Pulsed Microjet Control of Supersonic Impinging Jets:

A Reduced-Order Model

J. Choi∗ A. M. Annaswamy†

Massachusetts Institute of Technology, Cambridge, MA, 02139

F. S. Alvi‡

Florida A & M University and Florida State University, Tallahassee, FL

In recent years it has been demonstrated that direct microjet injection into the shearlayer of the main jet disrupts the feedback loop inherent in high speed impinging jet flows,thereby significantly reducing the adverse effects. The amount of noise reduced by microjetactuation is known to be dependent on nozzle operating conditions. In this paper, activecontrol strategy using pulsed microjets is suggested to maintain a uniform reduction of thesetones over the entire range of operating conditions. It was observed that the pulsed microjetwas able to completely eliminate the impinging tones at all operating conditions as the dutycycle increased. Moreover, a low frequency pulsing (16.4 Hz) brought about additionalreduction compared to high frequency pulsing, which is primarily due to the presence ofa low frequency mode in the flow-field. A two-mode feedback model that captures boththe low and high-frequency Rossiter mode as well as the interactions between the twomodes is presented. The effect of pulsing is modeled using a input-shaping controller thataccomplishes noise-reduction through a suitable redistribution of the acoustic excitationover the high and low frequencies.

Nomenclature

d Nozzle diameter of mainjet, mD Diameter of lift plate, mh Distance from lift plate to ground plane, m

I. Introduction

The Short Takeoff and Vertical Landing (STOVL) aircraft experiences discrete and high amplitude acous-tic tones that are produced via a feedback process. The feedback occurs from the strong interaction betweenthe ground and the high speed jet emanating from a STOVL aircraft nozzle. Instability waves are gener-ated by the acoustic excitation of the shear layer near the nozzle exit, which then convect down and evolveinto spatially coherent structures. Upon impinging on the ground, these structures generate acoustic waves,which in turn excite the shear layer at the nozzle exit, thereby closing the feedback loop. (see Fig.1 andref1,2). The high amplitude impingement tones are undesirable not only due to the associated high ambientnoise, but also due to the accompanied unsteady pressure loads on the ground plane and the nearby surfaces.While the high noise levels can lead to structural fatigue of the aircraft surfaces in the vicinity of the nozzles,the dynamic loads on the impingement surface can lead to an increased erosion of the landing surface as wellas a dramatic lift loss during hover.

Suppression of these tones requires destabilizing the feedback loop,1 which can be realized by suitablydisturbing the shear layer near the nozzle exit by several passive3–5 and active control methods.1,6, 7 Of these,

∗Graduate Research Assistant†Senior Research Scientist, Member AIAA‡Associate Professor, Senior Member AIAA

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12th AIAA/CEAS Aeroacoustics Conference (27th AIAA Aeroacoustics Conference)8 - 10 May 2006, Cambridge, Massachusetts

AIAA 2006-2601

Copyright © 2006 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.

the technique in ref1 appears most promising from the point of view of efficiency, flexibility, and robustness.The method in ref1 introduces microjets along the periphery of the nozzle exit which modify the shear layerat its most receptive location thereby efficiently attenuating the impingement tones.

Nozzle

Large-scale instability waves

Acoustic waves

Wall

Figure 1. Schematic diagram of an impinging jet and possible feedback path

While it has been shown in the previous work1 that an open-loop control strategy that employs themicrojets is effective in suppressing the impingement tones, the amount of suppression is dependent to alarge extent on the operating conditions. Since in practice, the flow conditions are expected to changedrastically, a more attractive control strategy is one that employs feedback and has the ability to control theimpingement tones over a large range of desired operating conditions.

Recently, we introduced a pulsed microjet as a new actuator to suppress impinging tones.8 The rationalefor doing this is that for a given mass flow rate, pulsed injection can generate larger momentum thansteady continuous microjet injection, which is consequently expected to have a stronger impact on the noisereduction mechanism.

