# Scale effects on the performance of sawtooth spoilers in transonic rectangular cavity flow

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## Abstract

An experimental study was conducted on the effectiveness of sawtooth spoilers at suppressing acoustic tones within a rectangular cavity with a length-to-depth ratio of five and a width-to-depth ratio of two, operating at a freestream Mach number of 0.71. Whereas previous research has focussed on the ratio of spoiler height to boundary-layer thickness (*h*/*δ*), this study also considers the effect of the ratio of cavity length to boundary-layer thickness (*L*/*δ*). Using a combination of unsteady pressure measurements and particle image velocimetry, it was established that consideration of the magnitude of both parameters is important when designing passive control methods for transonic cavities. A correlation was developed which suggests that in order to suppress fully the cavity tones, a critical spoiler height, *h* _{cr}, is defined such that *h* _{cr}/*δ* = 0.065 ≤ *L*/*δ* ≤ 0.082.

## Keywords

Particle Image Velocimetry Shear Layer Particle Image Velocimetry Measurement Broadband Noise Particle Image Velocimetry Data## List of symbols

*D*Cavity depth (m)

*f*Frequency (Hz)

*h*Height of spoiler (m)

*K*Ratio of disturbance velocity in the shear layer to the freestream velocity

*L*Cavity length (m)

*M*Mach number

*n*Acoustic mode number

- \(u_i\)
Velocity vector components in Cartesian coordinates \((u_1,u_2,u_3) \equiv (u,v,w)\)

*W*Cavity width (m)

*x*Streamwise distance from the leading edge of the cavity (m)

*y*Distance from the cavity exit plane normal to the floor (m)

*z*Spanwise distance from the cavity centre line (m)

*δ*Boundary-layer thickness (to \(0.99u_{\infty }\)) (m)

*γ*Ratio of specific heats

- OASPL
Overall sound pressure level (dB)

- SPL
Sound pressure level (dB)

- SWS
Sawtooth spoiler

- \(()_{\infty }\)
Freestream value

- \(()_{\text {cr}}\)
Critical value

- \(()_{\text {mean}}\)
Mean value

- \(()_{\text {rms}}\)
Root-mean-square value

## 1 Introduction

Cavities are increasingly appearing in aerospace applications, such as weapons bays on high-speed aircraft. When the bay doors are open, the resulting cavity flow can produce high acoustic and aerodynamic loads that may lead to structural failure. The control of cavity flow oscillations is, therefore, of practical significance and has, consequently, been the subject of numerous studies (Shaw et al. 1988; Perng and Dolling 2001; Ukeiley et al. 2004; MacManus and Doran 2008; Lawson and Barakos 2009; Vikramaditya and Kurian 2009a, b; Lada and Kontis 2011; Martinez et al. 2012). The flow control method can be passive or active depending on the external energy input (Cattafesta et al. 2008). Active control, such as blowing (Lusk et al. 2012), has the advantage that closed-loop feedback can be included to adapt to different frequencies. On the other hand, passive control methods involve simple geometric modifications and do not involve any direct energy input (although energy may be extracted from the flow or may be needed to deploy a passive device).

*f*

_{ n }is the frequency corresponding to the

*n*th mode,

*K*is the ratio of disturbance velocity in the shear layer to the freestream velocity (generally taken as 0.57), and

*α*is an empirical constant employed to account for the phase lag between the passage of a vortical disturbance past the cavity trailing edge and the formation of an upstream travelling disturbance. The value of

*α*depends on

*L*/

*D*and is evaluated as:

*α*= 0.062 (

*L*/

*D*). The ratio of specific heats,

*γ*, is assumed to be equal to 1.4 for air.

Previous research has shown that the geometrical scale of the cavity has a decisive influence on the noise levels and flow type of a cavity (Thangamani et al. 2014). The relevance of the scale of a model can be seen in two different aspects. First, at a given Mach number, cavities can be exposed to different thicknesses of the incoming boundary layer depending on factors like the position of the weapons bay on the aircraft and the altitude at which the aircraft is flying. Secondly, it determines the factors to be considered when a control technique that has been tested in a wind tunnel only at sub-scale is implemented at full-scale. Scaling effects have been recognised as contributory factors when anomalies are identified in cavity flow data. Shaw (1989) and Ross (2001) in their experiments with different cavity scales and a fixed boundary-layer thickness attributed the cause of discrepancies in their observations to the variation in boundary-layer thickness with respect to the cavity length.

