[american institute of aeronautics and astronautics 45th aiaa/asme/sae/asee joint propulsion...

American Institute of Aeronautics and Astronautics

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45th AIAA/ASME/SAE/ASEE Joint Propulsion Conference and Exhibit AIAA 2009-51162-5 August 2009, Denver, CO

Fueling Study on Scramjet Operability Enhancement

Kuo-Cheng Lin1, Chung-Jen Tam2 Taitech, Inc.

Beavercreek, OH 45430

Kevin Jackson3, Paul Kennedy3, Skip Williams4, Dell Olmstead5, MacKenzie Collatz6

Air Force Research Laboratory Wright-Patterson Air Force Base, OH 45433

ABSTRACT

The effects on the operability of a scramjet combustor of diverting fuel to injectors downstream of the cavity flameholder were investigated both experimentally and numerically. The subject features a recessed cavity flameholder, flush-wall low-angle primary injectors upstream of the cavity, and flush-wall normal secondary injectors downstream of the cavity on the body wall. The flight conditions of interest were simulated with Mach 1.8 and 2.2 facility nozzles to cover Mach 3.5 to 5.0 flight conditions. Unheated ethylene was selected as the fuel for both the primary upstream injectors and the secondary downstream injectors. It was found that downstream fuel addition raises the peak pressure, pushes the shock train a short distance toward the isolator entrance, and generates more thrust for a given upstream fuel flow rate. The overall combustion efficiency, however, decreases. For scramjet operation, downstream fuel addition may be able to generate additional thrust without causing isolator unstart. For a given amount of total fuel, diverting a portion of the fuel to downstream injectors pulls the shock train away from the isolator entrance, reduces thrust generation, and decreases both peak pressure and combustion efficiency. For scramjet operation, diverting fuel to downstream injectors may avoid isolator unstart, with a trade –off in combustor performance. For the present flowpath, there is no significant difference in combustor performance or the effectiveness of shock train control among the selected downstream fueling configurations. The further downstream fueling configuration, however, may not couple well with the upstream combustion zone and may also leave less available distance for the injected fuel to burn completely. With an excessively high fuel flow rate, there exists the possibility that the downstream fuel plumes may quench the upstream combustion zone.

NOMENCLATURE Isp = specific impulse ER = fuel equivalence ratio ΔF = net thrust increase Mnozzle = facility nozzle Mach number Mflight = simulated flight Mach number mAIR = air flow rate P = pressure

P0 = total pressure q = dynamic pressure T0 = total temperature x = free stream direction φ = fuel equivalence ratio ηc = combustion efficiency

INTRODUCTION

For a scramjet combustor operating at low flight Mach numbers, an isolator module of an optimal length is required to separate the inlet from combustion-induced pressure increase. The pressure rise generates a shock train system inside the isolator to smoothly elevate the pressure from the isolator entrance to the combustion zone. An excessive pressure increase over a given isolator length can push the shock train out of the isolator, which may then create the undesired unstart condition and limit the combustor operability. Several strategies to expand combustor operability have been proposed and evaluated. Among them, mass removal from wall boundary layer through bleed slots has shown to be effective in preventing inlet and isolator unstart.1,2 Placement, geometry, and control of the bleed slots should be carefully designed so that response time and the amount of bleed mass is optimized to ensure the desired performance.

__________________________________________ 1Senior Research Scientist, 1430 Oak Court, Suite 301, Beavercreek, OH 45430, Associate Fellow AIAA, corresponding author, [email protected] 2Senior Research Scientist, Associate Fellow AIAA 3Aerospace Engineer, Aerospace Propulsion Science Branch, Member AIAA 4Deputy for Science, Aerospace Propulsion Division, Senior Member AIAA 5Deputy Branch Chief, Aerospace Propulsion Technology Branch, Member AIAA 6Deputy Branch Chief, Aerospace Propulsion Science Branch

45th AIAA/ASME/SAE/ASEE Joint Propulsion Conference & Exhibit2 - 5 August 2009, Denver, Colorado

AIAA 2009-5116

Copyright © 2009 by the American Institute of Aeronautics and Astronautics, Inc.The U.S. Government has a royalty-free license to exercise all rights under the copyright claimed herein for Governmental purposes.All other rights are reserved by the copyright owner.


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Following the concept of variable geometry inside a scramjet flowpath, placement of a backward-facing step with a swept ramp in the isolator to enhance isolator performance was studied both numerically and experimentally.3,4 A properly configured ramp could be placed in the isolator flowpath, when needed, to expand combustor operability without isolator unstart. Therefore, the swept ramp design should be retractable, with little travel distance and few seals in practical applications. Inserting parts into the flowpath, however, would require precise movement and robust control, which can be very challenging in high-speed propulsion systems. Heat release distribution within the combustor is another potential strategy to enhance scramjet operability. The idea is to divert a certain amount of fuel toward the downstream section of the combustor, so that the pressure rise due to combustion heat release will not push the shock train too far toward the isolator entrance. With this concept in mind, this study intends to enhance the operability of a cavity-based scramjet flowpath by staging fueling from flush-wall injectors downstream of the cavity flameholder. The objectives of this study are 1) to prove the concept of the proposed operability enhancement strategy and 2) to assess the tradeoff in terms of combustor performance. Both experimental and numerical approaches were utilized in this study.

