat points of intersection. Wave angles are constructed to yieldlocal Mach numbers in general accord with the pitot pressuremeasurements, as shown by the dashed curves in Figs. 12aand 12b. These comparisons indicate that the general char-acter of the oblique shock structure has been depicted, andthat both the compression and expansion processes in realityconsist of a multiplicity of oblique waves which produce con-tinuous, rather than step changes, in P'tIPtQ. Note that theinteraction of the probe shock with the duct wave structurealso leads to a lack of precise definition of wave locations.The calculated value of static pressure across the initialcompression wave is 2.4, and the corresponding local Machnumber is 1.98, which is in very close agreement with theseparation criteria given by the simple modeling151

Ms2ep = 0.58M* (3)

Here, M^p is the Mach number downstream of a single obliqueshock, and Psep/P4 is the corresponding pressure ratio. Theline depicting the boundary of the separated zone was roughlyapproximated from the pitot pressures.

An interesting feature—the similarity in the shape of thepressure traces—was pointed out in Ref. 63 and used as the

• Shock wave• Expansion wave• Streamline

Boundary layer separation 'Measured pitot pressure exitCalculated pitot pressure based on

flow structure depicted above

r= 0.125in.

c)

r= 0.375in.

17 18 19 20 21Distance from isolator entrance

(in.)

22

Fig. 12 Shock train structure: a) simplified representation of shockstructure, b) comparison with pitot measurements at r = 0.125 in.,c) comparison with pitot measurements at r = 0.375 in.

basis for developing the engineering design model. This fea-ture can be appreciated when all of the data for a given M4are superposed by shifting the origin of the pressure rise toa common point. Figure 1363 shows the curves for the datafor four values of M4. For each set of data, the trace for alower overall pressure rise case is simply a portion of thecurve for the highest pressure rise case. Of primary interestin developing an engineering design model is the overall lengthof the shock train, not the details of the shape, so just theend points of the various data sets were used.

It was necessary to introduce some heuristic arguments toexplore the possibility of collapsing all of the data to a singlecurve. The logic that was used included the following consid-erations:

1) For the same overall pressure rise, but a different valueof M4, the length of the structure should vary as the wave-length. Linear aerodynamic theory yields an M\ - 1 wave-angle dependence.

2) If all other influencing factors are held constant, theflowfields should be geometrically similar; thus, SJD wouldbe constant.

3) The length scale should depend on the momentum deficitin the viscous layer, and should be proportional to the totalflow. Moreover, in the limit as the viscous layer approacheszero thickness, Sr should approach zero because a step rise inwall pressure could be accommodated. This situation was ob-served in a combustion experiment described in Ref. 144 inwhich the boundary layer was removed just upstream of theshock. A (0/D)N variation was assumed, where 0 is the mo-mentum thickness of the boundary layer, a measure of themomentum deficit. For the range of experimental data thatwas available, N = 3 gave the best results.

4) The character of the boundary layer should influence thestructure of the shock train, in particular, the initial separa-tion. Thus, a Reynolds number Re dependence would beexpected. As Ref increases, the boundary layer can withstanda larger Psep/P4 before separating, the initial shock angle issteeper, and St is smaller. The form Re® was introduced, andK - \ was obtained from a regression analysis of the data.

The degree of success of this approach is shown in Fig. 14,where all of the data points from Fig. 13 and Ref. 63 have

2 4 6 8 10Axial distance from onset of pressure rise, S (in.)

12

Fig. 13 Shock-train pressure rises for a translated reference point.

Fig. 14(1)].

Shock-train pressure rise vs correlation parameter [see Eq.


been normalized by the defined parameters. Some scatter ispresent, but a simple quadratic relationship

St(M24 -

+ 170(P5/P4 - I)2

= 50(PS/P4 -

(4)

as shown by the correlation curve, is adequate for an engi-neering design model. Note that the empirical correlationdefines not only the endpoint St of the pressure rise curve,but the entire wall pressure distribution, simply by replacingSt with S and PS/P4 with PJP4.

To apply this model to the design of the isolator and com-bustor, it is also necessary to locate the origin of the shocktrain relative to the combustor entrance, i.e., the isolator exit.For a sudden change in surface contour, such as a step be-tween the isolator and the combustor, no ambiguity arises.For other configurations, the location of the farthest upstreamfuel injector is used as the reference point for Sd (Fig. 10).The length Sd is obtained from Eq. (4) by setting PS/P4 =PsepAP4, which is the pressure ratio that will separate theboundary layer at conditions corresponding to M4. The sep-aration pressure ratio can be obtained from experiments orfrom data correlations such as Eq. (3). The minimum requiredisolator length S0 is the difference between the total shock-train length and the distance the shock train extends into thecombustor, St — Sd.

The invention of the dual combustor ramjet introduced anew design requirement for a combustor-inlet isolator in acoannular duct. It was necessary to determine whether theengineering design model for circular configurations wouldalso hold for coannular configurations. A cold flow testapparatus68 with a Mach 1-2.5 central duct flow to simulatethe gas generator (Fig. 9) was used. The supersonic inlet flowwas provided by annular coverging-diverging nozzles with ini-tial Mach numbers M4 of 1.66, 2.35, and 2.89.

