Proceedings of 12 th ONR Propulsion Conference, Salt Lake City, UT, August 4-6, 1999

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1 Proceedings of 12 th ONR Propulsion Conference, Salt Lake City, UT, August 4-6, 1999 Controlling Reacting and Non-Reacting Compressible Flows using Counterflow P.J. Strykowski, D.J. Forliti, R.D. Gillgrist University of Minnesota Minneapolis, Minnesota Research has examined the active control of compressible non-reacting and reacting flows using a novel countercurrent flowfield. Studies have shown that countercurrent shear can be used to successfully control the formation of coherent structures and in the process provide proportional authority of entrainment and mixing. The nature in which counterflow might be employed to control turbulent flame speed has been examined in a countercurrent-swirl combustor. The goal is to increase volumetric heat release and turndown by exploiting the mixing enhancement characteristics and high turbulent Reynolds numbers of countercurrent shear layers. Counterflow is also being considered as an active control strategy for the nonmechanical fluidic vectoring of jet exhaust. One of the principal limitations of fluidic control is the tendency for jet attachment. The work examines this mode of operation and how it can be eliminated through appropriate nozzle-collar design. Fundamental Studies of Confined Countercurrent Shear Layers Nearly three decades after the work of Liepmann & Laufer, the world of turbulent free shear flows was turned upside-down by the observation by Brown & Roshko 1 that so-called coherent structures were in large part responsible for the flow development. Brown & Roshko demonstrated that the stoichastic description of turbulent flows was incomplete and must be modified to allow for the interaction of the mean velocity field with inherent flow structures having integral scales. The mean flow characteristics of turbulent shear layers have been well documented over an extensive parameter space, and have been summarized in several studies most notably by Birch & Eggers 2 and Ho & Huerre. 3 The influence of velocity ratio and initial conditions dominate much of the literature, however systematic studies of the independent effects of density ratio and compressibility have received renewed attention, most notably due to the work of Brown & Roshko. The significance of velocity ratio (defined here as U 2 /U 1, where U 1 and U 2 are the primary and secondary streams, respectively) cannot be underestimated. This is, in part, due to the strong dependence of shear layer growth rate on U 2 /U 1, but also because the velocity ratio plays an important role in many shear flow control strategies. However, with the exception of some isolated experiments reported in Abramovich, 4 the literature has traditionally addressed turbulent shear layers whose streams are comprised of fluids traveling in the same direction relative to a stationary disturbance source such as a splitter plate. Turbulent shear layers having velocity ratios less than zero describe an important class of flows, namely those dominated by separation and flow reversal. This class of countercurrent turbulent shear layers constitutes both a challenging experimental problem as well as an important practical one, and will be the focus of the present investigation for both reacting and non-reacting flows. Facilities Two separate facilities were used to conduct the experiments described here. The first, known as a countercurrent-swirl combustor, was discussed in detail at the 1998 ONR Propulsion Meeting. The second was designed to examine a spatially developing countercurrent shear layer using particle image velocimetry (PIV) and will be highlighted here; a schematic of the facility is shown in Fig. 1. The primary stream U 1 was driven through standard flow conditioning hardware before delivering a nominally uniform stream to the test section at velocities up to 18 m/sec. The secondary counterflowing stream, labeled U 2 in Fig.1, was developed using a vacuum system connected to an adjacent plenum chamber. The secondary flow velocity was controlled using a dual valve assembly with one valve upstream of the compressor intake and the other one downstream of the plenum chamber. The test section consisted primarily of four Plexiglas side walls extending up to 3 cm from the jet lip. The primary flow exit consisted of a maximum opening (adjustable) of 1 cm by 5 cm bordering the secondary flow opening of the same size. Side walls adjacent to the primary and secondary flows were adjustable to compensate for pressure gradients as well as provide sufficient curvature for thrust vector control to be discussed later in this Work performed under ONR contract N