EXPERIMENTAL METHOD Test Article

The experiment was carried out on the thrust stand inside Research Cell 18 at Wright-Patterson Air Force Base. This facility was designed for fundamental studies of supersonic reacting flows using a continuous-run direct-connect open-loop air flow, supported by the Research Air Facility. The test rig consists of a natural-gas-fueled vitiator, interchangeable facility nozzle (Mach-1.8 and 2.2 currently available), modular isolator, modular combustor, and exhaust pipe, as illustrated in Fig. 1. The rig is mounted to a thrust stand capable of measuring thrust up to 2000 lbf. A series of compressors capable of providing up to 30 lb/s of air, with total pressures and temperatures up to 750 psia and 1600 R, respectively, supply air to the facility. An exhaust system with a pressure as low as 3.5 psia lowers and maintains the backpressure for smooth starting and safe operation. Combined with the currently available Mach-1.8 and 2.2 facility nozzles, the air vitiator was fine-tuned to simulate discrete flight conditions from Mach 3.5 to 5 at flight dynamic pressures up to 2000 psf (Table 1). The relatively low simulated flight Mach numbers represent scramjet takeover conditions, at which dual-mode combustion takes place.

Table 1 Simulated flight conditions of the present study

Mflight q (psf) T0 (R) P0 (psia) Mnozzle mair (lb/s) 5.0 2000 1913 208.3 2.2 6.87 5.0 1000 1950 103.0 2.2 3.36 5.0 500 1992 51.6 2.2 1.67 4.5 2000 1766 211.2 2.2 7.25 4.5 1000 1792 103.5 2.2 3.53 4.5 500 1828 51.4 2.2 1.74 4.0 2000 1389 102.3 1.8 5.75 4.0 1000 1398 53.5 1.8 3.00 4.0 500 1424 26.4 1.8 1.47 3.5 2000 1253 106.9 1.8 6.33 3.5 1000 1253 52.1 1.8 3.09 3.5 500 1278 25.7 1.8 1.50

Figure 1. Schematic of Research Cell 18 combustion facility at WPAFB.


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Figure 2. Schematic of the combustor flowpath and key interior features.

Figure 3. Integrated 3-D schematic of the upstream injection block, cavity flame holder, and downstream injection block on the body wall.

The scramjet flowpath of the present study consists of a heat-sink rectangular isolator, a rectangular combustor featuring a recessed cavity flameholder and flush-wall low-angle injectors, and a truncated nozzle. The isolator has a rectangular cross-section with a height of 1.5 in, a width of 4.0 in, and a length of 25.75 in. The combustor has a total length of 36 in and a constant divergence angle of 2.6 degrees. The truncated nozzle has a length of 8 in and a constant diverging angle of 11 degrees. Figure 2 illustrates the entire flowpath and the arrangement of crucial components. The interior surface of the entire flowpath is covered with thermal barrier coating for additional thermal protection. Two water-cooled combustor side-wall inserts can be replaced with quartz windows for flame visualization and optical measurements. The recessed cavity flame holder is located at the divergent top wall, which is designated as the body side of the scramjet-powered vehicle. A schematic of the cavity flame holder and both upstream and downstream body-side injection sites is shown in Fig. 3. This flame holder spans the entire flowpath width and has a forward-facing ramp to effectively interact with the shear layer originating from the cavity leading edge. A cavity with a length-to-depth ratio (L/D) of 5 was utilized for the present study. The present cavity is similar in general features to those used in Refs. 5-9. Two conventional spark plugs, located at the base of the cavity, are used as the baseline ignition source. There are 8 cavity fuel injectors located at the cavity ramp to provide cavity fuel injection parallel to the cavity base. Four banks of injectors (I-1 – I-4), two banks each on the top (body) and bottom (cowl) walls, were placed upstream of the cavity flameholder to provide various fueling options. The design for the upstream gaseous fuel injectors was adopted from the study of Mathur et al.5 with appropriate scaling of the orifice size. The injection angle is 15 degrees relative to the freestream flow. For the present study, only the I-2 upstream injection site, which has four orifices, was activated to provide the primary fueling. Three banks of injectors (I-5 – I-7) were located downstream


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of the cavity flamefolder to provide secondary fueling options. The injection angle of the secondary fuel injectors is normal to the combustor wall. There are 7 orifices for each downstream injection site. For the I-5 injection site, configurations with three orifices (I-5-3 configuration) or four orifices (I-5-4) were tested. Only four orifices were activated for the I-6 injection site. The I-7 injection site was not utilized for the present study. Combinations of these injector banks provide robust fueling schemes to enhance combustor operability.10,11 Unheated ethylene was used as the fuel for both main injectors and cavity fueling ports.

Basic Instruments Pressure taps and thermocouple ports were strategically positioned throughout the entire rig for instrumentation and health monitoring. The data acquisition system consists of a CAMAC-based crate (128 analog inputs, 16 analog outputs, 48 digital inputs and 32 digital outputs channels), a 256-channel electronic pressure scanning system (Pressure Systems Incorporated) and a 64-channel thermocouple scanning system (Scanivalve Corporation).