The character of the pressure rise curve was similar to thatof the curves for the cylindrical duct, suggesting that the shock-train structure is also similar. By simply changing the "width"scale from D to h, the duct wall separation, and translatingall the shock trains to a common origin, the correlation modelfor the cylindrical duct gives reasonably good results for St(Fig. 15). The cylindrical duct model is also adequate fordetermining Sd when the gas generator flow is overexpandedat the point where the two jets intersect. At very high gasgenerator flows, however, the inner jet will tend to "pump"the annular jet and shift the shock train downstream to thepoint where the static pressures in the two jets match. In thisflow situation, the static pressure at the discharge of the gasgenerator is calculated from the total pressure and the arearatio of the nozzle and is entered into Eq. (4), and the valueof S calculated is set equal to S0 to determine the anchor pointof the shock train.

The extension of the engineering design model to ductswith rectangular cross sections is a topic of current study.Figure 16 shows some preliminary results from "cold flow"

A = Mach 1.66• = Mach 2.35• = Mach 2.89

• EQ. 4• U. Illinois: M= 1.6• JHU/APL engine B1:M= 1.9O JHU/APL cell 1:Ms 3.3

2 3 4 5 6

(St[M?-1lR^/D[q/D]1/2)(1(r3)

Fig. 16 Pressure rise correlation for rectangular ducts.

1 D Approach zone

Flow station 4 f g b

Fuel——I

0 2 4 6 8 1 0

Fig. 15 Shock-train pressure distributions in coannular ducts.

Control Boundary

Fig. 17 Flow processes in combustor with supersonic diffusive flames.

test apparatus at the University of Illinois64 and from direct-connect, isolator-combustor, and semifreejet engines tests atJHU/APL and the NASA Langley Research Center. Wallpressure traces were similar in character to those observed inthe axisymmetric cold flow test apparatus configurations. Forthe data correlation in Fig. 16, the characteristic duct dimen-sion h is taken as the smaller of the width or height of therectangular cross section.

The boundary-layer momentum thickness at the isolatorentrance is calculated or measured, and the largest 0 in thecenterline of a surface is used in the data correlation model.(Momentum thickness is defined as the equivalent height ofinviscid duct flow that would be needed to compensate forthe viscous loss of momentum in the boundary layer.) Thereis somewhat more scatter about the "engineering design" curvethan was experienced in the axisymmetric data sets, perhapsdue to the asymmetry in the viscous-inviscid interaction. Ad-ditional data and analyses are needed to determine whethermodifications to the engineering design model are warranted(e.g.,Ref. 67).

Modeling of the Flow Structure in a Supersonic CombustorThe model used to analyze the flow structure in supersonic

combustors is shown in Fig. 17. The inflow conditions for thefuel and air are either assumed to be uniform, or are pre-scribed by measurements or computations of the incomingstreams on the transverse upstream surfaces of the controlvolume. When conditions at the outflow surface, station 5,can be assumed to be uniform or approximated by definedproperty distributions, integral solutions for specified ER, A5/A47 and combustion efficiency T?C, can be obtained. Modelsare needed for the wall shear, heat transfer, and pressureforce. If the integrated values of each of those terms arespecified, then a unique solution for the downstream prop-erties can be obtained. An infinite number of lateral walldistributions, only some of which could be physically realiz-able, will yield solutions. It is for this reason that entirelydifferent modeling assumptions can yield very similar calcu-lated outflow conditions. The challenge, however, is to devisea model which yields a plausible description of the flow struc-ture which can then be used to define boundary conditions


for finite difference solutions and, in turn, guide combustordesign.

Reference 144 introduced an interesting approach to find-ing unique realistic solutions. The presumption is made thatthe flow approaching the combustor exit plane is one-dimen-sional, and that the wall pressure distribution in this regioncan be represented as the exponential

PA^-i = const (5)

which had first been suggested by Crocco.152 If it is furtherassumed that the flow at station 5 is isentropic, then the math-ematical representation for the derivatives at station 5 is

(6)- MI)]P/A

which can be rearranged as

Mf = /5 + £5(1 - y5)] (7)

In the earlier studies that used this approach, it was assumedthat the shock-train structure was confined to a constant cross-sectional area isolator. With this assumption, the wall pressureat the combustor entrance is Ps. Moreover, if the s = constantpressure-area distribution is applied to the entire combustorwall, then the momentum equation is greatly simplified, i.e.

PA —PA -L (1 — C\IP A — (P IP \P A 1-I 4/*-4 5^^5 V /L 5^^5 \ s 47 4"^^4j

lw = p5u25A5 - p4u2

4A4 - pfu2fAf (8)

where p is the density and pfuJAf is the fuel momentum.

Combustor area ratio, A s/A4

1.0 1.1 1.2 1.3 1.4 1.5 1.6 1.8 2.0 2.2 2.4 2.6 3.0 3.5 4.0

1.0 1.2 1.4 1.6 1.8 2.0 2.2 2.4 2.6 2.8 2.9Total ternperture ratio.Tt5/Tt4

Fig. 18 Shock-train pressure rise as function of heat release (M4 =2.50, y = 1.4, r» = 0).