Pulsing of jet flows has been attempted in reference.9–13 Wiltse and Glezer introduced an open-loopcontrol strategy in ref13 via high frequency forcing in the inertial subrange of a free shear layer on a lowspeed flow. They found that broadband velocity fluctuations were reduced at low frequency but increasedat high frequencies. Stanek et al.11,12 and Sinha et al.10 adopted the high frequency forcing technology forcontrol of the cavity flows and Raman et al.14 reported results applied for control of impinging tones. Morerecently, Kastner and Samimy9 reported reducing a resonant peak using HTFA (Hartmann Tube FluidicAcutator), a very high speed actuator for controlling the impinging jet noise. This actuator primarilyworked in a blowing-mode, required fairly large mass-flow rates, and worked over a fairly narrow range offrequencies whose selection required considerable tuning. Here, we pursue a low speed pulsing strategy whichis far below the natural frequency (∼4.6 kHz) of the system. The actuator used modulates the flow at theexit of the microjet using a rotating cap. Saw-tooth structures placed in the inner race of the rotating capblock and unblock the microjet holes as the cap rotates and simulates an on-off micorjet action. A similarpulsing actuator design was used to control a free jet in reference.15 However, as demonstrated in sectionII-C, the design proposed here in significantly more efficient due to the location of the actuator.

Noise reduction using pulsed microjet injection is dependent on several control parameters such as dutycycle, mass flow rate, and phase difference between adjacent microjets and pulsing frequency. Among theseparameters, the duty cycle and pulsing frequency play a major role in suppressing impinging tone. A changein the duty cycle from 43% to 75 % was shown in ref16 to result in an additional reduction of 6 db. In thispaper, we consider the effects of pulsing at low frequencies. We will show in particular that pulsing at lowfrequencies leads to an additional reduction of about 2 dB.

A two-mode lumped parameter model is proposed in this paper that captures both the dominant im-pinging tone and the low-frequency mode discussed above. In addition, the effect of pulsing is modeledthrough an input-shaping controller which suitably redistributes the acoustic excitation over the high andlow frequencies.

The paper begins with an explanation of the experimental setup in section II, followed by the control

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of impinging jet using pulsed microjets. Parameters which affect the amount of noise reduction are to bediscussed in this section further. An experimentally tuned model will be suggested in section III and themechanism of low frequency mode will be identified from the flow visualization technique, PIV.

II. Experimental Setup

A. Test Configuration and Facility

The following experiments were carried out at the supersonic STOVL jet facility of the Fluid MechanicsResearch Laboratory located at the Florida State University. A schematic of experimental setup with asingle impinging jet is shown in Fig.2. This facility is used primarily to study jet induced phenomenon onSTOVL aircraft hovering in and out of the ground effect.1,2 A circular plate of diameter D (25.4 cm ∼ 10d)was flush mounted with the nozzle exit and, henceforth referred to as the ‘lift plate’, represents a aircraftplanform. A 1 m × 1 m × 25 mm aluminum plate is mounted under the nozzle, which serves as the groundplane simulating the hovering situation by fixed it to the desired position. Further facility details can befound in ref.2

h (variable)

Lift Plate

Ground Plate

Nozzled

Dz

Microphone

Figure 2. Test geometry

The supersonic impinging jet was produced by an axisymmetric, convergent-divergent (C-D) nozzle witha design Mach number of 1.5. The throat and exit diameters (d, de) of the nozzle are 2.54 cm and 2.75cm (see Fig.2 & 3). The divergent part of the nozzle is a straight-walled conic section with a 3◦ divergenceangle from the throat to the nozzle exit. A Validyne pressure transducer measures the stagnation pressurein the settling chamber just upstream of the nozzle. Although tests were conducted over a range of NozzlePressure Ratios (NPR,where NPR = stagnation pressure/ambient pressure), the results discussed in thepresent paper are limited to NPR = 3.7 that corresponds to an ideally expanded Mach 1.5 jet. The nozzletotal pressure was maintained within ± 0.2 psi of the desired conditions.