NUMERICAL APPROACH All simulations were performed using the CFD++ code, a general-purpose CFD tool developed by Metacomp Technologies.12 CFD++ uses a finite-volume numerical framework, with multi-dimensional Total Variation Diminishing (TVD) schemes and Riemann solvers for accurate representation of supersonic flows. Several types of Riemann solver are available; the Harten-Lax-van Leer-Contact (HLLC) Riemann solver with minmod flux limiting was used in the simulations described here. Multi-grid acceleration is available to provide a fast and accurate solution methodology for both steady and unsteady flows. A variety of one-, two-, and three-equation turbulence models are available for RANS calculations, along with large eddy simulation (LES) and hybrid RANS/LES options. Unless otherwise specified, turbulence was modeled using the two-equation cubic κ-ε model. This model has non-linear terms that account for normal-stress anisotropy, swirl, and streamline-curvature effects. At solid surfaces, an advanced two-layer wall function with equilibrium and nonequilibrium blending was employed to reduce grid requirements. Chemically reacting flows can be modeled with general finite-rate kinetics models. The code supports both structured (quadrilateral and hexahedral) and unstructured (triangle, prism, and tetrahedral) grids. A Message Passing Interface (MPI) is used to take advantage of modern parallel-processing computers. The numerical solutions were considered to be converged based on the residual history and the steadiness of the mass flow rate. The mass flow rate should not change; it should be constant along the whole length of the isolator/combustor. In the case with fuel injection, the mass flow rate behind the injection location should be equal to the sum of the air and fuel mass flow rates Building on the semi-global model developed by Singh and Jachimowski,13 a recently developed model, using quasi-steady state (QSS) approximations14 was employed for the simulations of ethylene-fueled supersonic combustion in the present study. This new model is referred to subsequently as the Princeton v2.0 model. This reduced model was based on a detailed ethylene oxidation mechanism from Qin et al.15

that consists of 70 species and 463 elementary reactions. First, a skeletal reduction was applied to identify and eliminate unimportant species and reactions. For the skeletal reduction, perfectly stirred reactor (PSR) and auto-ignition calculations were used over a range of conditions yielding a skeletal mechanism consisting of 33 species and 205 elementary reactions. Then a time scale reduction technique (QSS approximation) was applied. The computational singular perturbation (CSP) method was used to identify the QSS species again using PSR and auto-ignition as the data bases. A species was assumed to be a QSS species if its worst case normalized time scale was shorter than a specified threshold value. The QSS species were then removed from the skeletal mechanism and an internal algebraic loop was used to solve the concentrations of the QSS species. The final reduced mechanism consisted of 22 species. The removal of the short time scales effectively reduces the stiffness of the system as well as the number of differential equations.

RESULTS AND DISCUSSION Flame Visualization

Figure 4 shows the appearance of the flame inside the present scramjet flowpath at various fueling schemes. The flight condition of Mflight=3.5 and q=1000 psf was simulated, in order to avoid excessive thermal shock, which might damage the quartz window. The flow is from right to left. For the I-2-only fueling scheme in Fig 4(b), the flame originates from the cavity shear layer and spreads gradually into the freestream core flow. There is no combustion within the region immediately behind the cavity rear-facing step. As 50% of the fuel is directed from the I-1 injectors to the I-5 downstream injectors in Fig. 4(c), initial flame appearance is similar, except for the region of orange luminosity behind the cavity step. The soot coating on the window is the result of previous tests and was


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partially burned out after this test. As the flame propagates downstream, a darker region, as highlighted in Fig. 4(c), represents the fuel plumes injected from the I-5 downstream injectors. The fuel plumes, which are too rich to burn immediately after injection and, therefore, appear as the darker region, stay within the combustion zone from the upstream fuel. The downstream fuel plumes actually serve as the blockage to lift the upstream combustion zone. With no freshly entrained freestream air to interact with the downstream fuel plumes, the resulting combustion is expected to be less intensive. Please note that the flame appearance in Fig. 4 is line-of-sight and, therefore, can be dominated by flame distributed along the sidewall window.

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Figure 4. Photographs of the flame inside the scramjet combustor operated at various fueling conditions. (a) Flowpath without freestream air to illustrate the fueling locations, (b) ERI-2=0.60, (c) ERI-2=0.30, ERI-5=0.30, (d) ERI-2=0.10, ERI-5=0.41. Mflight=3.5, q=1000 psf. Flow is from right to left. Figure 4(d) shows the flame appearance with ERI-2=0.10 and ERI-5=0.41. With less fuel injected from the I-2 upstream injectors for this condition, flame spreading in the transverse direction is limited. Also, combustion appears to be complete before the I-5 downstream fuel plumes, as evidenced by the faded blue luminosity. With more freestream air not consumed by the I-2 fuel plumes, the downstream fuel plume can still be ignited by the upstream residue flame or by hot combustion products. Flame spreading in the transverse direction near the combustor exit appears limited in Fig. 4(d), since there is less distance in the freestream direction for the ignited downstream fuel flumes to propagate transversely. In addition, it is highly likely that the downstream fuel plumes may not burn completely within the remaining distance inside the combustor.