The aim of this article is to identify the capability of a sawtooth spoiler to suppress cavity acoustic tones under different scale conditions. The two scaling parameters considered are the ratio of the spoiler height to boundary-layer thickness, *h*/*δ*, and the cavity length to boundary-layer thickness, *L*/*δ*.

## 2 Experimental set-up

This section details the wind tunnel facility and test rig, including a description of the cavity models and the sawtooth spoilers. The experiments were conducted in a transonic wind tunnel with the cavity models attached to the side wall of the test section.

### 2.1 Transonic wind tunnel

^{3}reservoir that supplies the air to the wind tunnel. The tunnel has a run time of approximately 8 s for a computer-controlled blowing pressure of approximately 500 kPa (gauge) at a tunnel stagnation pressure of 102 kPa. Tests were conducted at freestream Mach numbers of 0.71 and 0.85 with mean tunnel working section static pressures of 73.1 and 63.5 kPa, respectively. The unit Reynolds numbers were calculated to be 12.3 × 10

^{6}and 13.1 × 10

^{6}m

^{−1}, respectively. The uncertainty in the centre-line Mach number was obtained from five tunnel runs of 65,536 samples each and calculated at 95 % confidence level as ±0.01.

*δ*, was estimated by measuring the stagnation pressure profile at a distance 45 mm upstream of the leading edge of the cavity model and was defined as the height from the wall to where \(u / u_{\infty } = 0.99\). The boundary-layer thickness was determined from the measured profile to be 8.0 mm (see Fig. 2).

### 2.2 Cavity rig

*L*×

*W*×

*D*= 80 mm × 40 mm × 16 mm (cavity A); 160 mm × 80 mm × 32 mm (cavity B); and 320 mm × 160 mm × 64 mm (cavity C). The

*L*/

*D*and

*L*/

*W*ratios were, therefore, five and two, respectively, for all three cavities. The upstream edge of each cavity was maintained in the same streamwise position in the wind tunnel by inserting one of the three different cavity models (A, B or C) into the test rig (Fig. 3).

*L*/

*D*for all tests, a 2-mm pocket was machined into the front wall of each cavity to which the spoilers were fixed. Table 1 shows the different sawtooth spoiler geometries and the different cavities on which they have been tested. The reference axis system used for the experiments is shown in Fig. 4; the origin is at the centre of the leading edge of the cavity. The

*x*-axis is parallel to the flow direction; the

*y*-axis is normal to the cavity exit plane.

Sawtooth spoiler geometries tested: pitch is defined as the distance between adjacent ‘teeth’

Total width (mm) | Pitch (mm) | | | Cavity |
---|---|---|---|---|

40 | 5 | 2 | 0.25 | A |

40 | 10 | 8 | 1 | A |

40 | 10 | 16 | 2 | A |

80 | 10 | 4 | 0.5 | B |

80 | 10 | 8 | 1 | B |

80 | 10 | 16 | 2 | B |

160 | 10 | 8 | 1 | C |

160 | 20 | 16 | 2 | C |

160 | 20 | 24 | 3 | C |

160 | 20 | 32 | 4 | C |

*x*/

*L*= 0.1. Three such rows were placed at planes

*z*/

*W*= 0,

*z*/

*W*= 0.1625 and

*z*/

*W*= 0.325, which are referred to as “Centreplane” (CP), “Offsetplane1” (OP1) and “Offsetplane2” (OP2), respectively. In total, there were 27 static pressure tappings for the cavity model, forming a measurement grid on the floor as shown in Fig. 5. Flow visualisation of a cavity with the same

*L*/

*D*and

*L*/

*W*ratios by Atvars et al. (2009) indicated that the mean cavity flowfield is reasonably symmetric about the centreline. Hence, symmetry is assumed, and the pressure tappings have been made only for half of the floor plane.

### 2.3 Instrumentation

#### 2.3.1 Pressure measurements

The unsteady pressure measurements in the experimental tests were made with an integrated pressure measurement device, the Scanivalve ZOC22B, and data were collected from the pressure tappings made on the cavity floor. The analogue signal from the ZOC22B was acquired by a modular amplifier system (DEWE-RACK) which has an analogue-to-digital converter. The signal was amplified, conditioned and digitised before being sent to a computer. A National Instruments DAQ card 6036E was used as the interface between the computer and the DEWE-RACK. For the spectral analysis of the unsteady pressure samples, a total of 65,536 samples, averaged over four tunnel runs, were recorded at a sampling rate of 12.5 kHz. The power spectra were plotted using an FFT algorithm to determine the amplitude of the different frequency components. The maximum error in amplitude of the tones was determined to be 0.208 dB.