Solutions to the thereby simplified equations have servedas the basis for the design of several scram jet combustors andprovided considerable insight into their operating character-istics. For example, Fig. 1859 shows the relationship betweenthe maximum pressure rise in the shock train, PS/P4 and A5IA4, for given total temperature increases, Tt5/Tt4. The extremesensitivity of PS/P4 to A5/A4 in these results at M4 = 2.5 aretypical for scram jets operating in the M0 = 4 to 8 regime.The impact of this sensitivity on combustor design is immense.High PS/P4 is detrimental to good combustor design. It greatlyincreases the length requirement of the isolator and simul-taneously increases the design loads. By simply increasing thecombustor area at a given ER (or Tt5/Tt4), the maximumpressure rise is substantially reduced. For example, for Tt5lTt4 = 1.83 increasing A5/A4 from 1.4 to 2.0, decreases PS/P4from 7.125, the value corresponding to a normal shock, to4.6. The corresponding length of required isolator is reducedby a factor of 3.16. There would also be a loss in total pressurePt5/Pt4 of about 8%, but the benefits would far outweigh thisloss.

When the shock train extends into the combustor to pointd in Fig. 17, and the pressure area distribution in the two- orthree-dimensional mixing and combustion zone, d to e, doesnot follow the same exponential relationship e5, then a morecomplex expression for the / P cL4 is needed. One that wassuggested in Ref. 152 is

P dA = KPs(Ad - A4) +

- S5)(PSAS - PeAe)

K,(Ae - Ad)

K = P AAIPs(Ad - A,)

(9)

(10)

obtained from Eq. (4), and the geometry of the isolator andthe portion of the combustor upstream of station d. K-^ ac-counts for a nonlinear pressure-area distribution in the two-to three-dimensional mixing and combustion zone. Some workhas been done on the modeling of K^ and defining the locationof station e, but at present it appears that little is gainedbeyond the simple assumption that stations d and e are coin-cident and that £ = e5 over the entire region d to 5.

Although solutions to the integral equations are readilyobtainable with suitable models, no information regarding theinternal characteristics of the flow are forthcoming. If, how-ever, the integral method is coupled with a finite differencemethod,59'101 very useful additional information can be ob-tained. The procedure is outlined in Fig. 19. The first step isto solve the integral equations for the given initial conditionsand combustor geometry. This yields the axial pressure dis-tribution P(x) for modeled values of wall shear TW and heat

Fig. 19 Method for combining integral and finite difference methods of computation for supersonic combustor analysis.


transfer Qw, and an assumed value for the combustion effi-ciency T]C. With P(x) so-defined, a simplified finite-differencemethod that avoids calculating flow in the separated regionsis made to obtain TW, Qw, and T/C analytically. Inherent to thefinite-difference calculations are models for kinetic rates andturbulent mixing. The integral solution is then repeated usingthese values of TW, Qw, and r/c to obtain a new P(JC), etc. Theiteration ends when the flow area A(x), obtained from thefinite-difference solution, agrees with the geometric areadownstream of station s. In the region of the shock train thecomputed A(x) generally does not match the geometric A(x).

In principle, a finite-difference solution could be obtainedusing the geometric A(x), and thereby avoid depending onP(x) from the integral method. In practice, the presence ofthe separated zone would have to be accurately calculated,and the simplifying assumptions that permit the use of theboundary layer or parabolized forms of the Navier-Stokesequations could not be made. The form of the conservationequations would be elliptic and solutions would have to beobtained using time-dependent techniques. References 154and 155 describe routines for such techniques that requirehours of CPU time on the largest computers. Given the un-certainties that still remain in adequately describing kineticprocesses, their associated rate constants, the transport prop-erties, and the models for turbulence, the simpler, far less-expensive approach has considerable merit.

At present, solutions have been obtained that were basedon the assumption that radial and circumferential pressuregradients are negligible, thereby permitting use of the bound-ary-layer form of the conservation equations. Additionally,it is more expeditious to first solve for the "core" flow in thecombustor, iterate for P(x), and then solve for the boundarylayer using edge conditions from the core flow and a densegrid point spacing in the radial direction. Details of the mod-eling and calculated procedures are given in Refs. 59 and 129.

Whereas, there are numerous contemporary examples thatestablish the veracity of the foregoing modeling of combustionprocesses (e.g., Refs. 63 and 142), results from a few of thepioneering tests of the mid 1960s will be presented herein.Not only is this in keeping with the historical tone of thisPaper, but there were some uniquely fascinating features ofthese tests and test apparatus.

Figure 20144 shows the apparatus and results for a test withan inflow Mach number of 1.95. At this M4, in an effectivearea ratio A5/A4 = 2.27 combustor, the heat release from aneffective equivalence ratio EReff = ERi7c will produce a shocktrain having PS/P4 corresponding to a normal shock. In thisearly test, the need for, and accordingly, the provision of, anisolator was not understood. Indeed, preliminary runs thatwere made without the boundary-layer bleed installed werevitiated, in that the shock train invariably receded into theair supply nozzle, thereby producing poorly defined initialconditions that were atypical of those that would be expectedin an actual engine. To prevent this, a boundary-layer bleedsystem was installed. The flow through the bleed system wasgoverned by the local pressure in the bleed, similar to thesituation that exists in the so-called "educated slot" in theinlet of an air breathing engine. The flow is proportional tothe local pressure in the bleed slot. In the absence of com-bustion, the local pressure is low and the bleed flow is small.In the presence of combustion, the pressure is high, i.e., P5,in this case and the bleed flow into this annular educated slotis high. With the pumping that was provided, the entire up-stream boundary layer was totally removed. Thus, the shocktrain collapsed to a single normal shock with St — 0. Thispoints out the possibility of reducing the length of, or elim-inating an engine isolator if adequate bleed could be providedsince, in accordance with Eq. (4), 0 - 0.