Sixteen microjets fabricated using 400 µm diameter stainless tubes were used as actuators for active flowcontrol. These are flush mounted circumferentially around the main jet as shown in Fig.3(a). While theorientation of the jets can be varied between 0 and 90◦, most of the experiments reported in this papercorrespond to the microjets at either 20◦ or 30◦ with respect to the nozzle axis. The supply for the microjetswas provided by compressed nitrogen cylinders through a main and four secondary plenum chambers. In thismanner, the supply pressures to each bank of microjets could be independently controlled. The microjetswere operated over a range of NPR = 5 to 7 where the combined mass flow rate from all the microjets wasless than 0.5% of the primary jet mass flux.

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Microjets (dm =400µm)

de =27.5 mm

Lift plate

Kulite 1

2

3

45

6

(a)

Secondary plenumchambers

To microjets

Control valves

Primary Plenum

(b)

Figure 3. (a) Lift plate/microjet layout, (b)Microjet feed assembly.

B. Pressure Measurements

Near-field noise was measured using B&KTM microphones placed approximately 25 cm away from the jet.The microphone signal was measured with an estimated uncertainty of ± 1 dB. The distribution of unsteadyloads on the lift plate was measured by six high frequency response miniature KuliteTM pressure transducers(model: XCS-062), placed axisymmetrically around the nozzle periphery plate, at r/d =1.3 from the nozzlecenterline (Fig.3). The transducer output were measured using National Instruments digital data acquisitioncards (PC-MIO-16E-1 card coupled with SC 2040 Sample and Hold card) and LabV iewTM software at asampling rate of 70 kHz. Standard statistical analysis techniques were used to obtain the spectral contentand the Overall Sound Pressure Level (OASPL) from these measurements.

C. Pulsing Using a Rotating Cap

To obtain a more consistent noise reduction over a larger range of jet operating conditions, we examineda different control strategy, which consists of a technique that pulses the microjet flow. The rationale forintroducing pulsing is that, for a given mass flow rate, a pulsed injection can generate more momentum thansteady continuous microjet injection, and hence can perhaps have a stronger impact on the jet shear layer,thus disrupting the feedback mechanism more effectively and hence reducing the noise more significantly.16

Flow modulation was introduced using direct modulation at the exit of the microjet using a rotatingcap (see Fig.4,5). This cap consists of several teeth which block and unblock the microjet holes as the caprotates, simulating an on-off microjet action.

The effect of the pulsed microjets through the rotating cap was quantified by spinning the motor atdifferent speeds and measuring the unsteady total pressure at the microjet exit using a Kulite mounted in atotal pressure probe configuration. It was observed that the rotating cap produces a fairly large amplitudeperturbations up to 300 Hz.

In addition to providing a direct method of pulsing the microjet flow, it is also of interest to be able tovary different parameters of the pulsed flow such as amplitude, frequency, duty-cycle, and phase. This canbe accomplished by varying the design parameters of the rotating cap. The pulsing amplitude is directlyproportional to the supply pressure delivered to microjet chamber, while pulsing frequency is solely controlledby the rotation speed of the cap. Therefore, these two parameters can be easily and electronically varied bychanging the microjet pressure and the motor speed. The duty cycle(dc) is changed by varying the numberand diameter of holes of the rotating cap, a phase difference between two adjacent microjet pulses is variedby changing the number of holes in the cap to be different from that of microjets.

It should be pointed out that the swirl caused by the cap rotation itself does not significantly affect the

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Rotating Cap

Figure 4. Conceptual diagram for rotating cap actuator.

