51st AIAA Aerospace Sciences Meeting including the New Horizons Forum and Aerospace Exposition
07 - 10 January 2013, Grapevine (Dallas/Ft. Worth Region), Texas
AIAA 2013-0103
Suppression of Vortex Shedding from a Circular Cylinder by
using a Suction Flow Control Method
Wenli Chen1
Department of Aerospace Engineering, Iowa State University, Ames, Iowa, 50011, USA;
School of Civil Engineering, Harbin Institute of Technology, Harbin, Heilongjiang, China, 150090, China
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Hui Hu2 ()
Department of Aerospace Engineering, Iowa State University, Ames, Iowa, 50011, USA
Hui Li3
School of Civil Engineering, Harbin Institute of Technology, Harbin, Heilongjiang, 150090, China
An experimental study was conduct to suppress the vortex shedding from a circular
cylinder by using a suction flow control method. The suction flow control was accomplished
using two suction holes located on the test cylinder model at an angle of 90.0 degrees in
relation to the oncoming flow direction. In addition to measuring the pressure distributions
around the test model, a high-resolution digital particle image Velocimetry (PIV) system was
used to conduct detailed flow field measurements to quantify the transient shedding
behavior of the unsteady vortex and turbulent flow structures in the wake of the test
cylinder model under different suction flow control conditions. The effectiveness of the
vortex shedding suppression using the suction flow control method was evaluated based on
the quantitative measurement results, which include the mean and fluctuating pressure
coefficient distributions around the test cylinder model, the lift and drag coefficients of the
test cylinder model with and without suction flow control, the instantaneous velocity and
vorticity fields, the mean velocity and turbulence kinetic energy distributions around the
cylinder model, and the mean velocity profiles in the wake of the test model under different
suction flow control conditions. The measurement results of the present study revealed that
the suction flow control method exhibited excellent performance in suppressing the
alternative vortex shedding from the circular cylinder model (i.e., change the alternative
vortex shedding mode into symmetric mode). As a result, the fluctuations of the lift
coefficients and drag coefficients of the cylinder model were found to be reduced
dramatically by using the suction flow control method.
Nomenclature
Cpi
= wind pressure coefficient of the point i on the surface of the cylinder model, i = 1 − 62
Cx
Cy
D
Fx
Fy
L
=
=
=
=
=
=
pi
drag coefficient of the test model
lift coefficient of the test model
diameter of the cylinder model
aerodynamic drag force of the test model
aerodynamic lift force of the test model
length of the test model
= wall static pressure of the point i on the surface on the cylinder model
1
Visiting scholar, Department of Aerospace Engineering; Assistant professor, School of Civil Engineering.
Associate Professor, Department of Aerospace Engineering, AIAA Associate Fellow, Email: huhui@iastate.edu
3
Professor, School of Civil Engineering.
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American Institute of Aeronautics and Astronautics
2
Copyright © 2013 by Wenli Chen, Hui Hu and Hui Li. Published by the American Institute of Aeronautics and Astronautics, Inc., with permission.
p∞
= static pressure in front of the cylinder model
Re
= Reynolds number
U0
= velocity at the inlet
x
y
ρ
= coordinate along the in-flow direction
= coordinate along the cross-flow direction
= air density, assumed to be 1.205kg/m3
θi
=
∆θi
= angle difference between two neighbouring pressure taps, ∆θi = 11.25° .
pressure tap angle relative to the
x coordinate
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I. Introduction
I
n recent years, long-span suspension bridges and cable-stayed bridges have been widely constructed worldwide
due to their superior structural performance and elegant appearance. As key components of cable-stayed bridges,
inclined cables often vibrate under wind, rain and traffic loads, such as the vortex-induced vibration (VIV) (Main
and Jones, 1999; Matsumoto et al., 2003; Zuo et al., 2008; Zuo and Jones, 2010) and rain-wind induced vibration
(RWIV) (Hikami and Shiraishi, 1988). Although the VIV of the cable is a self-limiting vibration, it frequently
occurs at lower wind speeds. Zuo et al. (2008) observed the VIVs of cables on the Fred Hartman Bridge in Houston,
Texas, USA through a long full-scale measurement investigation. Frequent wind-induced vibration may induce
fatigue damages to the cables, and mitigation is necessary to reduce the vibration.