Another interesting feature of the test apparatus was theplatinum gauze strips that were attached at one end to thecombustor wall just downstream of the fuel injector ports.The catalytic effect on the hydrogen fuel produced ignition

Pitot probe rakeTo

Boundary layer bleed exhaust

10 Holes, 0.098-inch-diameter

a) Schematic of hydrogen combustor test apparatus

o.u ,

4.0 ^——V I = 2.004

3.0 — X

2.0 ———^5.£

1.0 — —

0.8

\ M =

•̂\. •

M4 = 1.95T t4=2185'RP4 = 17.4PSIAMs= 0.579Ttf = 1715*Rpp n *r\A

;1the• Wa

oryII stati

\ER Theory=

%N̂

Entropy limitthepry

cpres

0.424

Sy

^S>

sureI

110

aT1.0 1.2 1.4 1.6 1.8 2.0 2.2 2.<

A/A 3 area ratio

b) Combuster pressure distribution

0.5

1.0

1.5

2.0

TheoryI = 2.00

c•

pi°IS-!•

I °SolI • b0|<i

4 ' Op«

Labol°pIII

dpoDW ce

an povece

nts arenterline -

nts arenterline

"TheoryI = 2.0(

°• Wall

Po•

o•

0_ _• _

•0

)4~~

0.2 0.4Exit pitot/inlet total,

0.6 1.4

Pt15

/Pt4

1.6Exit machnumber, M5

1.8

c) Combustor exit profiles

Fig. 20 Comparison of experimental and theoretical results at= 1.95.

at these relatively low temperature conditions (T4 = 1240°R).Reference 156 discusses results of the effectiveness of a widevariety of platinum catalytic ignitors in supersonic combus-tors.

The wall static pressure trace for the burning runs showsthe abrupt pressure rise to P/P4 = 4.27, the value corre-sponding to a normal shock at M4 = 1.95. With the removalof flow in the boundary-layer bleed, the effective area ratioof the combustor was about 10% greater than the geometricarea ratio. Applying the integral method with A5/A4 = 2.27and EReff = 0.424, and assuming the exponential pressure-area distribution holds over the entire combustor, yields s =2.004. Note the precise prediction of the exit pressure P5/P4and the close correspondence between the experimental andtheoretical values at the intersection of the total heat releaserate and entropy limit curves. It is also of interest to note that


if it is assumed that the flow has become nearly unidimen-sional at A/A4 = 1.38, the theory would yield M = 1 in thisplane. The lower solid curve is computed for e = 2.004. Theupper solid curve is the locus of end-point states (P/P4, A/A4)for the measured heat release. The "entropy limit" curve isthe locus of end-point states for various heat release rates.The M = \ curve is the locus of sonic point conditions atdifferent P/P4, A/A4 values.

Figure 20c compares measured pitot pressures and localMach numbers deduced from it with the theoretical values inthe combustor exit plane. The measurements show that thecore flow is near to uniform with an average value slightly inexcess of the theoretical value of M5 = 1.67. In retrospect,this was probably the first demonstration of the dual modeengine concept, wherein the incoming flow passed through anormal shock, became subsonic, and reaccelerated through asonic plane to a supersonic end state.

Figure 21 shows comparison of experimental and theoret-ical wall pressure distributions in a combustor where the heatrelease produces a shock-train pressure rise that is equivalentto that of a single oblique wave. By the time this test wasmade (1967), the need to include an isolator in the test ap-paratus was clearly understood. Bulk combustion efficienciesdetermined from steam calorimeter measurements of thecombustor exhaust gases were 0.81 and 0.92 for ER = 0.78and 0.49, respectively. All direct-connect tests at JHU/APLuse this method to determine the heat released in the com-bustor and, in turn, the required quantity of fuel that wouldhave to burn to local thermodynamic equilibrium at the com-bustor exit. The ratio of required-to-actual fuel flow is thenj]c. Gas sampling measurements from these tests and manyothers142 show that T/C values less than one are generally dueto lack of complete mixing. Indeed, it is difficult to concludefrom measurements of hydrogen-fueled scramjet combustorshaving T}C > 0.5, that losses in rjc are due to slow kinetics.

0Do

T'4CR)413541504150

Pt4

(psia)456456456

M4

3.223.223.22

T4

CR)158115901590

P4(psia)7.527.527.50

\11581160-

ER

0.780.49

0

He

0.810.92-

Model fuelStation Injector 1 42-10.37

Air-flow

• Disk calorimeterA Ring calorimeter

2.67 5.67 23.68 26.67

"^___jj.3.876 ID

f 2.36Isolation cylinder Upstream 1.25

and adapter ring Downstreamcalorimeter ring calorimeter

IT

Note:Dimensions are in inches

a) Schematic illustration of combustor

b) Pressure distributionFig. 21 Comparison of theoretical and experimental pressure distri-butions in short cyl-cone combustor at M4 = 3.22.