Effects of Downstream Fuel Addition In order to explore the coupling between upstream flame and downstream fueling, Fig. 5 shows the distribution profiles of wall static pressure within the isolator and the combustor, with a gradual increase in downstream fueling. The simulated flight condition has a flight Mach number of 4.5 and a dynamic pressure of 1000 psf. The fuel flow rate for the upstream I-2 injectors was kept at 0.20. The I-5 downstream injection site with three injectors equally spaced in the spanwise direction was activated for downstream fuel injection. With the I-2-only fueling at such a low fuel flow rate (F08295AV), the pressure rise due to combustion is mainly confined within the combustor. With ERI-

5=0.05 (F08295AW), there exists a very small pressure increase from the I-2-only fueling downstream of the I-5 injection site. The peak pressure remains the same.


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Figure 5. Wall static pressure distribution profiles within the isolator and combustor with various fuel splits between I-2 and I-5 injectors. Mflight=4.5, q=1000 psf. As ERI-5 is increased to 0.14 (F08295AX), a secondary peak in pressure distribution downstream of the I-5 injection site can be observed. The fact that the secondary peak is not located at the I-5 injection site indicates that the injected fuel is not burned immediately. The downstream fuel plumes probably do not have high enough momentum to penetrate through the upstream combustion zone initially. After traveling about 5-in in the freestream direction, however, sufficient fuel plume penetration and fuel/air mixing are achieved for combustion. For the condition with ERI-5=0.24 (F08295AY), the secondary pressure rise occurs at the I-5 injection site, probably due to the immediate combustion of the I-5 fuel plumes. With the fuel plumes penetrating better and interacting with the fresh freestream air, combustion is expected to take place sooner. In addition, the deeply penetrated fuel plumes provide the aerodynamic blockage to the freestream air with an effect similar to air throttling, resulting in an increase to the primary peak pressure around the cavity. Finally, a further increase in ERI-5 to 0.34 (F08295AZ) significantly raises the primary peak pressure, with the secondary pressure rise still taking place at the I-5 injection site.

Figure 6. Net thrust increase due to combustion for the test conditions in Fig. 5. Mflight=4.5, q=1000 psf.

Based on the observations from Fig. 5, the degree of overall combustion within the present flowpath increases with I-5 downstream fueling. Figure 6 demonstrates the measured net thrust increase due to combustion from the load cell readings. The net thrust increase is the difference between load cell readings before flame ignition with the fuel


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already injected at the desired flow rate and after steady sustained combustion with the air throttle completely removed. For the present study, the measured net thrust increase is used as an indicator of relative combustor performance. As can be seen in Fig. 6, fuel addition from the I-5 downstream injectors increases the net thrust. The specific impulse, defined as the net thrust increase divided by the total fuel flow rate, actually decreases with downstream fueling, as illustrated in Fig. 7. It appears that the fuel injected from the downstream injector may not react efficiently, since a portion of the injected downstream fuel remains unburned and there is a limited distance within the combustor for complete combustion.

Figure 7. Specific impulse for the test conditions in Fig. 5. Mflight=4.5, q=1000 psf.

For the test conditions in Fig. 5, the extent of the primary combustion zone is limited by a relatively small fuel flow rate from the I-2 injection side. The downstream fuel plumes can penetrate through the primary combustion zone with ease. Figure 8 shows two test series with a significant amount of fuel injected from the upstream I-2 injectors, while more fuel is gradually added to the I-5 downstream injectors. For the test series with ERI-2 around 0.60 (0.58-0.64) in Fig. 8(a), the primary peak pressure around the cavity region increases and the pre-combustion shock train moves toward the isolator entrance as more fuel is injected from the I-5 injection site. The secondary pressure rise, however, does not occur at the I-5 injection site. Instead, the secondary pressure rise due to the downstream injector takes place further downstream, indicating that the I-5 fuel plumes may not penetrate quickly through the significantly extended primary combustion zone to react with fresh freestream air. Consequently, the fuel injected from the downstream injectors may not burn within the limited combustor length. The problem of incomplete combustion of the downstream fuel can also be seen for the test series with ERI-2 around 0.80 (0.79-0.80) in Fig. 8(b). The downstream fueling contributes an insignificant pressure rise. The shock train moves less than 5” toward the isolator entrance as downstream fuel at up to as much as ERI-5=0.47 is added to the flowpath. The combustor operability in terms of the unstart limit shows a significant improvement with downstream fueling, since the overall ER increases from 0.80 to 1.27 in Fig. 8(b). The combustor performance, however, may not improve proportionally. Based on the observations so far, fuel addition at the downstream location should be applied with the extent of the primary combustion zone in mind. Figure 9 shows the measured net thrust increase and the specific impulse for the test conditions in Fig. 8. Also included in Fig. 9 are two test conditions with I-2-only injection at ERI-2=0.90 and 1.10, to illustrate the effect of downstream fuel injection. Those test conditions with similar total ER are identified for easy comparison in Fig. 9. As observed from Fig. 6, the net thrust increases as the amount of downstream fuel increases for both test series in Fig. 9(a). At the same overall ER, the injection condition with a larger portion of fuel delivered through the I-2 upstream injectors generates a higher thrust increase. Both conditions with I-2-only injection at ERI-2=0.90 and 1.10 exhibit the highest net thrust increase for each overall ER group. In terms of the measured specific impulse, the test condition with a large portion of fuel injected through the I-5 downstream injectors exhibits relatively poor performance, as shown in Fig. 9(b). For the ERI-2=0.60 test series, the


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specific impulse shows a reverse trend versus the overall ER, once ERI-5 is greater than 0.29. For the ERI-2=0.80 test series, the specific impulse decreases as more fuel is injected through the I-5 downstream injectors, as onserved in Fig. 7.