Viscous damper devices have been widely used to control the VIVs and RWIVs of the stay cables of cablestayed bridges (Main and Jones, 2001; Persoon and Noorlander, 1999; Zhou and Xu, 2007). Main and Jones (2001)
obtained an additional optimal damping ratio to the stay cable provided by the viscous damper. Wang and Xu (2007)
employed an active stiffness control method to suppress the RWIVs of prototype stay cables. Li et al. (2007) and
Liu et al. (2007) investigated the theory, algorithm and model test of the semi-active control of stay cables’ windinduced vibrations using magnetorheological fluid dampers, which can effectively reduce the cable vibrations and
increase the stay cables’ system damping ratios.
The control mechanisms of each type of damper device provided additional damping ratios to the stay cables and
increased their ability to resist wind-induced vibrations, but the methods did not change or reduce the wind loads on
the stay cables. Aerodynamic methods, including passive and active approaches, can be used to reduce the wind
loads on the stay cables. They can simultaneously increase the natural damping ratios of the stay cables at certain
times. Matsumoto et al. (1992) and Flamand (1995) proposed the use of protuberances to suppress the RWIVs of
stay cables. Gu and Du (2005) conducted wind tunnel tests using double-spiral wires to mitigate vibration. Owen et
al. (2001) attached hemispherical bumps to the circular cylinder surface to control the vortex-induced vibration
amplitudes of the cable models.
Many active aerodynamic approaches were also used to control the flow separation and flow-induced vibration.
Wu et al. (2007) developed a moving-wall, e.g., appropriate travelling transverse wave, control strategy to manage
the unsteady separated flow over a circular cylinder. This method allows the global flow to remain attached when
there is a strong adverse pressure gradient and eliminates vortex shedding. Modi (1997), Munshi et al. (1997) and
Patnaik et al. (2002) investigated the use of momentum injection to control the flow field around the airfoils, flat
plates, rectangular prisms, D-section prisms and circular cylinders; the authors found that this method could
effectively resist the wind-induced vortex and galloping instabilities. Grager et al. (2011) used a dynamic burst
control plate to suppress stall on an airfoil by preventing the low-Reynolds-number leading edge separation bubble
from bursting.
Suction flow controls were also widely used to suppress the flow separation of airfoils (Qin et al., 1998;
Greenblatt et al., 2006; Gbadebo et al., 2008; Chng et al., 2009) and circular cylinders (Seal and Smith 1999;
Fransson et al., 2004; Patil and Ng, 2010) by tests or CFD numerical simulations. In this paper, a steady suction
flow control method is adopted to mitigate the vortex shedding of a circular cylinder. Further details and
comprehensive study on this method to control the vortex shedding of a circular cylinder will be submitted as a
journal paper in the near future.
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II. Experimental Setup and PIV System
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The experimental study was conducted in a closed-circuit low-speed wind tunnel located in the Aerospace
Engineering Department of Iowa State University. The tunnel has a test section with a 1.0 × 1.0 ft (30.48 × 30.48
cm) cross section and the walls of the test section are optically transparent. The test model used in the present study
was a circular cylinder. The circular cylinder had a diameter d=0.0508 m and a width equal to the test section width
of the wind tunnel. Its roots were mounted on two sides of the wind tunnel wall. During the experiments, the wind
tunnel speed was adjusted to 8.0 m/s, which corresponds to a Reynolds number Re = 2.8 × 104 .
Figure 1 shows the schematic of the experimental setup used in the present study for the two-dimensional PIV
measurements. The flow was seeded with 1~5 µm oil droplets. Illumination was provided by a double-pulsed
Nd:YAG laser (New Wave Gemini PIV 120-15) adjusted on the second harmonic and emitting two pulses of 120 mJ
at the wavelength of 532 nm with a repetition rate of 8 Hz. The laser beam was shaped to a sheet by a set of mirrors,
spherical and cylindrical lenses. The thickness of the laser sheet in the measurement region is about 1.0 mm. The
high-resolution 12-bit (1376×1040 pixels) CCD camera (SensiCam, CookeCorp) was used to perform twodimensional PIV image recording. The CCD camera and double-pulsed Nd:YAG lasers were connected to a
workstation (host computer) via a Digital Delay Generator (Berkeley Nucleonics, Model 565), which controlled the
timing of the laser illumination and image acquisition.