The theoretically obtained wall pressure distributions wereobtained using the shock-train modeling and again applyings = c from station 5 to 5. Agreement of theory and experimentis reasonably good and the general features of the flow aresubstantiated. Contrast this to some of the other analyses ofdata that have been repeated in the literature, in particular,the simple, one-dimensional method. In this method the con-servation equations are solved at a finite number of axialstations in the flow, beginning at station 4. At each stationthe flow is assumed to have uniform properties, and any changein pressure above that due to a modeled friction and geo-metrical area change is attributed to heat release. In somecases, notably those with very low EReff, a plausible rate ofheat release is deduced. In others, the deduced results areridiculous. If this method was applied to the pressure distri-bution shown in Fig. 21, the deduced result would be thatwhen the flow reached a station slightly downstream of 5, theamount of heat released would have passed through a max-imum and then would proceed to fall in the major part of thecombustor. However, the final conditions in the flow wouldbe the same as with the shock-train integral method if thevalues of integrated wall shear were the same. Moreover, theone-dimensional method cannot predict the wall pressure dis-tribution for a specified EReff which is the crux of the integralmodel. It can only deduce the integrated value of wall pressureforce.

In these tests, autoignition of the hydrogen was readilyobtained. Although T4 in the undisturbed flow was somewhatbelow 1800°R, the value generally required for autoignition,the strong disturbances caused by the cross-stream injectionproduced local zones of very high temperature. Total tem-peratures in these tests correspond to about M0 = 7.5 flight.

The ultimate use of the modeling and analysis is the guid-ance it provides in engine design. The shock-train and com-bustor modeling that have been described played an importantrole in the design and development of the SCRAM engine.56

Figure 22 is a schematic illustration of a sectional viewthrough one of the modules of a scramjet engine built byJHU/APL. The engine was tested in the Ordnance Aero-physics Laboratory, Daingerfield, Texas, at Mach numbersof 5 and 5.8 in 1968, and at the JHU/APL Avery PropulsionLaboratory at Mach numbers of 7-7.3 in the early 1970s. Thetable lists the important dimensions of the five configurationsthat were tested. In the taper and step configurations, theinlet was directly connected to the injector-combustor. Thelong isolator was designed in accordance with Eq. (4) to ac-commodate the strongest shock train which occurred with highER at M0 = 5.8. The shorter isolator was designed to handleshock trains at M0 = 7-7.3, where the strengths were con-siderably lower. The tests without an isolator were made toconclusively prove the necessity of the added duct length.

Figure 23 shows the thrust coefficients for the first fourconfigurations listed in the table on Fig. 22. The fuel in thesetests was HiCal 3D, a blend of high-density liquid boranes.In the short taper configuration, the maximum thrust coef-

4.5 Capture radiusAll dimensions i

Cowl

V Step/

5 Exit radius

I

Model

TaperStepLong-isolator-taperLong-isolator-stepShort-isolator-step

Ld

0.000.00

13.5013.503.50

LC

22.0022.0022.0022.0018.25

Ln

9.009.009.009.00

15.00

ExitNozzle

15* Conical15* Conical15* Conical15* ConicalContoured

Aex^*

1.00001.00001.00001.00001.2346

* Model exit-to-inlet area ratioModule locationlooking upstream

Dashed lines indicate conical nozzleTaper-tapered combustorStep-step combustor

Fig. 22 Schematic of APL free-jet engines.


Oblique shock .solutions

Symbol ModelTaperStepLong-isolalor-taper or step

0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0Fuel-air equivalence ratio, ER

Fig. 23 Thrust coefficient of APL free-jet engine at Mach 5.0.

ficient CT = Tlq^A^ where AR = 78.54 in.2, occurred at ER= 0.15. Increasing the ER to 0.3 produced a region of reverseflow which was evidenced by small flame zones on the externalsurfaces of the engine downstream of the cowl crotches. Fur-ther increase caused a complete engine unstart and very largenegative values of CT (not shown). Putting in an abrupt stepmitigated the adverse effect, moving the peak CT to ER =0.37. Reference 151 gives compelling arguments to show thata step permits more heat release before reaching the samePS/P4. With the isolator installed in either the taper or stepconfiguration, the engine operated satisfactorily over the en-tire ER range. Similar satisfactory operation was demon-strated at M0 5.8 with the long isolator and at Mach 7-7.3with the shorter isolator. Moreover, the overall operatingcharacteristics of the engine and the performance were inaccord with that predicted by the model presented herein.

Concluding RemarksThe intent of this Paper was to show, by example, that

applied research in supersonic combustion has provided theinformation which enables the design and development ofsome scramjet engines. It seems appropriate to close with afew suggested research activities to address the remainingtechnical issues which must be resolved to realize the fullpotential of the scramjet.