(a) (b) Figure 8. Wall static pressure distribution profiles within the isolator and combustor with various fuel splits between I-2 and I-5 injectors. (a) ERI-2 maintained around 0.60 with ERI-5 gradually increased, (b) ERI-2 maintained around 0.80 with ERI-5 gradually increased. Mflight=4.5, q=500 psf.

(a) (b) Figure 9. Net thrust increase due to combustion and specific impulse for the test conditions in Fig. 8. Mflight=4.5, q=500 psf. Figure 10 demonstrates another negative impact of excessive downstream fueling on the flame structure inside the scramjet flowpath. For the F08301AS test condition (ERI-2=0.15, ERI-5=0.15), the secondary pressure rise starting from the I-5 injection site shows that the injected downstream fuel burns properly, as observed in Fig. 5. For the F08301AU test condition (ERI-2=0.15, ERI-5=0.49), however, instead of an increased secondary pressure size, a steep drop in wall static pressure after the I-5 injection site indicates that the fuel injected downstream does not burn as expected. At ERI-5=0.49, the local fuel/air mixture around the I-5 injection site may be too rich, or the fuel plumes


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may penetrate too deeply to couple with the limited combustion propagating from upstream. It is also possible that the downstream fuel may experience the condensation phenomenon, due to the large drop in ethylene injection pressure. With a small amount of fuel injected from the I-2 upstream injectors for the condition in Fig. 10, an excessive amount of fuel injected through the I-5 downstream injectors can actually quench the flame propagating from the cavity flameholder, even though a substantial amount of air is still available to react with the downstream fuel plumes.

Figure 10. Wall static pressure distribution profiles within the isolator and combustor with two fuel splits between I-2 and I-5 injectors to illustrate the effect of excessive downstream fueling. ERI-2 maintained around 0.15. Mflight=3.5, q=1000 psf.

Effects of Fuel Diversion on Combustor Performance The effects on pressure rise and shock train location of diverting a portion of a given amount of fuel to the downstream injectors are shown in Fig. 11, with four different testing series using the I-5 or I-6 downstream injectors. The general trends observed from these test series include the reduction in peak pressure rise and the pullback of the shock train to a downstream location inside the isolator, as a portion of the fuel is diverted to the downstream injectors. From the operability point of view, the strategy of injecting more fuel from the downstream injectors leaves a sufficient margin within the isolator to contain the shock train and, therefore, helps to prevent isolator unstart; the combustor can actually operate at a higher total ER without risking unstart. For example, the shock train can be pulled back from close to the isolator entrance with the F08301AK fueling condition (ERI-2=0.61, ERI-6=0) to the rear section of the isolator with the F08301AM fueling condition (ERI-2=0.20, ERI-6=0.36) in Fig. 11(b). Combustion quenching due to excessive downstream fueling, as illustrated in Fig. 10, is observed for the F08301AE test condition (ERI-2=0.11, ERI-6=0.51) in Fig. 11(a). Excessive fuel diverted to the downstream injectors should be avoided, to mitigate the negative effect on overall combustor performance. It can also be observed that the shock train location inside the isolator is mainly affected by the I-2 upstream fueling, with more fuel from the I-2 injectors creating a shock train location close to the isolator entrance. The measured net thrust increase due to combustion from the load cell readings and the specific impulse for the test conditions in Fig. 11 are tabulated in Table 2. With more fuel diverted to the downstream injectors, both the measured thrust increase and the specific impulse decrease. Therefore, the increase in combustor operability comes with the price of reduced combustor performance.

Effects of Downstream Injection Placement Figure 12 shows the effect of spanwise injector placement at the I-5 downstream injection site on combustor performance. The I-5-4 and I-5-3 configurations engage 4 and 3 orifices at the I-5 injection site, respectively. Injectors at the I-5-4 configuration are aligned with the 4 orifices at the I-2 injection site, while the injectors at the I-5-3 configuration are positioned between I-2 injectors in the spanwise direction. The I-5-3 configuration was originally designed to spread the flame between the I-2 fuel plumes at the downstream location. Also, the I-5-3 configuration provides downstream fueling slightly away from the combustor sidewalls, where extensive


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combustion takes place due to low momentum corner and boundary layer flows.10 As can be seen in Fig. 12, there exists no significant difference in pressure rise and shock train location between the I-5-3 and I-5-4 configurations for similar fuel splits between the I-2 and I-5 injection sites. The measured thrust increase for those test conditions in Fig. 12 also shows similar values, as tabulated in Table 3. With the cavity flameholder to enhance mixing and combustion, flame distributions downstream of the cavity exhibit no distinct separation between I-2 fuel plumes.