Figure 1. Schematic of the experimental setup
Figure 2 shows the assemblage of the test model used in present study. The test model was employed in present
test as shown in figure 2(a). For the test model, four different suction sections were employed with the spaces of
L/8@L/4@L/8. The suction holes have the same diameter of 3 mm. Two suction holes were symmetrically
distributed about the x axis at each suction section, i.e., the angle of suction holes: 90.0 (270.0) degree, as shown in
figure 2(b) which illustrates the suction hole distribution. Omega FMA-2613A was used in the testing to control the
suction flow rates of the main suction tube which was separated into 8 branches: 8 suction tubes which were
connected with 8 suction holes. The flow rate was corrected to a set of standard conditions which are 25ºC and one
standard atmosphere pressure of 101.325 kPa. Five suction flow rates of main suction tubes: 20, 40, 60, 80 and 95
L/min were adopted and the corresponding mean velocities of each suction hole were 5.89, 11.79, 17.68, 23.58,
28.00 m/s, respectively.
To measure the pressure distribution on the cylinder model, a digital pressure measurement system based on the
model DSA3217 pressure scanner (Scanivalve Corporation) was used in the wind tunnel tests. As shown in figure
2(b) and (c), total 62 pressure taps were arranged on the model in two sections, 32 for pressure section 1 and only 30
for pressure section 2 because two pressure taps are replaced by two suction holes at each suction angle. Pressure
section 1 was located on the middle of the test model and pressure section 2 was shared with one suction section as
shown in figure 2(b). The sampling rate was 300 Hz and the sampling time was 100 s for the pressure measurement
of each tap. The blockage ratios for all section models were 16.67 % and the measured pressures on the cylinder had
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been corrected combining the solid blockage and the wake blockage (Barlow et al. 1999). All pressure taps are
connected by polyvinyl chloride (PVC) tubes of 0.5 m long to the pressure measurement system. The effect of the
phase distortion and magnitude variation alone on measured peak pressures was found to be small for short PVC
tubes of length less than 2 ft. (0.61 m) (Irwin et al. 1979).
360
°
Wind speed direction
Suction holes Suction holes Suction holes Suction holes
1 and 2
3 and 4
5 and 6
7 and 8
270
180
Pressure Pressure
section 1 section 2
90
0
Pressure section 1
°
Suction hole
°
Pressure tap
°
360
°
270
°
180
90
°
0
(a) Test models
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Pressure section 2
°
°
°
(b) Pressure taps and suction holes
y
Wind
θi
U0
Cy
1
D∆θ i
2
∆θ i
Cylinder section
x
Cx
D
(c) Pressure section 1
Figure 2. Setup of the test model
III. Experimental results and discussions
Figure 3 shows the distributions of the mean pressure coefficients of the test model at two suction pressure
sections, respectively. It can be seen that the separation point should be at least in front of 90.0 degrees and the
suction flow control induces great change to the distributions of the mean pressure coefficients, especially in the
flow separation region. As the flow rate increases, the absolutes of the mean pressure coefficients in the flow
separation region decrease. It is indicated that the drag forces of the test model decrease under the suction flow
control, even at a small flow rate of 20 L/min. But, the control effect will not be improved and tends to be steady
when the suction flow rate exceeds 40 L/min. For pressure section 1, the position with the minimum mean pressure
coefficient has no change under suction flow control with different flow rates. But, for pressure section 2, the point
with the minimum mean pressure coefficient moves to next point when the flow rate is larger than 40 L/min. It can
be deduced that the separation point may move backward or the flow separation is delayed at this suction section. It
should be further investigated combined with the results of PIV measurement.
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Figure 3. Mean wind pressure coefficient distribution of the test model
Figure 4 shows the distributions of the fluctuating pressure coefficients of the test model at two suction pressure
sections, respectively. Similar with the analysis of the mean pressure coefficients, the small flow rate of 20 L/min
can obtain better control effect, but the control effect tends to be steady when the suction flow rate exceeds 40
L/min. The fluctuating pressure coefficients of most taps decrease to 1/3 compared with the case without the suction
flow control.
