Separation Control of Axial Compressor Cascade by Fluidic-Based Excitations

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Xinqian Zheng

Yangjun Zhange-mail: [email protected]

State Key Laboratory of Automotive Safety andEnergy,

Tsinghua University,Beijing 100084, China

Weidong Xing

Junyue Zhang

National Key Laboratory of Diesel EngineTurbocharging Technology,

Datong,Shanxi 037036, China

Separation Control of AxialCompressor Cascade byFluidic-Based ExcitationsFlow separation control was investigated on a compressor cascade using three types offluidic-based excitations: steady suction, steady blowing, and synthetic jet. By solvingunsteady Reynolds–averaged Navier–Stokes equations, the effect of excitation param-eters (amplitude, angle, and location) on performance was addressed. The results showthat the separated flow can be controlled by the fluidic-based actuators effectively and thetime-averaged performance of the flow field can be improved remarkably. Generally, theimprovement can be enhanced when the amplitude of excitation is increased. The optimaldirection varies with each type of excitations and is related to physical mechanismsunderlying the separation control. For two types of steady excitations, the most effectivejet location is at a distance upstream of the time-averaged separation point and thesynthetic jet is just at the separation point. �DOI: 10.1115/1.4002407�

IntroductionIn the modern aero-engine design, the tendency toward weight

eduction and pressure ratio enhancement leads to stronger ad-erse pressure gradient in axial compressors, making unsteadyow separation appears easily. The separated flow constrains bothfficiency and pressure ratio. Therefore, it is increasingly urgent toddress the control of flow separation in axial compressor.

Almost all the targets of separation control focus on accelerat-ng the low momentum fluid in boundary layer on adverse pres-ure gradient, notwithstanding the methods are different. Thereas been extensive work in research and development on separa-ion prevention �1,2�. These methods include surface-based actua-ors, e.g., vortex generators, riblets, slotted flaps, and fluidic-based

ethods, e.g., steady suction, steady blowing, and synthetic jet.Most of the flow control research was focused on the single

irfoil �3,4�. Several research groups have begun to investigate thepplication of flow control to compressor blades by using fluidic-ased methods. Merchant et al. �5,6� designed and tested an aspi-ated compressor stage, which achieved a total pressure ratio of.4 at a tip speed of 457.2 m/s with a suction flow rate of 3.5% ofhe inlet mass flow. Schuler et al. �7� designed an aspirated fantage for pressure ratio of 1.6 at a low tangential tip Ma of 0.7.he mass flow was removed through slots on the suction surfacef the rotor/stator and the rotor shroud and stator hub. Steadylowing was carried out on a highly loaded compressor cascade8�. Both the profile and the corner boundary layer could be con-rolled, and the performance was improved. Culley et al. �9� con-rolled the stator vane suction surface separation in a low-speed

ultistage compressor using steady and unsteady injection on theuction surface. These results suggested that air injection from theuction surface was relatively ineffective in mitigating endwalloss. Furthermore, a number of papers investigating an approachor tip critical instabilities are focused on injection �10,11�.

The referred research contributed to the understanding of floweparation control in axial compressors �12�. However, to the au-hors’ knowledge, little study has been performed on separationontrol comparatively by using various types of fluidic-basedethods in axial compressors. In the present study, three types ofuidic-based methods, steady suction �SS�, steady blowing �SB�,

Contributed by the International Gas Turbine Institute �IGTI� of ASME for pub-ication in the JOURNAL OF TURBOMACHINERY. Manuscript received October 3, 2007;nal manuscript revised February 12, 2010; published online April 21, 2011. Editor:
avid Wisler.
ournal of Turbomachinery Copyright © 20

ded 21 Jun 2011 to 166.111.50.173. Redistribution subject to ASM

and synthetic jet �SJ�, were investigated regarding their ability tocontrol the separation inside an axial compressor cascade by solv-ing the unsteady Navier–Stokes equations. In view of complexflows in axial compressors, two-dimensional case was consideredonly in this paper. Specifically, flow conditions were fixed atMa=0.5 and i=10 deg.