#### 2.3.2 Velocity measurements

Water was used for seeding the flow. Filtered clean water was pressurised using a three-stage pump and sprayed through a rake of *Bete* atomising nozzles positioned across the span of the settling chamber of the wind tunnel. The rake consisted of three atomisers located at equal intervals at the mid-height position of the test section. The nominal seeding particle diameter based on a seeding pressure of 2000 psi is quoted by Ritchie (2005) as 10 μm.

All the PIV measurements given in this work were obtained from the mid-plane of the cavity (*z*/*W* = 0). The double-pulse frequency of the laser was set at 15 Hz with a pulse separation of 6 μs. The images were acquired using a 1016 × 1008 pixel resolution Kodak ES 1 Megaplus camera fitted with a Nikkor 60 mm f2.8 lens. A total of 700 image pairs acquired over 10 tunnel runs were used to create a single mean velocity plot. A Dantec Flow Map 500 acquisition controller was used as the hardware box, which worked in conjunction with Flow Manager (v3.21) software. Both the camera and laser were connected to the hardware box, and the timing and acquisition functions of both were controlled using the Flow Manager software. Due to the large size of the cavity model and limited resolution of the camera, the PIV data for cavity C were created from two separate sets of acquisitions. Image pairs for cavity segments *x*/*L* = 0 to *x*/*L* = 0.5 and *x*/*L* = 0.5 to *x*/*L* = 1 were acquired separately and combined together to get the final data for the whole length of the mid-plane. Post-processing of all PIV data was done using DaVis software from LaVision. A three-pass window deformation algorithm, with initial and final interrogation sizes of 32 × 32 pixels and 12 × 12 pixels and a 50 % overlap between subsequent interrogation areas, was used to create the vectors. The time-averaged data were obtained from the ensemble average of 700 image pairs.

The uncertainties in PIV measurements stem from a number of factors which include errors occurring due to the particle flow dynamics, bias errors, random errors and statistical error in the mean and rms velocities, which are strongly related to the sample size. To remove outliers, the median filter proposed by Westerweel (1994) was used to locate spurious vectors and replace them by interpolation. On average, over 95 % of the vectors in a given field were found to be valid prior to replacement and interpolation. The accuracy with which the position of the correlation peak is determined can be affected by bias errors. A source of bias error in PIV measurements is the velocity gradients present inside the interrogation area. This is because particles with higher velocities tend to leave the interrogation area between the pulses, biasing the result towards lower velocities. This has been minimised by selecting displacement limits between the pulses and by using window deformation algorithms as described above to increase the number of matched particles. Progressively decreasing the interrogation window size reduces the loss of particles from the interrogation area, leading to improved accuracy and better signal-to-noise ratios (Keane and Adrian 1992). Bias error can also occur due to the phenomenon of peak locking. Peak locking occurs because of the bias of the estimation of the position of the correlation peak towards an integer number of pixels. Peak locking is attributable mainly to the choice of the sub-pixel estimator used to locate the correlation peak. In the current experiments, a Gaussian sub-pixel estimator was used which was shown by Westerweel (1997) to be superior in terms of mitigating peak-locking effects. Random errors in PIV can arise from background noise, shot noise (which can vary from pixel to pixel) and random errors associated with properly identifying the sub-pixel displacement. Although these errors can play an important role in the determination of instantaneous PIV results, their effect on ensemble averages of large numbers of samples (as is the current case) can be considered negligible owing to the random nature of their occurrence. Therefore, random errors do not degrade the accuracy of statistics derived from larger PIV ensembles (Christensen 2004). Overall, the maximum error in the mean and RMS velocities was determined to be 2.8 and 2.6 %, respectively.

## 3 Results and discussion

*h*/

*δ*= 1. Spoilers with a height approximately equal to the boundary-layer thickness have been tested by a number of researchers including Rossiter (1966), Shaw et al. (1988), Ashworth (2008) and Lawson and Barakos (2009). Figure 7 shows the power spectra of cavities A, B and C with and without the use of sawtooth spoilers. The first three predicted Rossiter tones, R1, R2 and R3, calculated from Eq. 1, are also shown in the plots. It can clearly be observed that the effectiveness of the spoilers is not the same for all three cavities.