Mixing EnhancementAt high hypersonic speeds, the fuel (hydrogen) must have

an injection angle that is close to coaxial with the air since itprovides a significant portion of the exit momentum. Coaxialmixing is intolerably slow when the relative velocities of thetwo streams approach zero. This occurs at flight speeds of M0== 15, and perhaps could be called the hypersonic "pinchpoint." New concepts are needed to produce effective spread-ing angles which can be less than 1 deg at the hypersonic pinchpoint to required values of 3-4 deg.

Structurally Compatible Injection SystemsMost of the devices and techniques that provide good initial

fuel-air distributions and/or promote mixing at moderate speeds(M0 < 10) are unsuitable at high speeds. They produce localzones having extremely high-heat transfer rates. Typical ex-amples are 1) external corners that produce vortical structuresbut locally thin the protective air boundary layer; 2) discrete-hole cross-stream injectors that produce strong shocks, sep-arated flows, and reattachment zones; and 3) swept instreaminjectors which have excessive leading-edge heat transfer rates.Techniques that exploit the cooling capability of the fuel toprotect the injectors and combustor walls while simulta-neously producing the desired fuel-air distribution are needed.

Fuel DensificationThe energy/unit volume of cryogenic hydrogen is so low

that all-hydrogen fueled systems for many applications maynot be possible (e.g., single-stage transatmospheric acceler-ators). The possibility of a complementary or additive com-

ponent of higher density (e.g., a dense hydrocarbon, boron,boranes, etc.), appears to be attractive. Research on injec-tion, mixing, and combustion of these bipropellant candidatesin the scramjet environment is incumbent to establish feasi-bility.

Turbulence ModelingIt will not be possible to produce conditions in ground tests

for a significant portion of the scramjet flight corridor. Con-sequently, computational techniques will have to be relied onfor many crucial design issues. The greatest deficiency in CFD,at present, is turbulence modeling capable of predicting tran-sition in the inlet, heat transfer, shear and mixing in the com-bustor, and possible relaminization in the nozzle.

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98Gutmark, E., Wilson K. J., Parr, T. P., and Hanson-Parr, D. M.,"Combustion Enhancement in Supersonic Coaxial Flows," AIAA/ASME/SAE/ASEE 25th Joint Propulsion Conf., AIAA Paper 89-2788, Monterey, CA, July 1989.

"Gilreath, H. E., and Schetz, J. A., "Research on Turbulence inHypersonic Engines," APL Technical Review, Vol. 2, No. 1, 1990,pp. 127-139.

1{)0Sullins, G. A., Schetz, J. A., and Schadow, K. C., "ScramjetMixing Studies," APL Technical Review, Vol. 2, No. 1, 1990, pp.140-149.

1()1Schetz, J. A., and Billig, F. S., "Studies of Scramjet Flowfields,"AIAA/SAE/ASME/ASEE 23rd Joint Propulsion Conf., AIAA Paper87-2161, San Diego, CA, June-July, 1987.

102Schetz, J. A., Billig, F. S., and Favin, S., "Analysis of SlotInjection Beneath a Thick Hypersonic Boundary Layer," AIAA/ASME/SAE/ASEE 25th Joint Propulsion Conf., AIAA Paper 89-2457, Monterey, CA, July 1989.

103Schetz, J. A., Thomas, R. H., and Billig, F. S., "Gaseous In-jection in High Speed Flow," Ninth International Symposium on Air-breathing Engines, Athens, Greece, AIAA, New York, Sept. 1989.

104Sherman, A., and Shetz, J., "Breakup of Liquid Sheets and Jetsin a Supersonic Gas Stream," AIAA Journal, Vol. 9, No. 4, 1971,pp. 666-673.

105Schetz, J., Kush, E., and Joshi, P., "Wave Phenomena in LiquidJet Breakup in a Supersonic Crossflow," AIAA Journal, Vol. 18,No. 7, 1980, pp. 774-778.

106Less, D., and Schetz, J., "Penetration and Breakup of SlurryJets in a Supersonic Stream," AIAA Journal, Vol. 21, No. 7, 1983,pp. 1045, 1046.

107Forde, J., Molder, S., and Szpiro, J., "Secondary Liquid Injec-tion into a Supersonic Airstream," Journal of Spacecraft and Rockets,Vol. 3, No. 8, 1966, pp. 1173-1176.

108Reichenbach, R., and Horn, K., "Investigation of InjectantProperties on Jet Penetration in a Supersonic Stream," AIAA Jour-nal, Vol. 9, No. 3, 1971, pp. 469-472.

109Yates, C., and Rice, J., "Liquid Jet Penetration," Research andDevelopment Programs Quarterly Rept., U-RQR/69-2, Applied PhysicsLab., Johns Hopkins Univ., Laurel, MD, 1969.

110Catton, I., Hill, D., and MacRae, R., "Study of Liquid JetPenetration in a Hypersonic Stream," AIAA Journal, Vol. 6, No.11, 1968, pp. 2084-2089.

H1Less, D. M., and Schetz, J. A., "Transient Behavior of Liquid


Jets Injected Normal to a High-Velocity Gas Stream," AlAA Journal,Vol. 24, No. 12, 1986, pp. 1979-1986.

112Hautman, D. J., and Rosfjord, T. J., "Transverse Liquid In-jection Studies," AIAA Paper 90-1965, July 1990.