2 report. The upper wall adjacent to the secondary stream was also removable allowing free entrainment of air when studying single-steam shear layers and jets. Olive oil droplets generated with a Laskin nozzle recorded with a Kodak Megaplus camera (132x135 pixels). Synchronization of the two lasers, image shifter, and camera was accomplished with a TSI LaserPulse synchronizer. To Vacuum System Suction Chamber Honeycomb Upper Flexible Wall U 2 U 2 Jet Facility U 1 U 1 Lower Flexible Wall Figure 1: Schematic of countercurrent wind tunnel facility. aerosol generator are used as the light scattering seed particles. The details of the Laskin nozzle are given in Gillgrist. 5 Particle size measurement studies (Gerbig and Keady, 6 and Crosswy 7 ) have shown that this type of aerosol generator tends to produce polydisperse particle distributions predominately in the submicron diameter range. As described in Westerweel, 8 it is important that PIV images do not contain gradients of seed concentration. The seed flow rates to the primary and secondary streams were adjusted such that no seed gradients were observed. Because of the homogeneous seeding, the image does not present any qualitative information about the flow, i.e. no visualization of flow structure is present. The lower wall was also equipped with a quartz window for introduction of the laser sheet. Two aligned Continuum Surelite I-1 lasers (each laser is capable of 2 mj/pulse at a wavelength of 532 nm) were used to illuminate the test section. The laser beams are transformed into a thin light sheet using a cylindrical and spherical lens of focal lengths of -5mm and 1mm, respectively. The lasers and optics were configured such that the beam waist was located below the image capture region. This produced a light sheet in the test section which was thin enough to generate intense particle images, yet thick enough to reduce the loss of particle image pairs due to displacements normal to the plane of the light sheet. A spinning mirror image shifting system was used to generate spatially uniform particle displacement images. The shifter is also used in experiments to resolve the directional ambiguity, which is a consequence of the autocorrelation mode used in data processing. The digital images were Figure 2: Velocity-vector field and iso-u' (% of U 1 ) contours for U 2 =. Images extend over a region of x/h of 2, and y/h of 1.2; the splitter plate location is indicated by the arrow. The time-averaged velocity-vector field obtained in the absence of imposed counterflow, and the corresponding contours of the rms u-velocity component are shown in Fig. 2 (a minimum of 2 images where captured to ensure convergence of turbulence quantities.) Viewing the velocity-vector field from a small angle with respect to the page, the observer can see the tradition uniform core flow and a very slight reverse flow on the low-speed side of the shear layer, owing to the influence of the upper wall on the natural entrainment field. The corresponding turbulence contours indicate intensity levels up to approximately 15% of the mean velocity. When the vacuum pump is activated, the secondary counterflow is established as seen in the time-average velocity vectors of Fig. 3. The velocity ratio of U 2 /U leads to rms turbulence levels of over twice the magnitude seen in Fig. 2. The contours of Fig. 3 indicate u-fluctuations of over 3% (a maximum level of ~4%). A comparison of the iso-contours indicates that the countercurrent shear layer is not only more disturbed, but diffused over a considerably larger spatial extent. Although the time-averaged behavior is useful for defining the global parameters affecting the shear 5% 1% 15% 15% 1% 5%

3 layer development, the principal advantage of PIV is the ability to examine instantaneous flow features. Figure 4 shows independent snapshots of the flow; the upper traces are without imposed counterflow (i.e. natural entrainment); the middle traces are for U 2 /U ; the lower traces are for U 2 /U significant vortical motion which will lead t o concomitant transport in mass and momentum. U 2 /U 1 = U 2 /U 1 = -.15 ~ 4% 15% 2% 25% 3% 3% 2% 15% 1% 5% Figure 3: Velocity-vector field and iso-u' contours for U 2 /U 1 = -.3. Images extend over a region of x/h of 2, and y/h of 1.2. As counterflow is increased, the scale of the resulting disturbances increases considerably. Although the time-averaged velocity field appears well behaved in Fig. 3, the instantaneous flowfield is clearly void of a defined potential core in the secondary stream. The formation of large structures in the flow is undoubtedly caused by high spatial amplification rates in the shear layer. This is, in part, due to the increased shear U, but more critically the reduced average convection velocity in the flow which should scale as (U 1 +U 2 )/2. The role that global instabilities may play in the flow development also cannot be ruled out, but must be considered geometry specific. Instantaneous vorticity computed from the PIV data is presented in Fig. 5. Grayscale levels indicate two important features of the countercurrent shear layer. First, the peak vorticity in the countercurrent shear layer is not significantly higher than that in the single-stream shear layer, despite the increased average shear. Second, the transport of moderate vorticity is seen throughout the countercurrent shear layer. Hence the flow structures accompany U 2 /U 1 = -.3 Figure 4: Instantaneous velocity-vector fields showing flow structure with increasing counterflow. U 2 /U 1 = U 2 /U 1 = -.3 Figure 5: Instantaneous iso-vorticity contours.