2 Numerical Methods

2.1 Discretization Methods. The code used in this simulationwas based on a 2D unsteady compressible finite difference schemeto solve the Reynolds-averaged Navier–Stokes equations in con-servative formulation. A third-order TVD Runge–Kutta time-marching scheme was used for time integration because time ac-curacy was critical for this simulation. It was generally recognizedthat high-order spatial schemes were recommended especially fortime-dependent simulations. A fifth-order accurate, generalizedcompact scheme �13� was used for spatial difference of inviscidfluxes and a sixth-order compact finite difference scheme �14� forviscous fluxes. For turbulence model, the standard algebraicmodel of Baldwin and Lomax was adopted �15�.

2.2 Grid and Initial Value. In the present paper, an H-typegrid was adopted, which was generated by solving relevant ellipticdifferential equation. The grid influences were analyzed to makesure that the computation results were independent of spatial dis-cretization. The influences of different factors were analyzed, suchas the forward and backward extensions of the computational do-main, the number of grid points, and the local densification of gridpoints. After a thorough consideration of compromise betweencomputation work and computation precision, the grid adopted inthe present paper was as follows. The computational domain cov-ers six chord lengths, one chord length before the cascade andfour chord lengths behind the cascade; the number of grid pointswas 321 �axial direction��65 �circumferential direction�. Forimproving computation precision, grid resolution was refined lo-cally near the leading and trailing edges and the cascade walls, inparticular, within the boundary layers, the first level grid pointsabove the wall were located at y1

+�10.The problem how to select initial values had also been studied.

Commercial software NUMECA was first used to obtain a con-verged steady solution, which was then taken as the initial flow-field for our unsteady simulation. This greatly accelerated conver-gence of the unsteady simulation. For different initial flow-fields,
the unsteady simulations could invariably lead to periodic or qua-
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iperiodic converged solutions, and the influences on the time-veraged performances were also small. The influence of the ini-ial values lies mainly in the converging speed. Data wereecorded only when the computation was converged to a periodicr quasiperiodic solution.

Because for this kind of massively separated and unsteady flowhe initial transient stage is quite random and has a profound effectn the detailed fluctuation history of the flow, it is impractical toeach an exact grid and initial value independence. Overall inac-uracy due to the grid and initial solution was smaller than 0.5%.

2.3 Boundary Conditions. The present paper adopted the ap-roach of modifying the boundary conditions on the profile sur-ace to simulate the fluidic-based actuators. The following veloc-ty at the excitation location was imposed:

uwall = Vex cos�2�fet� · cos �

vwall = Vex cos�2�fet� · sin � �1�

n which � is the excitation direction angle, i.e., �=0 deg meanshe excitation is tangential to the suction surface and pointing tohe trailing edge, and �=90 deg means excitation is normal to theuction surface. Equation �1� denotes the method of synthetic jethen fe�0; fe=0 means the methods of steady excitations, i.e.,

teady suction ��0� and steady blowing ��0�, respectively.Steady boundary conditions were given at the inlet and outlet

oundaries. Total temperature, total pressure, and incidence wereiven at the inlet, whereas only the static pressure was fixed at theutlet while velocity and density could be derived by extrapola-ion. Nonslip, impermeability, and adiabatic boundary conditionsere imposed on all the profile surface points except for the lo-

ig. 1 Comparisons of static pressure distribution on profileurface between the measurements and the simulation results

Fig. 2 Cascade geometry an

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cations of blowing and suction. Periodic conditions were imposedon the boundaries of the extended domains before and behind thecascade.

2.4 Validation of the Program Reliability. Validation of thesimulation’s reliability is of key importance in numerical simula-tion. Comparisons are shown in Fig. 1 for static pressure distribu-tion on the profile surface between the measurements obtained inthe planar cascade experiments and the relevant simulation results�time-averaged values�. Overall, good agreement is obtained be-tween the experimental and the simulation results, indicating thereliability of the simulation setup. However, there are some dif-ferences, e.g., the simulation results are somewhat smaller thanmeasurements at the points behind the position of 50% chord. Themain reasons can be explained as follows: �1� Simplified simula-tion model—even though planar cascades are nonrotating partswith weak 3D effects, however, they are actually 3D straight-blade cascades with end-wall, whereas the simplified simulationmodel is a strictly 2D one. Hence, the determinative conditions ofphysics and mathematics have differences between numericalsimulations and experiments. �2� B-L model was adopted as tur-bulence model. This model is widely used in engineering applica-tions; however, practice shows that it is always some differencesbetween the simulation and experiment results.