113Wilson, E., Jr., and Berl, W. G., "Fuels." Applied Physics Lab.,TG 610-11, Johns Hopkins Univ., Laurel, MD, June 1974, Chap. 6.

114Kay, I. W., McVey, J. B., Kepler, C. E., and Chiappetta, L.,"Hydrocarbon-Fueled Scramjet, Piloting and Flame Propagation In-vestigation," Air Force Wright Aeronautical Lab. TR-68-146, Wright-Patterson AFB, OH, Vol. IX, May 1971.

115Kay, I. W., Peschke, W. T., and Guile, R. N., "Hydrocarbon-Fueled Scramjet Combustor Investigation," AIAA Paper 90-2337,July 1990.

116Slack, M. W., "Rate Coefficient for H + O2 + M = HO2 +M Evaluated from Shock Tube Measurements of Induction Times,"Combustion and Flame, Vol. 28, No. 3, 1977, pp. 241-249.

117Bussing, T. R. A., and Eberhardt, S., "Chemistry Associatedwith Hypersonic Vehicles," AIAA Paper 87-1291, June 1987.

118Warnatz, J., "Concentration-, Pressure-, and Temperature-De-pendence of the Flame Velocity in Hydrogen-Oxygen-Nitrogen Mix-tures," Combustion Science and Technology, Vol. 26, Nos. 5 and 6,1981, pp. 203-213.

119Hanson, R. K., and Salimian, S., "Survey of Rate Constants inthe N/H/O System," Combustion Chemistry, edited by W. C. Gar-diner Jr., Springer-Verlag, New York, 1984, Chap. 6, pp. 361-421.

l20Combustion Chemistry, edited by W. C. Gardiner Jr., Springer-Verlag, New York.

121Brabbs, T. A., Lezberg, E. A., Bittker, D. A., and Robertson,T. F., "Hydrogen Oxidation Mechanism with Applications to (1) theChaperon Efficiency of Carbon Dioxide and (2) Vitiated Air Test-ing," NASA TM 100186, Sept. 1987.

122Parthasarathy, K., and Zakkay, V., "An Experimental Inves-tigation of Turbulent Slot Injection at Mach 6," AIAA Journal, Vol.8. No. 7, 1970, pp. 1302-1307.

123Cary, A., and Hefner, J., "Film-Cooling Effectiveness and SkinFriction in Hypersonic Turbulent Flow," AIAA Journal, Vol. 10, No.9. 1972, pp. 1188-1193.

124Hefner, J., and Cary, A., "Swept Slot Film Cooling Effective-ness in Hypersonic Turbulent Flow," Journal of Spacecraft, Vol. 11,No. 3, 1974, pp. 351, 352.

125Billig, F. S., and Grenleski, S. E., "Heat Transfer in SupersonicCombustion Processes," Fourth International Heat Transfer Confer-ence, Heat Transfer 1970, Paris-Versailles, Aug. 1970, Vol. Ill, El-sevier, Amsterdam, 1970, FC 6.1 pp. 1-11.

126Billig, F. S., "Two-Dimensional Model for Thermal Compres-sion," Journal of Spacecraft and Rockets, Vol. 9, No. 9, 1972, pp.702, 703.

127Waltrup, P. J., Billig, F. S., and Stockbridge, E. S., "A Pro-cedure for Optimizing the Design of Scramjet Engines," AIAA/SAE14th Joint Propulsion Conf., AIAA Paper 78-1110, Las Vegas, NV,July 1978; see also Journal of Spacecraft and Rockets, Vol. 16, No.3, 1979, pp. 163-172.

128Waltrup, P. J., Billig, F. S., and Evans, M.C., "Critical Con-siderations in the Design of Supersonic Combustion Ramjet (Scramjet)Engines," AIAA/SAE/ASME 16th Joint Propulsion Conf., AIAAPaper 80-1284, Hartford, CT, June 30-July 2,1980; see also Journalof Spacecraft and Rockets, Vol. 18, No. 4, 1981, pp. 350-357.

129Schetz, J. A., Billig, F. S., and Favin, S., "Analysis of Mixingand Combustion in a Scramjet Combustor with Co-Axial Fuel Jet,"AIAA/SAE/ASME 16th Joint Propulsion Conf., AIAA Paper 80-1256, Hartford, CT, June 1980.

130Schetz, J. A., Billig, F. S., and Favin, S., "Flowfield Analysisof a Scramjet Combustor with a Co-Axial Fuel Jet," AIAA Journal,Vol. 20, No. 9, 1982, pp. 1268-1274.

131Schetz, J. A., Billig, F. S., and Favin, S., "Scramjet CombustorWall Boundary Layer Analysis," AIAA/SAE/ASME Joint Propul-sion Conf., AIAA Paper 81-1434, Colorado Springs, CO, June 1981.

132Billig, F. S., White, M. E., and Van Wie, D. M., "Applicationof CAE and CFD Techniques to a Complete Tactical Missile Design,"AIAA 22nd Aerospace Sciences Meeting, AIAA Paper 84-0387, Reno,NV, Jan. 9-12, 1984; see also "Numerical Methods for Engine-Air-frame Integration," Vol. 102, Progress in Astronautics and Aero-nautics, AIAA, New York, 1986, pp. 192-217.