4 Flame Speed Control The long-term objective of this research is to study the applicability and potential benefits of countercurrent shear layers in practical combustion devices. The high turbulence levels that have been observed in flows with counterflow provide the motivation for applying this control technology to reacting flow systems. The control of the turbulence intensity through counterflow, as originally proposed by Strykowski and Wilcoxon, 9 may provide both proportional control of the heat release as well as high volumetric heat release rates desirable for practical combustors. The work of Lonnes 1 illustrated the potential controllability of the turbulent flame speed using the countercurrent-swirl combustor (CSC), shown schematically in Fig. 6. Flame speed control was achieved through a combination of counterflow and swirl, with the swirl having the effect of confining the tubular flame away from the combustor walls. Although control of the flame speed was established, the actual magnitude of the chamber averaged turbulent flame speed was below approximately four times the laminar flame speed. The objective of the present research is to find a combustor geometry that retains the controllability of the heat release while achieving higher turbulent flame speeds and an overall increase in the heat release rate. "Rear" Ro r Flame Sheet x L "Front" Figure 6: Cross section of countercurrent-swirl combustor. The presence of turbulence and swirl in the CSC has the effect of wrinkling and smoothing the tubular flame, respectively. The positive radial pressure gradient caused by the swirl impacts the ability of the turbulence to convolute the flame, causing a reduction in the flame surface area and consequently lowering the heat release rate (Veynante and Poinsot 11 ). Beér et al. 12 demonstrated the laminarizing effect on flames due to a swirling environment, which was required to stabilize the flame inside the swirl burner. Thus the presence of swirl in the CSC is expected to limit the performance to lower turbulent flame speeds. The mean radial pressure gradient for a turbulent swirling axisymmetric flow is given by the following: ( ) 2 2 ( ) p ρ u + = θ uθ ρ + r r r u 2 u 2 r r 2 where p is the pressure, ρ is the mixed reactants density, u is the velocity with subscripts of r and θ for radial and azimuthal velocities, respectively. This pressure gradient can be minimized through the removal of the mean swirl. The present ideology for obtaining higher turbulent flame speeds is to remove the mean swirl (or to reduce to the greatest extent possible while maintaining a stable flame), and minimize any significant positive pressure gradient effects from the remaining terms in the above equation. A parametric study of the CSC geometry will assist in defining a combustor geometry having the optimum controllability of the turbulence intensity via counterflow. In particular, it must be shown that the front and rear annular jets have strong interaction since this is the proposed counterflow mechanism for turbulence control. A single component LDA system was used t o investigate the effect of rear mass fraction (RMF) on the flow field of the CSC for isothermal flow conditions. It was hypothesized that the ability for RMF to control the turbulence inside the combustor would be qualitatively the same with and without the flame present. The LDA system incorporated a Coherent Innova 7-4 Argon Ion CW laser with Dantec and TSI optical and electrical components. The laser and LDA optics were fixed to a mill bed traverse controlled by a Centroid CNC controller. The optics were oriented to measure axial velocities. Olive oil droplets of sub-micron scale generated with a Laskins nozzle (Gillgrist 5 ) were used for this LDA study. Velocity measurements were made for the geometry and flow conditions listed in Table 1, which was representative of the conditions studied by Lonnes. 1 Parameter Value Chamber length, L 12 inches Jet separation, δ.632 inches Air mass flow (no fuel flow).3 Kg/sec RMF.1,.2,.3 Front swirl angle 68 Rear swirl angle 86 Table 1: Conditions of CSC for LDA Experiments Radial profiles of the mean and rms fluctuation axial velocities at x=.5l are shown in Figs. 7a,b for the three RMF cases studied. As can be seen, the