3 Simulation ResultsThe modeled compressor cascade geometry is as follows: blade

chord C=80.6 mm, solidity �=1.5, inlet camber angle �1K=41.6 deg, exit camber angle �2K=1.23 deg, and stagger angle=19.88 deg. The imposed fluidic-based excitations were as-sumed to be completely 2D and the jet slot width is about 5% ofthe chord length. Figure 2 shows the main parameters involved.

The total loss coefficient is chosen as the indication of perfor-mance. Taking into account the mass and kinetic energy of the jetflow, the total pressure loss coefficient is defined as

=min · pin

� + mex · pex� − �min + mex� · pout

�

min · �pin� − pin� + mex · �pex

� − pex��2�

where pin� and pout

� are the mass-averaged total pressure of inletand outlet of computational domain, respectively. pin is the mass-averaged static pressure of inlet section. In Eq. �2�, the denomi-nator means the total kinetic energy entering the channel, and thenumerator represents the total pressure loss.

According to Eq. �2�, the total pressure loss coefficient withoutfluidic-based excitation is 0=0.237 and the separation point onthe suction side is about 26.2% of chord length from the leadingedge. The time-averaged performance is evaluated by using nor-malized total pressure loss coefficient / 0.

d excitations parameters

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For the three types of excitation methods, the jetted mass flows as follows:

mex��0 steady blowing

=0 synthetic jet or unexcitation

�0 steady suction� �3�

3.1 Steady Suction

3.1.1 Effect of Amplitude. For a given suction angle and loca-ion, Fig. 3 shows the relation between suction amplitude and lossoefficient. Results give that improvement of performance is di-ectly related to the suction amplitude. Even a small increase ofhe suction amplitude will result in an immense improvement in

he performance. As shown in Fig. 3, when A=10% �m=0.73%�,oss coefficient can be reduced by 31.2%. Therefore, for steadyuction, the separated flow could be under effective control with-ut high suction amplitude.

3.1.2 Effect of Angle. Figure 4 gives the variation of loss co-

fficient as suction angle varies at two locations of l=26.2%time-averaged separation point� and 29.9%. The results indicate

hat, for a given suction amplitude �A=8%�, the optimal efficiencyan be obtained at �=−90 deg, and other angles are symmetricbout the center of �90 deg. It can be explained by the fact that

ig. 3 The loss coefficient for different relative steady suctionmplitude „�=−90 deg, l=15.6%…

ig. 4 The loss coefficient for different steady suction direc-¯
ion „A=8%…
ournal of Turbomachinery


the suction mass flow rate reaches its maximum at the suctionangle of �90 deg when other suction parameters are fixed. Theresults can be proved herewith that a certain amount of the suctionflow rate is the prerequisite of obtaining a positive suction effect.

3.1.3 Effect of Location. Figure 5 gives the relation betweendifferent suction locations and total pressure loss coefficients.

Loss coefficient can be reduced by about 20% when 7%� l�30%. However, the improvement of performance increasesslightly when suction location is beyond 60% chord length.

Particularly, when steady suction is imposed at the location of

15.6% chord length instead of the separation point �l=26.2%�, theoptimal effect could be obtained with a reduction by 28.3%. Atthis location, separation does not yet occur and the flow is stillattached. However, the low momentum has been accumulated atthis location. It is at this location that the flow could keep attachedif low momentum fluid is sucked or high streamwise momentumfluid is injected, thereby delay or restrain the downstream separa-tion.

3.2 Steady Blowing

3.2.1 Effect of Amplitude. Three locations were selected on the

suction surface, one near to the leading edge �l=1.75%�, one at a

distance upstream of the time-averaged separation point �l=15.6%�, and another near to the separation point �l=26.2%�. Theblowing angle is fixed, and the blowing flow rate is controlled tom�0.25%. The results at the three locations are given by Fig. 6.The loss coefficient of the flowfield generally decreases with anincrease of the blowing amplitude.