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SYM. RUN CONFIG. ER F08301AA I-2-4/I-6-4 0.59/0.00 F08301AC I-2-4/I-6-4 0.31/0.31 F08301AE I-2-4/I-6-4 0.11/0.51

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SYM. RUN CONFIG. ER F08301AK I-2-4/I-6-4 0.61/0.00 F08301AL I-2-4/I-6-4 0.30/0.29 F08301AM I-2-4/I-6-4 0.20/0.36

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SYM. RUN CONFIG. ER F08295AL I-2-4/I-5-4 1.10/0.00 F08295AI I-2-4/I-5-4 0.80/0.29 F08295AE I-2-4/I-5-4 0.59/0.49

(c) (d) Figure 11. Wall static pressure distribution profiles within the isolator and combustor with various fueling and simulated flight conditions. (a) Mflight=3.5, q=1000 psf, ERtotal=0.60, I-2 and I-6 injectors, (b) Mflight=4.0, q=2000 psf, ERtotal=0.60, I-2 and I-6 injectors, (c) Mflight=4.5, q=500 psf, ERtotal=0.90, I-2 and I-5 injectors, (a) Mflight=4.5, q=500 psf, ERtotal=1.10, I-2 and I-5 injectors.

The effects of downstream injector placement in the freestream direction are demonstrated in Fig. 13. At similar fuel splits between the I-2 and downstream injectors, the I-6 downstream fueling scheme exhibits a slightly smaller pressure rise and a shock train location slightly downstream. The measured net thrust increase in Table 4, however, shows no significant difference between the two downstream injection schemes at the simulated flight condition of Mflight=5.0 and q=500 psf. For the two test conditions with I-6 downstream fueling in Fig. 13(a), the combustion of the I-6 fuel plumes can be clearly seen from the pressure increase immediately downstream of the I-6 injection site. Neither of the test conditions with I-5 downstream fueling in Fig. 13(a), however, exhibits a pressure rise due to


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combustion of the I-5 fuel plumes immediately downstream the I-5 injection site. The selection of the downstream injection location clearly affects the flame structures in the downstream section of the combustor.

Table 2 Net thrust increase and the specific impulse for test conditions in Fig. 11. Figure Run ERI-2 ERI-5 ERI-6 ERtotal ΔF (lbf) Isp (lbf/lbm)

Fig. 11(a) F08301AA 0.59 0 0 0.59 122 982 F08301AC 0.31 0 0.31 0.62 94 762 F08301AE 0.11 0 0.51 0.62 30 266

Fig. 11(b) F08301AK 0.61 0 0 0.61 230 966 F08301AL 0.30 0 0.29 0.59 211 861 F08301AM 0.20 0 0.36 0.56 180 872

Fig. 11(c) F08295AK 0.90 0 0 0.90 88 827 F08295AH 0.80 0.10 0 0.90 79 749 F08295AD 0.59 0.29 0 0.88 72 695

Fig. 11(d) F08295AL 1.10 0 0 1.10 102 796 F08295AI 0.80 0.29 0 1.09 80 626 F08295AE 0.59 0.49 0 1.08 75 590

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SYM. RUN CONFIG. ER F08295AD I-2-4/I-5-4 0.59/0.29 F08295AO I-2-4/I-5-3 0.59/0.27 F08295AF I-2-4/I-5-4 0.58/0.69 F08295AR I-2-4/I-5-3 0.56/0.72

Figure 12. Wall static pressure distribution profiles within the isolator and combustor with I-5-3 and I-5-4 downstream injectors to illustrate the effect of spanwise injection placement. Mflight=4.5, q=500 psf. Figure 13(b) shows another testing series to demonstrate the effect of downstream injector placement on combustor performance at the simulated flight condition of Mflight=3.5 and q=1000 psf. Unlike the test conditions for Fig. 13(a), the I-6 downstream injection scheme in Fig. 13(b) exhibits features of poor combustor performance, such as lower peak pressure rise, downstream shock train location, quenched combustion, and smaller net thrust increase (Table 4). It is interesting to observe the significant difference between F08301AQ (I-5 fueling) and F08301AE (I-6 fueling) in Fig. 13(b). While the pressure profile for the F08301AE test condition shows signs of quenched combustion (or unsuccessful ignition of downstream fuel plumes), the pressure profile for the F08301AQ test condition shows robust combustion for the downstream fuel plumes to significantly raise the peak pressure and also push the shock train forward.


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Based on the observations so far, it is recommended to place the downstream injectors close to the cavity flameholder, in order to provide better coupling between the downstream fuel plumes and the upstream combustion zone and to allow a longer distance and time for complete combustion of the downstream fuel. Excessive downstream fueling should be avoided, as it may quench the flame.

Table 3 Net thrust increase and the specific impulse for test conditions in Fig. 12. Run ERI-2 ERI-5-4 ERI-5-3 ERTotal ΔF (lbf) Isp (lbf/lbm)

F08295AD 0.59 0.29 0 0.88 72 737 F08295AO 0.59 0 0.27 0.86 70 605 F08295AF 0.58 0.69 0 1.27 89 576 F08295AR 0.56 0 0.72 1.28 88 617

0 10 20 30 40 50 60 70x (in)

0

5

10

15

20

25

30

35

p (

psi

a)


I-5 I-6

SYM. RUN CONFIG. ER F08297AS I-2-4/I-5-4 0.60/0.29 F08297BC I-2-4/I-6-4 0.59/0.29 F08297AT I-2-4/I-5-4 0.60/0.51 F08297BB I-2-4/I-6-4 0.59/0.52

0 10 20 30 40 50 60 70

x (in)

0

10

20

30

40

p (

psi

a)


I-6I-5

SYM. RUN CONFIG. ER F08301AO I-2-4/I-5-4 0.30/0.30 F08301AC I-2-4/I-6-4 0.31/0.31 F08301AQ I-2-4/I-5-4 0.11/0.50 F08301AE I-2-4/I-6-4 0.11/0.51

(a) (b) Figure 13. Wall static pressure distribution profiles within the isolator and combustor with I-5 and I-6 downstream injectors to illustrate the effect of injector placement in freestream direction. (a) Mflight=5.0, q=500 psf, (b) Mflight=3.5, q=1000 psf.