Pulley

Kulite

Motor

Rotating Cap

Small Lift Plate

Microjets

Bearing

Main jet

Tooth

Figure 5. Rotating cap design.

baseline performance. This is demonstrated in Fig.6, where the OASPL of the uncontrolled impinging jet iscompared to that while the cap was rotating without any microjet action. As can be seen in the figure, thetwo OASPLs are almost identical.

D. Results Using Pulsed Microjets

Using the above experimental setup, studies of the impinging jets were conducted with and without pulsing.The unsteady pressure measurements could not obtained on the lift plate due to the vibrations from thespinning cap incorporated in the small lift plate (see Figs.4 and 5). Instead the noise level was measured by amicrophone located at 25 cm away from the nozzle axis. The results obtained when the rotating cap was spunat a frequency of f=121 Hz and φphase = 0, and dc = 42%, 74%, respectively, are shown in Fig.7(a) and (b).These results show that the impinging tones are completely eliminated by the pulsed microjets. They alsoshow that (a) both pulsing and steady microjet action yield about the same amount of pressure reduction,and since the supply pressures were the same, it implies that the pulsing action allows noise reduction tooccur at 42% of the mass flow rate needed for the steady case. (b) They show that a significantly largerreduction can be obtained from the pulsing action under certain duty cycles, which follows from Fig.7(b).

The rotating cap was spun over a range of frequencies from 60 to 150 Hz, corresponding results of which

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are shown in Fig.8. For dc = 42%, the amount of noise reduction is quite independent of fpulsing over thisrange. In the next section, we will discuss about the effect of low frequency pulsing below 60 Hz in detail.

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0123456789

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Freq (Hz)

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Figure 8. Noise reduction for different pulsing frequencies, dc = 42% (x axis represents pulsing frequency, yaxis denotes noise reduction in dB scale)

III. Effect of Low frequency pulsing and A Reduced-order Feedback Model ofthe Impinging Jet

In this section, we discuss the effect of low frequency pulsing and propose a reduced-order feedback modelof the impinging jet that captures these effects.

A. Identification of A Low Frequency Mode

As noted in the previous section II-D, the amount of noise reduction was independent of the input pulsingfrequency in the range of 60 ∼ 150 Hz, as shown in Fig.8. However, in subsequent experiments, an additionalnoise reduction was possible by pulsing at low frequencies around 20 Hz. In repeated trials (seen in Fig.9),an additional 1 ∼ 2 dB reduction was always achieved by low frequency pulsing injection. This indicatesthe impinging jet flow has a global mode in this low frequency region. This low frequency peak can also beobserved in the spectral plot of uncontrolled impinging jet noise, shown in Fig.10.

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Test 1 Test 2

∆dB

Figure 9. Noise reduction for different pulsing frequencies, dc = 56%, h/d = 3.5 (x axis represents pulsingfrequency, y axis denotes noise reduction in dB scale)

In almost all investigations of the impinging jets, the spectral plot of the uncontrolled impinging jetnoise is distinguished by a dominant peak ∼ 4.6 kHz which corresponds to the impinging tone (shownin Fig.10.(b)). In the higher frequency range (> 4.6 kHz), the harmonics of this impinging tone appear

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0 50 100 150 200 250 300 350102

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Figure 10. (a) Spectral plot (low frequency regaion) (b) Spectral plot (High frequency region) of uncontrolledflow, h/d = 3.5, NPR = 3.7

repeatedly since the acoustic fluctuation is not a perfectly sinusoidal shape.8 Fig.10.(a) which is the spectralplot of low frequency band shows a moderate peak at ∼ 10 Hz. The plot (FFT size is 1024) was obtainedusing 20 seconds of microphone data with 2048 Hz sampling rate filtering through low pass filter with 1 kHzcutoff frequency. Hence, the corresponding resolution is 2 Hz, which is accurate enough to capture the lowfrequency mode. Seen in Fig.11, peaks in the low frequency region are also observed in other data measuredfrom the microphone, unsteady pressure fluctuation of ground and lift plate respectively. Moreover, thereis a strong correlation between the spectral plot and amount of noise reduction. This proves that the lowfrequency peak is not due to experimental errors but a meaningful mode that should be considered for modelconstruction. In the following section, we model both the dominant (Rossiter) mode at a high frequency,and the low-frequency mode. The possible mechanism of the low frequency mode will be investigated in thefollowing section in detail.