3.2.2 Effect of Angle. Figure 7 shows the relation betweenblowing angle and cascade performance. The effectiveness ofsteady blowing control is in a wide range. For a given blowingamplitude, the blowing flow rate will increase with an augmentedangle, and blockage is likely to appear, which leads to decreasethe static pressure ratio. Blockage is present when ��30 deg andthe results is not given accordingly.

The effectiveness of tangential blowing ��30 deg� is similarto that of the normal suction for the reason that when blowingtangentially, i.e., blowing streamwise, high momentum fluid canfully join with the boundary layer flow, making low momentumfluid in boundary layer accelerate on the one hand, and avoidserious disturbance to main flow on the other hand. At the same

Fig. 5 Loss coefficient for different steady suction location„A=8%, �=−90 deg…

time, for a given slot width, tangential blowing brings high mo-

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entum and low mass flow rate injection, the blockage broughtith it will not decrease the static pressure ratio while diminishing

he separated region.

3.2.3 Effect of Location. Figure 8 shows the effect of the lo-ation of excitation on the performance. Again, loss coefficient

an be reduced by about 20% when 7%� l�30%. Particularly,hen steady suction is imposed at the location of 15.6% chord

ength instead of the separation point �l=26.2%�, the optimal ef-ect can be obtained with a reduction by 29.1%. In addition, seenrom Figs. 5 and 8, it is surprised that the similar results arebtained with the excitations by steady suction and steady blow-ng.

3.3 Synthetic Jet. Steady suction and blowing are belong toteady methods. Synthetic jet, the characteristic of zero net massux and nonzero momentum flux, is a typical unsteady method.he main difference between steady and unsteady methods is the

mportance of excitation frequency. As this paper focuses on theffect of amplitude, angle, and location, the role of excitationrequency can be referenced in Ref. �16�. When the excitationrequency was approximately equal to the characteristic frequency

f vortex shedding, i.e., f =1, the optimal effect of unsteady exci-ation on separation flows can be obtained. Therefore, relative

xcitation frequency is chosen at f =1.

ig. 6 Loss coefficient for different relative steady blowingmplitude at three locations „�=2 deg…

ig. 7 Loss coefficient for steady blowing direction „A=90%,
=26.2%…
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3.3.1 Effect of Amplitude. Figure 9 shows the effect of excita-tion amplitude on the loss coefficient. It is demonstrated that theperiodic excitation has a negligible effect on the flow dynamicswhen the excitation amplitude is relative small. Therefore, posi-tive results obtained must satisfy the condition that the relative

excitation amplitude is above a certain threshold value �A�10%�. Basically, the aerodynamic performance increases mono-

tonically with excitation amplitude up to about A=35%. If the

excitation amplitude goes beyond the limit �A=35%�, the effect ofperiodic excitation is weakened slightly.

3.3.2 Effect of Angle. Figure 10 shows the effect of excitationangle on the loss coefficient. The same best positive excitationeffects are obtained in the range of 40 deg��140 deg. Be-yond this range, the excitation effects decrease sharply. However,up to the extreme case of �=0 deg or �=180 deg, i.e., the exci-tation direction tangent to the suction surface, positive effectscould be still obtained. There is another typical feature, the curvesin Fig. 10 almost take the normal to the suction surface of �=90 deg as their symmetry axis, their left parts and right parts aregenerally symmetric. Therefore, the best positive excitation ef-fects are obtained when the excitation direction is normal to thesuction surface and the positive effects remain invariability in arather wide range of �.

The nature of synthetic jet is periodic suction and blowing, and

Fig. 8 Loss coefficient for different steady blowing location„A=90%, �=2 deg…

Fig. 9 Loss coefficient for different synthetic jet amplitude „f¯
=1, �=90 deg, l=26.2%…


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combination of steady suction and pulsed jet in temporal char-cteristic. Not only can it suck out the low momentum fluid inoundary layer, but also it can facilitate the interaction of un-teady vortices. Therefore, synthetic jet is not sensitive to the jetngle.

3.3.3 Effect of Location. Figure 11 shows the loss coefficientor different excitation locations. The best results were obtainedhen synthetic jets are excited just at the time-averaged separa-

ion point. If the excitation location moves backward or forwardurther from the time-averaged separation point, the positive effectf excitation is reduced slightly. Generally, the positive results arebtained in a wide range of excitation location, which is verymportant for engineering application because the separation points varied with the flow conditions.