With the use of spoilers, there is a significant decrease in both the broadband and tonal amplitudes for cavities A and B. Cavity A has a maximum reduction of 5 dB in its dominant tone (R1), and tones R2 and R3 are reduced to broadband levels. For cavity B, there is a maximum reduction of 12.7 dB for its dominant tone (R2) which is no longer dominant. R1 is now the dominant tone, and R3 has been shifted to a higher frequency at around 2 kHz. The spoiler has very little effect on the amplitude of the tones for cavity C. Although there is a reduction of 7.4 dB in the dominant tone, the power spectrum is not reduced to broadband. There are high-amplitude peaks present even after the use of the spoiler, although with reduced magnitude when compared to the clean cavity. This suggests that the same spoiler is unable to cope with the increase in flow unsteadiness caused by the increase in cavity scale (*L*/*δ*).

*u*supports the belief that the spoilers deflect the shear layer away from the cavity. Due to the presence of the spoiler, the shear layer is vertically displaced or ‘lifted’, which decreases its interaction with the aft wall. This leads to a decreased momentum transfer to the acoustic source region, which is considered the main reason for tone amplification. It is worth noting, however, that lifting of the shear layer alone would not be responsible for the decrease in momentum transfer. Spoilers introduce significant three-dimensionality into the shear layer and decrease its spatial coherence. They also thicken the shear layer due to vorticity induced in the streamwise direction. These factors also contribute to a reduced momentum transfer from the freestream to the vicinity of the aft wall.

*u*

_{mean}for cavities A, B and C were observed to be 18.1, 36.1 and 41 % of the freestream velocity for the respective clean cavities, these maxima have reduced to 4.9, 15.2 and 28 % for cavities with a spoiler. This indicates 73.2, 57.7 and 31.3 % reduction in the values of peak streamwise velocity with a spoiler for cavities A, B and C, respectively. The velocity reduction is not constant, therefore, for all three cavities. The fluctuating vertical velocity,

*v*

_{rms}, also showed a non-constant magnitude relative to the freestream velocity, with larger-scale cavities showing higher vertical velocity components (Fig. 11). Due to the use of the spoilers with the same

*h*/

*δ*, the length up to which the shear layer is vertically displaced should be equal. It is suggested that this is the reason for a stronger interaction between the shear layer and the cavity trailing edge as

*L*/

*δ*is increased.

*u*). The increase in negative velocities is due to the lifting of the shear layer and a decrease in pressure inside the cavity (Thangamani et al. 2014). The decrease in the positive streamwise velocity and increase in the negative streamwise velocity are not consistent across the different cavity scales as clearly shown in Fig. 12, which shows the percentage change in the values of maximum +

*u*and maximum −

*u*with the use of spoilers. Cavity A is most affected by the use of spoilers and cavity C the least affected. Where cavity tones are not fully suppressed, the change in negative streamwise velocities with the use of spoilers is likely to be responsible for the change in frequency of the tones observed in the power spectra due to a change in the velocity of the acoustic feedback wave.

Although spoilers with *h*/*δ* = 1 were able to attenuate totally the cavity tones and reduce the frequency spectrum to broadband level for small scales (cavity A), it was also found that the spoilers were incapable of suppressing cavity tones and unsteadiness at the larger scales (cavities B and C). This suggests that leading-edge spoilers cannot be scaled to the boundary-layer thickness alone to achieve suppression. When a change in *L*/*δ* results in a less-effective spoiler, the only practical solution is to increase its height, *h*. To deduce the correct spoiler height required for a particular value of *L*/*δ* requires the testing of different spoiler heights at different values of *L*/*δ*.