B. A reduced-order model

At the region near the nozzle exit, the evolution of shear layer instability Pshear to a large scale vorticalstructure can be modeled using a linear stability theory with hyperbolic velocity profile of main jet. But thedetermination of transfer function from the shear layer instability Pshear to impinging tone of the groundplane Pground based on the underlying physics is far too difficult to build. This is primarily due to the factthat the acoustic noise on the ground is produced by the process of the vortex annihilation while impinging onthe ground,8 which is a very nonlinear phenomenon. We therefore use an alternate, reduced-order, modelingapproach using input-output relationships. This model, shown in Fig.12, consists of a forward loop and afeedback loop. The forward loop represents the transfer function from Pshear to Pground where Pshear ispressure fluctuation of shear layer at the nozzle exit and Pground is the acoustic noise on the ground plane.Our goal is to retain the amplitudes at two most dominant frequencies, one being the largest peak at aRossiter mode, and the second being a low frequency peak, we choose the transfer function Gshear(s) to beof the form

Gshear(s) =Nmode∑

i=1

Ki

s2 + 2ζiωis + ω2i

(1)

where the ζi is the damping ratio, the ωi is the natural frequency, and Ki is the amplitude of the ith

mode, respectively. Nmode is the number of dominant modes in underlying system, which, in this case, istwo. The delay Ta = h/Cv represents the convective time-delay between the nozzle and the ground. Thefeedback loop is represented by the pure time-delay transfer function e−sTb where Tb = h/Ca is the timetaken for the reflected acoustic wave to travel from the ground to the nozzle exit, which in turn excitesthe shear-layer. The transfer function Gshear(s) represents the dominant effects of the shear-layer, and is

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C-D nozzle

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Figure 11. Spectral plot of (a) unsteady pressure on the ground plane (b) unsteady pressure on the lift plate(c) acoustic noise of microphone. (d) Acoustic noise reduction (∆dB) by pulsing actuation.

pground

Time to ground plane (Tb = h/Cv)

pshearN

Time to nozzle lip(Ta = h/Ca)

++

Gclosed(s)

Gshear(s) bsTe−

sTae−

Figure 12. Block diagram of feedback loop for uncontrolled impinging jet

described in more detail below. The noise input N represents all other input excitation that exist at otherfrequencies including broad band noise effects. This results in a closed-loop transfer function of the form

Gclosed(s) =Gshear(s)e−sTb

1−Gshear(s)e−s(Ta+Tb)(2)

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pground

Gshear(s)pshearNdist

++ bsTe−)(

)(

2

1

sZ

sZ

Gpulse(s)

Gf(s)

pint

f, φphase, dc

asTe−Pmic

RotatingCap

Figure 14. Block diagram of feedback loop for controlled impinging jet by low speed pulsing

Using these two dominant modes whose natural frequencies are 10 Hz and 4.6 kHz, the model of impingingtone system was built and compared to the experimental data seen in Fig.13(a). The blue lined indicatedabove is the microphone data measurements and the red line represent acoustic data produced by a model.The part that is not modeled, which is in the range of between 50 and 500 Hz is due to the broadband noise.Because it is a purely aeroacoustic phenomena, we can exclude its effects in consideration of the impingingtone mechanism, and is represented through the external input N . It should be noted that a feedback modelsimilar to that suggested in Fig.12 was introduced in Rowley et al.17 in the context of cavity tones.