It is speculated that the boundary layer close to the separationoint is sensitive to disturbance. Just around this specific localegion, the boundary layer is receptive to unsteady excitation. Theasic steady flow has a high velocity component upward at theocal region near separation point, which will amplify the effect ofny nonlinear streaming and thus make local oscillations an effec-ive actuator in delaying separation on the blade. This amplifica-ion mechanism, which cannot be found in attached flows, willnhance the primary entrainment and rolling-up coalescence in-erent in boundary layer, resulting in reduced separation region.

ig. 10 Loss coefficient for different synthetic jet direction „f1, l=26.2%, A=40%…

ig. 11 Loss coefficient versus synthetic jet location „f=1, A
30%, �=90 deg…


4 Flow Field StructureThe flow field without excitation, i.e., the baseline flow, is

shown in Fig. 12. It is obvious that a strong velocity gradientdivides the flowfield into main and separated flow. There is asequence of vortices generating, developing, and shedding in theseparated flow. The presence of a large scale separated vortexdeteriorates the diffusion capacity, and loss increases. Figures 13and 14 are the flow structures imposed by steady suction andsteady blowing, respectively. Compared with Fig. 12, the strongvelocity gradient weakens, and the flow velocity increases. Theseparated region is reduced remarkably, the scale of separatedvortex diminishes, and the streamlines become smoother. Theflowfield structure by synthetic jet can be referenced in Ref. �16�.

5 Concluding RemarksFor this paper, investigations on separation control in an axial

compressor cascade have been carried out using three types offluidic-based excitations: steady suction, steady blowing, and syn-thetic jet. The focus is on the influence of time-averaged aerody-namic performances when the amplitude, angle, and location ofexcitation are varied. The results can be summarized as follows:

Fig. 12 Instantaneous contour plots for Ma, and total pressureplus streamlines. Baseline flow.

�1� The effectiveness of above three type’s excitation varies

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significantly with the excitation parameters �e.g., jet angle,location, amplitude, or injected flow rate�. As a whole, thepositive effect is more remarkable when the excitation am-plitude is increased.

�2� For a given suction amplitude and location, normal suction��=−90 deg� can maximize the suction flow rate and ob-tain the best effect. The steady blowing energizes the lowmomentum region directly, so a tangential angle is effec-tive. Synthetic jet is not sensitive to the jet angle and thepositive effects remain invariability in a rather wide rangeabout 40 deg��140 deg.

�3� The optimal location of steady excitation �i.e., steady suc-tion and steady blowing� and unsteady excitation �syntheticjet� are different. The optimal location of steady excitation

is at a distance �l=15.6%� upstream of the time-averaged

separation point �l=26.2%�. The optimal location of un-steady excitation is just at the time-averaged separationpoint and positive results can be obtained in a wide range

ig. 13 Instantaneous contour plots for Ma, and vorticity plustreamlines. Steady suction „A=8%, �=90 deg, l=15.6%….

of excitation locations.

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AcknowledgmentThis research was supported by the National Natural Science

Foundation of China �Project No. 50806040�. The authors aredeeply grateful to Zhou Xiaobo for her efforts to make the simu-lation possible. X.Z. would also like to thank Joern Huentelerfrom RWTH Aachen University for his great effort to correct theEnglish grammar.

NomenclatureA � relative excitation amplitude A=Vex /c

C � chord of the cascade bladec � free stream velocity

fshed � characteristic frequency of vortex sheddingfe � excitation frequency

f � relative excitation frequency f = fe / fshedl � distance between L.E. and excitation location

l � normalized distance between L.E. and excita-

tion location l= l /Cm ¯

Fig. 14 Instantaneous contour plots for Ma, and vorticity plusstreamlines. Steady blowing for „A=75%, �=2 deg, l=15.6%….