*h*, decreases the amplitude of the cavity tones and the broadband noise. It can also be seen from Fig. 13a that increasing the spoiler height from

*h*/

*δ*= 0.25 to 1 for cavity A decreases the tonal amplitudes and broadband noise significantly. However, increasing the spoiler height from

*h*/

*δ*= 1 to

*h*/

*δ*= 2 does not have much effect. This shows that the reduction in cavity noise reaches a saturation level at a particular spoiler height. This happens when all the cavity tones get reduced to a minimum broadband level achievable. The same behaviour can also be noticed for cavity C when

*h*/

*δ*is increased from three to four. Once the tones are attenuated to a broadband level, any further increase in spoiler height has no effect on noise reduction. The value of spoiler height at which this saturation is attained will be called the critical spoiler height,

*h*

_{cr}. The value of OASPL reached after attaining the

*h*

_{cr}is not the same, however, for the different scales. This is because the OASPL obtained after attenuation of all tones is due to broadband noise. This broadband noise is created mainly due to turbulent fluctuations in the shear layer. With changes in scale or

*L*/

*δ*, the turbulent fluctuations could change. At larger cavity scales, the broadband noise is higher due to the growth and convection of the turbulent fluctuations in the streamwise direction.

*h*/

*δ*are shown in Fig. 14. It can be seen that the OASPL curves are dependent on the value of

*L*/

*δ*. It was discussed above that a spoiler with height equal to the boundary-layer thickness was insufficient to attenuate the tones for cavity C. However, when the spoiler height was raised to three times the boundary-layer thickness, the tones were completely attenuated to the broadband level (see Fig. 13c). This shows that the spoiler heights have to be adjusted with a change in

*L*/

*δ*. In the literature, the size required for a leading-edge control (

*h*

_{cr}here) has been assumed to depend only on factors such as Mach number,

*L*/

*D*and

*L*/

*W*. Here, however, the results indicate that the critical spoiler height,

*h*

_{cr}, required for a particular cavity is dependent on the factor

*L*/

*δ*, i.e. on the boundary-layer thickness relative to the cavity size. Hence, for a given cavity, \(h_{\text {cr}} = f(M_{\infty }, L/D, L/\delta \)). The value of

*h*

_{cr}is defined in this work as the lowest value of

*h*that achieves a condition where all the tones are completely eliminated and the OASPL is a minimum, or within ±0.25 dB of the minimum value. The value of

*h*

_{cr}/

*δ*for cavities A, B and C obtained from the current experiments can be observed as 1, 2 and 3, respectively, from Fig. 14. It can also be seen from Fig. 14 that increasing the value of

*h*/

*δ*beyond

*h*

_{cr}/

*δ*slightly increases the OASPL. This appears to be because a further increase in spoiler height increases the turbulent broadband noise generated by the spoiler.

*h*

_{cr}. For the clean cavities (Fig. 15a), it can be seen that the OASPL generally has higher values near the trailing edge. This increase is due to the amplification of the shear-layer disturbances in the downstream direction. For a fixed boundary-layer thickness, the OASPL values increase as the cavity size is increased; again this is due to the greater opportunity for downstream amplification. For the case with spoilers (Fig. 15b), the OASPL for each cavity is reduced to an approximately constant broadband level. The streamwise amplification of the disturbances is effectively suppressed.

*h*

_{cr}/

*δ*obtained are plotted against the corresponding values of

*L*/

*δ*including the point (0, 0) in Fig. 16. Uncertainty bars (based on 50 % of \(\Delta h\)) are given for the different values of

*h*

_{cr}/

*δ*because not all values of

*h*/

*δ*have been tested. Since

*h*

_{cr}/

*δ*is chosen as the lowest value of

*h*/

*δ*when all tones are eliminated, only a lower limit is taken. Figure 16 shows the data points obtained with their uncertainties. Using the points available, a linear approximation has been made. A linear regression curve, weighted according to the error at each point, is constructed. Due to the uncertainties, a confidence band (at 95 % confidence level) is also shown.

*y*= 0.082× and

*y*= 0.065×. From this linear approximation, it is clear that the value of

*h*

_{cr}is dependant only on the cavity length

*L*such that

*h*

_{cr}/

*L*is independent of the boundary-layer thickness, suppression is not achieved for a given cavity until this ratio is achieved. For \(h / L < h_{\text {cr}} / L\), cavity oscillations are still vulnerable to changes in the ratio

*L*/

*δ*. Once \(h / L = h_{\text {cr}} / L\) oscillations are suppressed and

*L*/

*δ*no longer plays a role in cavity unsteadiness.

Previous researchers such as Shaw et al. (1988), Geraldes (2005) and Lawson and Barakos (2009) have presented data on sawtooth spoilers similar to the one used in this work; however, it must be noted that although they observed cavity tone attenuations, Shaw et al. (1988) and Lawson and Barakos (2009) were unable to suppress successfully all the tones to broadband level [see, for example, Figure 11 in Shaw et al. (1988)].