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C. Effect of Pulsed Microjet-Control

The most immediate effect of the pulsed microjets is the alteration of the pressure Pshear in the immediatevicinity of the nozzle. The impact of these microjets is the reduction of the peak amplitude at 4.6 kHz, asshown in Fig.13(a) and (c). Such a reduction is possible only through one of two effects, one of which isvia damping, while the other is via a reduction of the input excitation at this frequency. The former caserequires the introduction of an input at the same frequency but with a different phase, an evidence of whichwas not available in the impinging jet. We therefore model the effects of the pulsed microjet-control via analternation of the input excitation, and is represented via a transfer function Gf (s) in Fig.14. This transferfunction is chosen as

Gf (s) =Z1(s)Z2(s)

(3)

where the polynomials Z1 and Z2 are such that they result in a redistribution of the input energy.As shown in Fig.13 (d) and (f), the uncontrolled system produces an input excitation which has a largeamplitude at the impinging frequency of 4.6 kHz. Relatively speaking, with a pulsed microjet, the input,which is now modified due to the presence of the pulsed microjet action, has a much smaller amplitude atthe same frequency. It should also be noted that the pulsed microjet increases the input amplitude at thelower frequency. However, due to the fact that the overall flow field is such that the lower frequency has amuch higher damping, the pressure response at this frequency is not increased despite the increase at theinput. This could be the reason for the effective noise reduction due to the pulsed microjet.

For comparison, the effect of a steady microjet is also modeled using the same transfer functions, theresults of which are illustrated in Fig.13 (b) and (e). A comparison of Fig.13 (d), (e), and (f) shows twofacts; first, the same input-shaping effect as in the pulsed microjet is exhibited in the steady microjet in thatthe input-excitation at the higher frequency is lowered and is increased at the lower frequency. However, thisshaping is not as effective as in the pulsed case; compared to the pulsed case, the excitation at the higherfrequency is decreased to a smaller extent. This in turn could be the reason why the corresponding noisereduction in the steady microjet case is smaller.

IV. Low Frequency Mode: A Possible Mechanism

The high frequency mode corresponding to 4.6 kHz is produced by impingement of a large scale vorticalstructure on the ground. It is a well known mode as predicted by many researchers such as Rossiter,18

Neuwerth19 and Powell20 in the context of feedback mechanism. However, the presence of low frequencymode (10 Hz) has not been reported hitherto in any investigations. The question then is raised as to whatthe possible mechanisms are that are responsible for such a low frequency behavior.

In the cavity test by Kegerise et al.,21 a low frequency mode was reported to be present without anydistinct physical noise source. For example, it is argued that two Rossiter modes fa,fb may produce anotherpeak which corresponds to the sum (fa + fb) or difference (fa − fb) of these frequencies when the nonlinearinteraction between Rossiter modes are prominent. Such a nonlinear interaction may not be the mechanism ofthe low frequency mode in this problem because the Rossiter modes of the impinging jet under considerationare at a fairly high frequency (fa= 4.6 kHz, fb = 6.1 kHz) and hence both the sum and difference of thetwo are far greater than the low frequency mode of 10 Hz. Moreover, the spectral plots of experimental datashown in Fig.13(a)∼(c) indicates that the peaks in the higher frequency region are quite harmonic, whichalso implies that the nonlinear interaction between Rossiter modes may be fairly weak.

The flapping (/or helical) motion of the impinging jet column is another possible mechanism of lowfrequency phenomena. Flapping mode of a plane jet (/ helical mode of axisymmetric jet) has a relativelyslow motion compared to the feedback loop, but the exact frequency of this mode is unknown. Previousresearch8 indicates that the uncontrolled flow structure is dominated by helical mode at h/d = 3.5, Ma =1.5, which supports the helical motion as a possible noise mechanism. However, in the repeated experiments,the helical structure is found to be switched to the axisymmetric mode once microjets are activated. If thehelical motion is the primary source of low frequency phenomenon, the frequency content of the controlledflow should not have the low frequency peak. In fact, low frequency mode in spectral plot in Fig.13(a)∼(c)remains unchanged even in the presence of pulsed injection controls the impinging jet. Hence, we canconclude that flapping motion cannot be the possible mechanism of low frequency mode.