Table 4 Net thrust increase and the specific impulse for test conditions in Fig. 13. Figure Run ERI-2 ERI-5 ERI-6 ERtotal ΔF (lbf) Isp (lbf/lbm)

Fig. 13(a)

F08297AS 0.60 0.29 0 0.89 66 677 F08297BC 0.59 0 0.29 0.88 67 712 F08297AT 0.60 0.51 0 1.11 73 491 F08297BB 0.59 0 0.52 1.11 76 622

Fig. 13(b)

F08301AO 0.30 0.30 0 0.60 125 994 F08301AC 0.31 0 0.31 0.62 94 762 F08301AQ 0.11 0.50 0 0.61 141 1037 F08301AE 0.11 0 0.51 0.62 30 266


13

(a) (b)

(c) (d) Figure 14. Numerical mass-averaged 1-D distribution profiles of (a) static temperature, (b) Mach number, (c) mixing efficiency, and (d) combustion efficiency for selected fueling conditions at Mflight=4.5 and q=1000 psf.

Table 5 Fuel flow rates and combustion efficiency at the exit plane of the truncated nozzle for the fueling conditions in Fig. 14.

Run ERI-2 ERI-5-4 ERI-5-3 ERI-6-4 ERTotal CFD

Combustion Efficiency F08295AB 0.61 0 0 0 0.61 0.72 F08295AD 0.59 0.29 0 0 0.88 0.52 F08295AO 0.59 0 0.27 0 0.86 0.53 F08297AE 0.60 0 0 0.31 0.91 0.52 F08295AK 0.90 0 0 0 0.90 0.66 F08295AH 0.80 0.10 0 0 0.90 0.60

Numerical Simulations for Selected Fueling Conditions

Figure 14 shows the numerical mass-averaged 1-D distribution profiles of wall static pressure, Mach number, mixing efficiency, and combustion efficiency for representative fueling conditions selected from the test matrix. The corresponding fuel flow rates and the combustion efficiency at the exit of the truncated nozzle for each fueling condition are tabulated in Table 5. The effects of downstream fuel addition, fueling diversion to downstream injectors at a given total fuel flow rate, and placement of downstream injectors at I-5-3, I-5-4, and I-6-4 injection sites on combustor performance are discussed.


14

The major findings from the numerical simulations shown in Fig. 14 agree well with those obtained previously from the experimental testing. First of all, downstream fuel addition increases the peak pressure and moves the shock train toward the isolator entrance with a reduction in combustion efficiency. (See F08295AB, F08295AD, F08295AO, and F08297AE.) At a given total fuel flow rate, the strategy of diverting a portion of the fuel to the downstream injection site decreases the peak pressure, pulls the shock train toward the combustor, and increases the combustor operability with a reduction in combustion efficiency. (See all test conditions except for F08295AB.) The more fuel diverted to the downstream injection site, the lower the combustion efficiency. Also, there appears to be no significant difference in flame structure and combustion efficiency among the I-5-3, I-5-4, and I-6-4 downstream fueling configurations for the same fueling split between up and downstream injectors. (See F08295AD, F08295AO, and F08297AE.) The F08295AO case with the I-5-3 configuration shows slightly better performance, probably due to the slightly lower fuel flow rate for the I-5-3 downstream injectors from the actual test condition. Figure 15 shows the numerical fuel equivalence ratio contours at several planes for the fueling conditions in Fig. 14. For those conditions with downstream fueling, distinct downstream fuel plumes with high fuel equivalence ratios exist within the plotting range in Fig. 15, indicating inefficient combustion for the downstream fuel.

(a) (b) (c)

(d) (e) (f) Figure 15. Numerical fuel equivalence ratio contours on symmetry plane and three cross-sectional planes at different freestream locations for the fueling configurations in Figure 14. The white line indicates the stoichiometry. Flow is from left to right.

CONCLUSIONS The effects on the operability enhancement of a scramjet combustor of diverting fuel to injectors downstream of the cavity flameholder were investigated both experimentally and numerically, using an AFRL research scramjet flowpath at Wright-Patterson Air Force Base. This flowpath features a recessed cavity flameholder, flush-wall low-angle primary injectors upstream of the cavity, and flush-wall normal secondary injectors downstream of the cavity on the body wall. The flight conditions of interest were simulated with Mach 1.8 and 2.2 facility nozzles to cover Mach 3.5 to 5.0 flight conditions. Unheated ethylene was selected as the fuel for both the primary upstream injectors and the secondary downstream injectors. The major conclusions of the present study are as follows:

1. Since both upstream and downstream injectors are located on the body wall in the present flowpath, downstream fuel plumes must interact and couple with the upstream combustion zone to establish stable and efficient combustion. Downstream fuel plumes with a low fuel flow rate stay within the upstream combustion zone without encountering freshly entrained freestream air for immediate combustion and, consequently, may not burn completely within the limited available distance inside the combustor. With an excessively high fuel flow rate, there exists the possibility for the downstream fuel plumes to quench the upstream combustion zone.