*h*/

*δ*used in the works by Shaw et al. (1988) and Lawson and Barakos (2009) is also shown in Fig. 16. Although Shaw et al. (1988) do not specifically quote boundary-layer thickness, there are sufficient data within the report to make an estimate; this has been calculated to be 5.2 mm, based upon a flat plate calculation. Referring to an earlier report on the F-111 weapons bay (Clark 1979), the full-scale boundary-layer thickness at the leading edge of the weapons bay was measured as 88.9 mm (4.4 mm at 4.9 % scale) and between 8.0 and 8.4 mm at 6.67 % scale (5.9–6.2 mm at 4.9 % scale), giving confidence in the estimated value. It is noted that the value of

*h*/

*δ*used by these researchers falls below the value of

*h*

_{cr}/

*δ*recommended by the linear model. By contrast, the value of

*h*/

*δ*used by Geraldes (2005) is close to the linear approximation, and a good level of suppression was achieved in his studies (Fig. 17). According to the linear approximation model presented here, a cavity with \(L/D = 5\) achieves suppression when the height of the sawtooth spoiler used is 6.5–8.2 % of the cavity length. This is irrespective of the value of \(\hbox {L}/\delta \) faced by the cavity and indicates that sawtooth spoilers with critical height have to be scaled according to the cavity dimensions and not according to the incoming boundary-layer thickness as previously thought. Once the critical spoiler height,

*h*

_{cr}, of a particular cavity is determined, then the spoiler can be scaled according to the cavity dimensions. It is presumed that the critical spoiler height is mainly dependent on factors like spoiler profile and

*L*/

*D*ratio of the cavity. For the current case, a spoiler with height equalling 6.5–8.2 % of the cavity length achieves critical height. A more accurate estimate of

*h*

_{cr}/

*L*can be estimated by testing a large number of the intermediate values in the

*h*/

*δ*—OASPL curves. Shaw et al. (1988) and Lawson and Barakos (2009) quote spoiler heights equalling 4.7 and 3.76 % of the cavity lengths, respectively, which are less than the critical spoiler heights required.

## 4 Conclusions

The influence of scaling on the effectiveness of a sawtooth passive control device for cavity tone suppression was studied. By using spoilers of the same *h*/*δ* ratio on cavities of different scales, it was found that there was significant variation in the control effectiveness. Spoilers with *h*/*δ* = 1 were able to suppress the tones effectively on smaller cavities, but they failed at larger scales. These results show that scaling the spoiler with respect to the boundary-layer thickness is not correct. It also shows that the factor *L*/*δ* has significant impact on the effectiveness of the passive control.

Different spoiler heights were tested for a given cavity (or *L*/*δ*), and it was found that the spoilers were effective in eliminating the cavity tones when the ratio of spoiler height to onset boundary-layer thickness, *h*/*δ*, reached a particular value. After this value, any further increase in spoiler height was found to have negligible effect. The spoiler height at which this phenomenon was achieved was termed the critical spoiler height, *h* _{cr}. By correlating the ratios, *h* _{cr}/*δ* and *L*/*δ*, it was found that they are approximately linearly related to each other. The linear model implied that the ratio *h* _{cr}/*L* is constant for a cavity of given *L*/*D* and is immune to the effects of *L*/*δ*. This implies that as long as *h* _{cr}/*L* is achieved, the cavity is expected to have no tones and exhibit the lowest OASPL possible for that particular *L*/*δ*. Once *h* _{cr}/*L* for a particular cavity specification (\(L/D,\, L/W,\, M_{\infty }\) and spoiler profile) is known, then the passive control method can be scaled according to the cavity dimensions. If the spoiler height is less than *h* _{cr}, then the cavity is not immune to the changes in *L*/*δ*. Finally *h* _{cr}/*L* is suggested to be a more suitable criterion for scaling a passive control device because it depends only on the physical cavity dimensions and not the flow conditions.