5 mean velocity profiles show the presence of counterflow. The mean and fluctuating velocity profiles are very similar for all three RMF cases, which would not be expected if the jet interaction was to strongly modify the mean counterflow patterns and turbulence levels. (a) Velocity (m/s) Mean Axial Velocity Profiles at x/l.5 RMF=.1 RMF=.2 RMF= The turbulence levels are in excess of 1 times the laminar flame speed for methane at an equivalence ratio of.8, which is representative of the work of Lonnes. 1 The chamber averaged turbulent flame speeds measured by Lonnes is significantly lower than that found by other researchers (e.g. Aldredge, Vauzi and Ronney 14 ) and theoretical models (e.g. Bradley 15 ) for this level of turbulence intensity. The lower turbulent flame speeds found by Lonnes are likely caused by the pressure gradient due to the high swirl, and not an effect of curvature/flame stretch. (a) Velocity (m/s) 1 5 (b) 6 5 r/r -5 RMF=.1 RMF=.2 RMF=.3 rms Velocity (m/s) RMF=.1 RMF=.2 RMF=.3 (b) RMF=.1 RMF=.2 RMF=.3 yr/r r/r r/r o Figure 7: Radial (a) mean and (b) rms u-velocity profiles at the axial center (x/l =.5) of the countercurrent-swirl combustor. Radial profiles of the axial velocity were also measured near the rear drive of the CSC (x=.1l) to observe the streamwise variation of the flow. The mean and rms fluctuation velocity profiles are shown in Figs. 8a and 8b, respectively. The mean axial velocity profiles show some dependence on RMF. The presence of the rear jet can only be detected in the mean velocity profile for RMF=.2 near r/r o ~.4. The absence of the peaks suggests a possible effect of a mismatch in the angular momentum balance between the front and rear jets, which could result in a radial inward or outward deflection of the rear jet; this trend was observed in the computational study of the CSC by Lanicek et al. 13 The rms axial velocity profiles near the rear show no trend with RMF, and the magnitudes are similar to those measured at the combustor midlength (x=.5l). rms Velocity (m/s) r/r r/r o Figure 8: Radial mean (a) and (b) rms u-velocity profiles near the rear (x/l =.1) of the countercurrent swirl combustor. The parametric study is presently continuing to study the effects of chamber length L, jet separation δ, and mean flow swirl in search of a combustor geometry which allows for turbulence control through the interaction of the two counterflowing annular jets. The study will then proceed with the optimized combustor geometry to study the flame speed control at potentially higher chamber averaged flame speeds. Thrust Vector Control and Jet Attachment The basic concept of counterflow thrust vectoring can best be illustrated by referring to the

6 sketch in Fig. 9. The primary jet exhausts from the nozzle between symmetric curved surfaces called collars. To achieve upward thrust vectoring a secondary counterflow must be established above the upper shear layer of the jet. This can be accomplished by applying a vacuum source to the gap formed between the upper curved surface and the primary nozzle. The action of counterflow in the upper shear layer gives rise to asymmetric entrainment and a cross-stream pressure gradient sufficient to vector the jet. Figure 9: Schematic of counterflow thrust vectoring nozzle. The ability of the jet to deflect in the presence of a stationary surface has been commonly referred to as the Coanda effect. As a nominally twodimensional jet issues from the nozzle, the shear between the jet and the surrounding quiescent fluid gives rise to lateral momentum transport which is accentuated by turbulent mixing of the shear layer. This process leads to momentum mixing between the two fluids and is commonly referred to as entrainment. The fluid entrained along the upper collar surface by the pumping action of the shear layer is constrained by the proximity of the collar, thus giving rise to subatmospheric pressures in that region. This differential pressure draws the jet toward the curved surface. If the entrainment rate is sufficiently high, the pressure continues to drop until the jet becomes attached to the surface. It is this change in the pressure field within the vicinity of a solid surface that is known as the Coanda effect. Due to the entrainment differential, a small degree of jet turning is possible, in principle, without a collar surface, however it is the addition of an extended collar surface that makes counterflow TVC a viable technology. First, it channels the secondary flow in parallel with the primary flow, effectively inducing a countercurrent mixing layer over the entire length of the collar. Second, it restricts the natural entrainment of the jet, intensifying the cross stream pressure gradient. Finally, it gives the pressure forces a surface over which to act. A longer collar has more surface area for the pressure forces to act on; thus a smaller pressure differential will impart an equal transverse force on the nozzle-collar assembly. In other words, with a longer collar, the