2. For a given upstream fuel flow rate, downstream fuel addition raises the peak pressure, pushes the shock

train a short distance toward the isolator entrance, and generates more thrust. The overall combustion


15

efficiency, however, decreases. For scramjet operation, downstream fuel addition may be a reasonable strategy to generate additional thrust without causing isolator unstart, when there is no significant margin for shock train movement and additional thrust is needed for critical scramjet maneuvers.

3. For a given amount of total fuel, diverting a portion of the fuel to downstream injectors pulls the shock

train away from the isolator entrance, reduces the thrust generation, and decreases both peak pressure and combustion efficiency. For scramjet operation, diverting fuel to downstream injectors may be an effective strategy to avoid isolator unstart with a trade-off of combustor performance.

4. For the present flowpath, there is no significant difference in combustor performance or the effectiveness

of shock train control among the selected downstream fueling configurations, which have variations in injector placement in spanwise and freestream directions. The further downstream I-6 fueling configuration, however, may not couple well with the upstream combustion zone and may also leave less available distance for the injected downstream fuel to burn completely.

ACKNOWLEDGEMENTS

This work was sponsored by the AFRL/Propulsion Directorate under contract number FA856008D2844 (Contract monitor: Robert Behdadnia). Assistance from the air facility group of the Air Force Research Laboratory is acknowledged. The authors would like thank Matt Streby and Steve Enneking (Taitech, Inc.) for their great assistance in rig operation, data acquisition, and hardware design and setup.

REFERENCES 1. Tam, C-J., Eklund, D., Behdadnia, R., and Jackson, T.A., “Investigation of Boundary Layer Bleed for

Improving Scramjet Isolator Performance,” AIAA Paper 2005-3286, 2005. 2. Iannelli, J., “A Distributed Wall Mass Removal System for Improving Scramjet Isolator Performance,” AIAA

Paper 2008-0065, January, 2008. 3. Lin, K.-C., Tam, C.-J., Jackson, K., Kennedy, P., and Behdadnia, R., “Experimental Investigations on Simple

Variable Geometry for Improving Scramjet Isolator Performance,” AIAA Paper 2007-5378, July, 2007. 4. Tam, C.-J., Lin, K.-C., Davis, D., and Behdadnia, B., “Numerical Investigations on Simple Variable Geometry

for Improving Scramjet Isolator Performance,” AIAA Paper 2006-4509, 2006. 5. Mathur, T., Gruber, M., Jackson, K., Donbar, J., Donaldson, W., Jackson, T., Billig, F., “Supersonic

Combustion Experiments with a Cavity-Based Fuel Injector,” Journal of Propulsion and Power, Vol. 17, No. 5, 2001, pp. 1305-1312.

6. Ben-Yakar, A. and Hanson, R., “Cavity Flame-Holders for Ignition and Flame Stabilization in Scramjets: An Overview,” Journal of Propulsion and Power, Vol. 17, No. 4, 2001, pp. 869-877.

7. Yu, K., Wilson K., and Schadow, K., “Effect of Flame-Holding Cavities on Supersonic-Combustion Performance,” Journal of Propulsion and Power, Vol. 17, No. 6, 2001, pp. 1287-1295.

8. Gruber, M., Baurle, R., Mathur, T., and Hsu, K.-Y., “Fundamental Studies of Cavity-Based Flameholder Concepts for Supersonic Combustors,” Journal of Propulsion and Power, Vol. 17, No. 1, 2001, pp. 146-153.

9. Gruber, M., Donbar, J., Carter, C., and Hsu, K.-Y., “Mixing and Combustion Studies Using Cavity-Based Flameholders in a Supersonic Flow,” Journal of Propulsion and Power, Vol. 20, No. 5, 2004, pp. 769-778.

10. Lin, K.-C., Tam, C.-J., Boxx, I., Carter, C., Jackson, K., and Lindsey, M., “Flame Characteristics and Fuel Entrainment Inside a Cavity Flame Holder in a Scramjet Combustor,” AIAA Paper 2007-5381, July, 2007.

11. Lindstrom, C. D., Jackson, K. R., Williams, S., Givens, R., Bailey, W. F., Tam, C.-J., and Terry, W. F., "Multiple Line-of-Sight Absorption Spectroscopy of a Supersonic Shock Train Part I: System Design, Validation, and Mach 2 Flow Results," AIAA Journal, submitted.

12. Metacomp, http://www.metacomptech.com/index.html, 2006. 13. Singh, D. J. and Jachimowski, C. J., “Quasiglobal Reaction Model for Ethylene Combustion,” AIAA Journal,

Vol. 32, No. 1, pp. 213-215, 1993. 14. Liu, J., Tam, C.-J., Lu, T., and Law, C. K., “Simulations of Cavity Stabilized Flames in Supersonic Flows Using

Reduced Chemical Kinetic Mechanisms,” AIAA Paper 2006-4862, July, 2006. 15. Qin, Z., Lissianski, V. V., Yang, H., Gardiner, W. C., Davis, S. G., and Wang, H., Proc. Combust. Inst., Vol.

28, pp. 1663-1669.

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