133Schetz, J. A., Billig, F. S., and Favin, S., "Simplified Analysisof Scramjet Flowfields," Journal of Propulsion and Power, Vol. 4,No. 6, 1988, pp. 172-180.

134Schetz, J. A., Billig, F. S., and Favin, S., "Modular Analysisof Scramjet Flowfields," Journal of Propulsion and Power, Vol. 5,

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136Lee, R. L., Turner, R., and Billig, F. S., "Particulate Mea-surements in the APL Fuel-Rich Ramjet-Combustor Supersonic-Ex-haust Flow," Western States Section/Combustion Inst., Pittsburgh,PA, Oct. 1980.

137Billig, F. S., Lee, R. L., and Waltrup, P. J., "Instrumentationfor Supersonic Combustion Research," 1981 JANNAF PropulsionMeeting, New Orleans, LA; see also CPIA Publication 340, Vol. II,May 1981, pp. 67-100.

138Northam, G. B., Greenberg, L, and Byington, C. S., "Evalu-ation of Parallel Injector Configurations for Supersonic Combus-tion," AIAA Paper 89-2525, July 1989.

139Anderson, G. Y., and Gooderum, P. B., "Exploratory Tests ofTwo Strut Fuel Injectors for Supersonic Combustion," NASA TN-D-7581, 1974.

140Rogers, R. C., and Eggers, J. M., "Supersonic Combustion ofHydrogen Injected Perpendicular to a Ducted Vitiated Airstream,"AIAA Paper 73-1322, Nov. 1973.

141Russin, W. R., "The Effect of Initial Flow Nonuniformity onSecond-Stage Fuel Injection and Combustion in a Supersonic Duct,"NASA TM X-72667, June 1975.

142Billig, F. S., Orth, R. C., and Funk, J. A., "Direct ConnectTests of a Hydrogen Fueled Supersonic Combustor," NASA CR-1904, Aug. 1971.

143Waltrup, P. J., and Billig, F. S., "Prediction of PrecombustionWall Pressure Distributions in Scramjet Engines," Journal of Space-craft and Rockets, Vol. 10, No. 9, 1973, pp. 620-622.

144Billig, F. S., "Design of Supersonic Combustors Based on Pres-sure-Area Fields," Proceedings of the Eleventh Symposium (Inter-national) on Combustion, Combustion Inst., Pittsburgh, PA, 1967,pp. 755-769.

145Avrashkov, V., Baranovsky, S., and Levin, V., "GasdynamicFeatures of Supersonic Kerosene Combustion in a Model CombustionChamber," AIAA Paper 90-5268, Oct. 1990.

146"Hypersonic Research Engine Project Technological Status, 1971,"NASA TM X-2572, Sept. 1972.

147"Hypersonic Research Engine Project—Phase II, Aerother-modynamic Integration Model Development—Final Technical DataReport," AP 75-11133, AiResearch Manufacturing Co. of California;NAS1-6666, NASA CR-132654, May 1975.

148Kepler, C. E., and Karanian, A. J., "Hydrogen-Fueled Varia-ble-Geometry Scramjet," United Aircraft Research Lab., Air ForceAeropropulsion Lab., TR-67-150, East Hartford, CT, Jan. 1968.

149"Investigations of the Low Speed Fixed Geometry SupersonicCombustion Ramjet," General Applied Science Lab., Air ForceAeropropulsion Lab., TR-66-139, Westbury, NY, March 1967.

150Orth, R. C., and Erdos, J. L, "The Use of Pulse Facilities forTesting Supersonic Combustion Ramjet (SCRAMJET) Combustorsin Simulated Hypersonic Flight Conditions," Ninth International Sym-posium on Air Breathing Engines, Vol. 1, Athens, Greece, Sept. 1989,pp. 596-604.

151Magar, A., "On the Model of the Free, Shock Separated Tur-bulent Boundary Taper," Journal of the Aeronautical Sciences, Vol.23, No. 2, 1956, pp. 181-184.

152Crocco, L., "One Dimensional Treatment of Steady Gas Dy-namic," Fundamentals of Gas Dynamics, High Speed Aerodynamicsand Jet Propulsion, Vol. 3, Princeton Univ. Press, Princeton, NJ,1958, pp. 105-130.

153Billig, F. S., "Scramjet Propulsion," Third Joint Europe/U.S.Short Course in Hypersonics, Aachen, Germany, Oct. 1990.

154Kumar, A., "Numerical Simulation of the Flow Through ScramjetInlets Using a Three-Dimensional Navier-Stokes Code," AIAA Pa-per 85-1664, July 1985.

155Kumar, A., "Numerical Analysis of the Scramjet-Inlet by UsingTwo-Dimensional Navier-Stokes Equations," NASA TP-1940, Dec.1981.

156Billig, F. S., and Grenleski, S. E., "Experimental Studies ofHydrogen-Air Ignition in a Supersonic Combustor," Proceedings ofthe AIAA Second Propulsion Joint Specialists Conference," Air ForceSystem Command TR RTD TR-66-1, Vol. 1, Sept. 1966; see alsoApplied Physics Lab./Johns Hopkins Univ. TG 848, Laurel, MD,Aug. 1966.

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