same vector angle is achievable with a smaller amount of vacuum from the secondary flow system. The net effect is a reduction in the secondary pumping requirements. While the collar is vital to the efficiency of the system, it poses a potential problem that can disrupt the continuity of the operating curve. This condition, a result of the hysteretic interaction between a free jet and a wall, is one in which the jet attaches to, and reaches a stable equilibrium on, the wall. It has been observed, under certain operating conditions, 16 that the jet attaches to the collar surface during vectoring. This is unacceptable from a design standpoint as, in a flight situation, this may cause a loss of control. Furthermore, to release the jet from the collar, the differential control pressure must be reduced beyond that which was required t o cause attachment. This hysteretic behavior makes a continuous vector-control system very difficult (if not impossible) to implement. Modeling Jet Attachment The important thing to bear in mind when designing a counterflow TVC system is that wall attachment can only occur if equilibrium can be sustained. This means that the entrainment mechanisms within the jet must be able to sustain the low pressure necessary to hold the jet attached to the wall. The shorter the collar, the sharper the jet must turn in order to attach; thus low pressure in the secondary stream will be required. Furthermore, with a short collar, the shear layer has less contact with the bubble region. This makes it more difficult for the pumping mechanism within the jet to generate the low pressure required to hold itself to the wall. Consequently, if the collar is sufficiently short, attachment will not occur. There is a unique equilibrium location where the attaching streamline intersects the collar. This location depends on the secondary gap width, the mixing dynamics of the shear layer, and the amount of secondary flow leaving (or entering) the recirculation zone. A large gap places the collar surface farther away from the jet, resulting in a longer attachment length. Enhanced shear layer growth rates mean the jet is more effective in sustaining the low pressure within the bubble region. Likewise, counterflow being drawn from the bubble region by a secondary pump assists

7 the jet in sustaining the low pressure necessary t o hold itself to the wall. These two latter conditions enable the jet to turn with a tighter radius of curvature, resulting in a shorter attachment length. If the collar is longer than this equilibrium attachment length, Latt, the jet will merely attach at the designated location, and follow the contour of the wall until boundary layer separation occurs. However, if the collar length is shorter than the equilibrium attachment length, i.e. L < Latt, the jet will not attach to the wall. When designing a collar, it is important to be able to estimate the attachment length based on the intended operating conditions. Experimental determination of jet attachment was made by instrumenting the collar with surface pressure taps and using PIV. The streamline pattern in Fig. 1 indicates the nature of the attachment process and the corresponding surface pressure distribution. The subatmospheric pressure developed in the separation bubble gives rise to a relatively constant pressure in the nozzle near field, followed downstream by jet attachment as indicated by the local minimum in the normalized pressure distribution. The attachment location, Latt, was assumed to be indicated by this minimum. Timeaveraged PIV images (not shown here; see Gillgrist 5 ) were used to provide detail of the flow field and t o corroborate the attachment location using the pressure distributions. Truncation of the collar downstream of attachment had virtually no influence in the attachment location indicating that the phenomenon was dominated by the shear layer entrainment in the jet near field. Once the collar was truncated at a length less than Latt, the attachment was avoided. A model was developed to predict jet attachment as a function of nozzle-collar geometry as well as jet Mach number and temperature. Conservation of mass and momentum was applied to a control volume located downstream of the jet exit. The attaching streamline is determined by the geometry (e.g. gap height, collar curvature) and the rate at which the shear layer entrains the surrounding fluid. The model assumes that the mixing characteristics of the shear layer are determined to leading order by the velocity ratio, density ratio, and convective Mach number, and that the effect of curvature on shear layer mixing is relatively