� time-averaged excitation flow rate m= �mex /min�



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Ma � Mach numberp � static pressure

p� � total pressuret � time, blade pitch

uwall � velocity in wall-tangent directionvwall � velocity in wall-normal directionVex � excitation amplitudex ,y � Cartesian coordinate in axial direction, Carte-

sian coordinate in circumferential direction� � excitation angle with respect to the direction

tangent to the blade suction surface � total loss coefficient

ubscriptsex � excitationsin � inlet

out � outlet

eferences�1� Greenblatt, D., and Wygnanski, I. J., 2000, “The Control of Flow Separation

by Periodic Excitation,” Prog. Aerosp. Sci., 36�7�, pp. 487–545.�2� Seifert, A., Bachar, T., Koss, D., Shepshelovich, M., and Wygnanski, I., 1993,

“Oscillatory Blowing: A Tool to Delay Boundary-Layer Separation,” AIAA J.,31�11�, pp. 2052–2060.

�3� Liu, Y., Sankar, L. N., Englar, R. J., Ahuja, K. K., and Gaeta, R., 2004,“Computational Evaluation of the Steady and Pulsed Jet Effects on the Perfor-mance of a Circulation Control Wing Section,” AIAA Paper No. 2004-56.

�4� Wu, J. Z., Lu, X. Y., Denny, A. G., Fan, M., and Wu, J. M., 1998, “Post-Stall
Flow Control on an Airfoil by Local Unsteady Forcing,” J. Fluid Mech., 371,


pp. 21–58.�5� Merchant, A. A., Drela, M., Kerrebrock, J. L., Adamczyk, J. J., and Celestina,

M., 2000, “Aerodynamic Design and Analysis of A High Pressure Ratio Aspi-rated Compressor Stage,” ASME Paper No. 2000-GT-619.

�6� Merchant, A., Kerrebrock, J. L., Adamczyk, J. J., and Brqunscheidel, E., 2004,“Experimental Investigation of A High Pressur Ratio Aspirated Fan Stage,”ASME Paper No. GT2004-53679.

�7� Schuler, B. J., Kerrebrock, J. L., Merchant, A. A., and Drela, M., 2000, “De-sign, Analysis, Fabrication and Test of an Aspirated Fan Stage,” ASME PaperNo. 2000-GT-618.

�8� Sturm, W., Scheugenpflug, H., and Fottner, L., 1992, “Performance Improve-ments of Compressor Cascades by Controlling the Profile and Sidewall Bound-ary Layers,” ASME J. Turbomach., 114, pp. 477–486.

�9� Culley, D. E., Bright, M. M., Prahst, P. S., and Strazisar, A. J., 2004, “ActiveFlow Separation Control of a Stator Vane Using Embedded Injection in aMultistage Compressor Experiment,” ASME J. Turbomach., 126, pp. 24–34.

�10� Suder, K. L., Hathaway, M. D., Thorp, S. A., Strazisar, A. J., and Bright, M.B., 2001, “Compressor Stability Enhancement Using Discrete Tip Injection,”ASME J. Turbomach., 123, pp. 14–23.

�11� Bae, J. W., Breuer, K. S., and Tan, C. S., 2005, “Active Control of Tip Clear-ance Flow in Axial Compressors,” ASME J. Turbomach., 127, pp. 352–362.

�12� Zheng, X. Q., Zhou, X. B., and Zhou, S., 2005, “Investigation on a Type ofFlow Control to Weaken Unsteady Separated Flows by Unsteady Excitation inAxial Flow Compressors,” ASME J. Turbomach., 127, pp. 489–496.

�13� Shen, M. Y., Niu, X. L., and Zhang, Z. B., 2001, “The Three-Point Fifth-OrderAccurate Generalized Compact Scheme and Its Applications,” Acta Mech.Sin., 17�2�, pp. 142–150.

�14� Lele, S. K., 1992, “Compact Finite Difference Schemes With Spectral-LikeResolution,” J. Comput. Phys., 103, pp. 16–42.

�15� Baldwin, B. S., and Lomax, H., 1978, “Thin Layer Approximation and Alge-braic Model for Separated Turbulent Flows,” AIAA Paper No. 78-257.

�16� Zheng, X. Q., 2006, “Investigation on the Conditions of Flow Pattern Trans-
form in Unsteady Vortex Flow-Field of Axial Compressors,” Ph.D. thesis.
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Separation Control of Axial Compressor Cascade by Fluidic-Based Excitations

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