## References

- Ashworth R (2008) DES of a cavity with spoiler. In: Peng SH, Haase W (eds) Advances in hybrid RANS-LES modelling, notes on numerical fluid mechanics, vol 97. Springer, Berlin, pp 162–171CrossRefGoogle Scholar
- Atvars K, Knowles K, Ritchie SA, Lawson NJ (2009) Experimental and computational investigation of an ‘open’ transonic cavity flow. Proc Inst Mech Eng G J Aerosp 223(4):357–368CrossRefGoogle Scholar
- Cattafesta LN, Song Q, Williams DR, Rowley CW, Alvi FS (2008) Active control of flow-induced cavity oscillations. Prog Aerosp Sci 44(7–8):479–502CrossRefGoogle Scholar
- Christensen K (2004) The influence of peak-locking errors on turbulence statistics computed from PIV ensembles. Exp Fluids 36:484–497CrossRefGoogle Scholar
- Clark R (1979) Evaluation of F-111 weapon bay aero-acoustic and weapon separation improvement techniques. Technical Report AFFDL-TR-79-3003, Air Force Flight Dynamics Laboratory, Wright Patterson Airforce BaseGoogle Scholar
- Geraldes P (2005) Instabilities in transonic cavity flows. PhD thesis, Cranfield UniversityGoogle Scholar
- Heller H, Holmes D, Covert E (1971) Flow-induced pressure oscillations in shallow cavities. J Sound Vib 18:545–553CrossRefGoogle Scholar
- Keane RD, Adrian RJ (1992) Theory of cross-correlation analysis of PIV images. Appl Sci Res 49:191–215CrossRefGoogle Scholar
- Lada C, Kontis K (2011) Experimental studies of open cavity configurations at transonic speeds with flow control. J Aircr 48(2):719–724CrossRefGoogle Scholar
- Lawson SJ, Barakos GN (2009) Assessment of passive flow control for transonic cavity flow using detached-eddy simulation. J Aircr 46(3):1009–1029CrossRefGoogle Scholar
- Lusk T, Cattafesta L, Ukeiley L (2012) Leading edge slot blowing on an open cavity in supersonic flow. Exp Fluids 53(1):187–199CrossRefGoogle Scholar
- MacManus DG, Doran DS (2008) Passive control of transonic cavity flow. J Fluid Eng 130(6):064501–064501-4CrossRefGoogle Scholar
- Martinez MA, Cicca GMD, Iovieno M, Onorato M (2012) Control of cavity flow oscillations by high frequency forcing. J Fluid Eng 134(5):051201–051201-11CrossRefGoogle Scholar
- Perng SW, Dolling DS (2001) Suppression of pressure oscillations in high-Mach-number, turbulent, cavity flow. J Aircr 38(2):248–256CrossRefGoogle Scholar
- Ritchie SA (2005) Non-intrusive measurements and computations of transonic cavity flows with applications to aircraft stores release. PhD thesis, Cranfield UniversityGoogle Scholar
- Ross J (2001) High speed acoustic measurements in cavities. Technical Report F49620-98-1-0167, DERAGoogle Scholar
- Rossiter JE (1966) Wind-tunnel experiments on the flow over rectangular cavities at subsonic and transonic speeds. Reports and Memoranda 3438, Aeronautical Research CouncilGoogle Scholar
- Shaw L (1989) Scale effect on the flow induced acoustic environment in cavities. Technical Report WRDC-TM-89-159-FIBG, Wright Research and Development Center, Wright Patterson Airforce BaseGoogle Scholar
- Shaw L, Clark R, Talmadge D (1988) F-111 generic weapons bay acoustic environment. J Aircr 25(2):147–153CrossRefGoogle Scholar
- Thangamani V, Knowles K, Saddington AJ (2014) The effects of scaling on high subsonic cavity flow oscillations and control. J Aircr 51(2):424–433CrossRefGoogle Scholar
- Ukeiley LS, Ponton MK, Seiner JM, Jansen B (2004) Suppression of pressure loads in cavity flows. AIAA J 42(1):70–79CrossRefGoogle Scholar
- Vikramaditya NS, Kurian J (2009a) Effect of aft wall slope on cavity pressure oscillations in supersonic flows. Aeronaut J 113(1143):291–300CrossRefGoogle Scholar
- Vikramaditya NS, Kurian J (2009b) Pressure oscillations from cavities with ramp. J Aircr 47(12):2974–2984Google Scholar
- Westerweel J (1994) Efficient detection of spurious vectors in particle image velocimetry data. Exp Fluids 16:236–247Google Scholar
- Westerweel J (1997) Fundamentals of digital particle image velocimetry. Meas Sci Technol 8:1379–1392CrossRefGoogle Scholar

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