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As mentioned at the beginning, the low frequency mode exists in all measurements including acousticdata, unsteady pressure signal of ground and lift. Due to this fact, we can assume that the noise source ofthe low frequency mode is perhaps located on the ground, and generates acoustic waves that travel towardthe lift plate. Hence, a closer investigation of the flow field near the impinging region may be called for.

Using a flow visualization with Particle Image Velocimetry (PIV) technique, the normalized vorticityfield was determined and is shown in Fig.15. At the center of impinging zone, we can clearly see a pairof vorticity field which is counter directional to that of main jet. This local vorticity field is produced bythe stagnation bubble which is formed by the fountain flow from the center of impinging region. In ref.,22

this stagnation bubble structure is prominent for an underexpanded impinging jet, and very weak for anoverexpanded jet. Because the nozzle was planned to be operating under ideally expanded condition, we cannot expect the strong formation of stagnation bubble. However, the nozzle experiences a slightly off designcondition during the above test. Seen in Fig.16, we can clearly see a conical weak shock cell structure formednear the nozzle exit. This indicates that the impinging jet is operating at a slightly underexpanded conditionwith a very low period. As a result, we can expect a periodic formation of stagnation bubble at the centerof the ground plane. This stagnation bubble can be the possible noise source because the properties of thebubble satisfies some characteristics in acoustic measurements: (i) the pulsation of this bubble is relativelyweak compared to the impingement of vortical structure and hence, the corresponding low frequency peakis much smaller than the high frequency peak, and (ii) the bubble still stays at the center while microjetactuation is applied and hence the low frequency mode doesn’t change while control is on. We thereforespeculate that the periodic formation of the stagnation bubble is a possible mechanism of low frequencymode. More careful investigations need to be carried out to confirm this speculation

No Control

Steady Pulsing (16.4 Hz)

Figure 15. Normalized Vorticity

V. Conclusion

In this paper, active control of supersonic impinging jets using microjets is considered. To achieveconsistent noise reduction for all operating conditions, a pulsing microjet was introduced as a new actuator,since the unsteadiness of the microjet flow can increase the forcing strength and thus have a more significantimpact on the shear layer of the main jet. The pulsing was accomplished by way of a saw-toothed rotating

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No Control

Steady Pulsing (16.4 Hz)

Figure 16. Axial turbulent intensity

cap that was incorporated in the lift plate which, when spined the saw-tooth periodically blocks and unblocksthe microjet flow. Using this method, we were able to pulse the microjet flow up to several hundred Hz.

Several parameters of the pulsing were varied to determine conditions under which the flow structureand noise reduction were most receptive to. It was observed that variations in the duty-cycle of the pulsingled to a maximum impact. It was also found that pulsing frequency plays an important role. Low frequencypulsing injection at around 10 Hz found to be causing additional 1∼2 dB reduction, which was due to thepresence of low frequency mode in the impinging flow field. A reduced-order model was constructed basedon low (10 Hz) and high (4.6 kHz) frequencies. The effect of the rotating cap was also included in the form ofan input-shaping controller. It was shown that the pulsed microjet control accomplishes noise reduction byextracting energy from the high frequency region to the low frequencies. A possible mechanism that causeslow frequency mode is attributed to a stagnation bubble which was periodically observed in the PIV to format the center of the impinging region of the jet

Acknowledgments

This work was supported by a grant from AFOSR, monitored by Dr. J. Schmisseur. We would liketo thank the staff of FMRL, for their invaluable help in conducting these tests. We are grateful to Mr.Choutapalli for his help in conducting the tests, and Robert Avant for his advice in designing the experimentalsetup. In addition, Mr. DePriest’s prompt support made it possible to implement several crucial tests.

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