unimportant. The model input also required specification of the secondary mass flow rate, which can be either counterflowing or coflowing to the primary jet. The attachment location is indicated by the angle α measured along the collar, where α = at x=. Experiment-model comparisons of the jet attachment location as a function of collar displacement (gap) and collar curvature are shown in Fig. 11. Despite the simplicity of the model, the attachment locations are captured quite well. The curves can be used for nozzle-collar design t o eliminate jet attachment, as long as the actual collar is shorter than the model predictions. Figure 1: Streamline pattern (upper) and collar pressure distribution (lower) during jet attachment. Figure 11: Jet attachment location for the case without counterflow; G is gap height, R is collar radius of curvature. Designing a counterflow TVC system for aircraft or missile propulsion requires that a collar geometry be found which is capable of achieving the required thrust vector angle while minimizing external drag, and secondary mass flow pumping demands. All designs must also provide attachment-free operation over the entire operating domain of the vehicle. The design process is set by the independent choice of jet exit Mach number and exhaust temperature, which in turn determine the shear layer entrainment

8 characteristics through density ratio and convective Mach number. These parameters are provided as input to the jet attachment model and provide the equilibrium attachment length Latt/H corresponding to the level of secondary flow present. References 1 Brown, G.L., Roshko, A., "On Density Effects and Large Scale Structures in Turbulent Mixing Layers," J. Fluid Mech., Vol. 64, 1974, pp Birch, S.F., Eggers, J.M. "A Critical Review of the Experimental Data for Developed Free Turbulent Shear Layers," NASA SP-321, 1973, pp Ho, C.-M., Huerre, P., "Perturbed Free Shear Layers," Ann. Rev. Fluid Mech., Vol. 16, 1984, pp Abramovich, G.N., "The Theory of Turbulent Jets," MIT Press, Cambridge, MA, Gillgrist, R.D., "A Fundamental Study of Thrust Vector Control using Counterflow," M.S. Thesis, University of Minnesota, March, Gerbig, F.T., Keady, P.B., "Size Distributions of Test Aerosols from a Laskin Nozzle," Microcontamination, July 1985, pp Crosswy, F.L., "Particle Size Distributions of Several Commonly used Seeding Aerosols," NASA CP-2393, 1985, pp Westerweel, J., "Fundamentals of Digital Particle Image Velocimetry," Meas. Sci. Tech., Vol. 8, 1997, pp Strykowski, P.J., Wilcoxon, R.K., "Mixing Enhancement Due to Global Oscillations in Jets with Annular Counterflow," AIAA J., Vol. 31, 1993, pp Lonnes, S.B., "Flame Speed Control using a Countercurrent Swirl Combustor," Ph.D. Thesis, University of Minnesota, May Veynante D., Poinsot T., Effects of Pressure Gradients on Turbulent Premixed Flames, J. Fluid Mech., Vol. 353, 1997, pp Beer J.M., Chigier, N.A., Davies, T.W., Bassindale, K., Laminarization of Turbulent Flames in Rotating Environments, Combustion and Flame, Vol. 16, 1972, pp Lanicek, L., Alizadeh, S., Jicha, M., Strykowski, P.J., "Enhancing Understanding of the Operation of the Dynamic Containment Combustor through CFD Modeling," Proceedings of 5 th ASME/JSME Thermal Engineering Joint Conference, March 15-19, 1999, San Diego. 14 Aldredge, R.C., Baizi, V., Ronney, P.D., Premixed-flame Propagation in Turbulent Taylor- Couette Flow, Combustion and Flame, Vol. 115, 1998, pp Bradley, D., How Fast Can we Burn?, 24 th Symposium (International) on Combustion, 1992, pp Van der Veer, M.R., Strykowski, P.J., "Counterflow Thrust Vector Control of Subsonic Jets: Continuous and Bistable Regimes," J. Propulsion & Power, Vol. 13, 1997, pp Biographical Sketches Dr. Paul Strykowski is a Professor of Mechanical Engineering at the University of Minnesota, Minneapolis, and has held appointments as an Adjunct Professor in Mechanical Engineering at Florida A&M University, Tallahassee. He received his B.S. degree from the University of Wisconsin (1982), and M.S. (1983), M.Phil (1985) and Ph.D. (1986) degrees from Yale University, all in Mechanical Engineering. Research interests include fundamental and applied aspects of non-reacting and reacting shear flow control. Mr. David Forliti is a Ph.D. candidate in the department of Mechanical Engineering at the University of Minnesota. He received his B.S. degree from the University of Minnesota and his M.S. degree from Florida State University, both in Mechanical Engineering. Mr. Robert Gillgrist is a M.S. candidate in the department of Mechanical Engineering at the University of Minnesota. He received his B.S. degree from Purdue University.