Figure 4. Fluctuating wind pressure coefficient distribution of the test model
The coordinate system of the model and the directions of the aerodynamic coefficients are schematically
illustrated in Fig. 2(c). The aerodynamic coefficients can be obtained by integrating the wind pressures over all the
taps; the lift and drag coefficients, C y and Cx , can be calculated with the following expression:
1
1
ρU 02 ∑ C pi ⋅ D∆θ i ⋅ cos θ i
1
2
2
i
Cx =
=
= ∑ C pi ⋅ ∆θ i ⋅ cos θ i
1
1
2 i
2
2
ρU 0 D
ρU 0 D
2
2
1
1
ρU 02 ∑ C pi ⋅ D∆θ i ⋅ sin θ i
Fy
1
2
2
i
Cy =
=
= ∑ C pi ⋅ ∆θ i ⋅ sin θ i
1
1
2 i
ρU 02 D
ρU 02 D
2
2
p − p∞
C pi = i
1
ρU 02
2
Fx
(1)
Figures 5 and 6 show the time histories of lift and drag coefficients of the cylinder model, respectively. The
fluctuations of lift coefficient histories are dramatically suppressed. If the cylinder model can freely vibrate, the
oscillation amplitude will also be suppressed under a suction flow control. So the suction flow control should have
the ability to suppress the flow-induced vibration according to the analysis of the lift coefficients. Figure 6 shows
that not only the amplitudes of drag coefficients are greatly suppressed, but also the means of drag coefficients are
obviously decreased. It is shown that the suction flow control has the function of the flow drag reduction.
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Figure 5. Time histories of lift coefficients of the test model (left: section 1 and right: section2); the gray
lines denote the cases without suction control and the red lines denote those with suction control
Figure 6. Time histories of drag coefficients of the test model (left: section 1 and right: section2); the gray
lines denote the cases without suction control and the red lines denote those with suction control
For further statistical analysis, the control effects of the root mean squares (RMSs) of lift coefficients and the
means of the drag coefficients are obtained as shown in figure 7. The control effect of the lift coefficients of pressure
section 2 is just slightly better than that of pressure section 1; and the lift coefficients at two sections can reach about
80% the amplitude reduction. But, for the drag coefficient reduction, the best control effect is about 30% on pressure
section 2, but only 25% on pressure section 1. It is indicated that the control effect of the drag coefficients along the
axial direction is different; the reason may be that the flow should have a three-dimensional feature along the axial
direction. It will be further studied in present paper by analyzing flow structures of different planes along the axial
direction through PIV measurements.
(a) RMS of lift coefficients
(b) Mean of drag coefficients
Figure 7. Comparison of control effects for aerodynamic coefficients at two pressure sections
The PIV measurements were simultaneously performed when the wind pressures on the cylinder were measured
for the test model. Two measured PIV planes are located on pressure section 1 and 2. Figure 8 shows the
instantaneous vorticity field on section 2 for the test model. Under the suction flow control, the alternating vortex
shedding was changed into a symmetric mode. Figure 9 shows the mean velocity fields on section 2 for the test
model. The results indicate the velocities in the wake have the same direction with the main flow. The maximums of
T. K. E. under the flow rates of 40 and 95/min are smaller than that under the flow rate of 20 L/min.
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(a) 0 L/min
(b) 20 L/min
(c) 40 L/min
(d) 95 L/min
Figure 8. Instantaneous vorticity field of the test model for section 2
(a) 0 L/min
(a) 20 L/min
(b) 40 L/min
(c) 95 L/min
Figure 9. Mean velocity field of the test model for section 2
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Figure 10. Velocity profiles in the wake of the test model
For further investigating the suction flow control for the wake velocities, the velocity profiles along the y axis
behind 1.5 D from the center of the cylinder models have been extracted as shown in figure 10. When the cylinder
model is without the suction flow control, some velocities are negative values which indicate these velocity
directions are opposite to the main flow direction. Under a small flow rate of 20 L/min, the negative velocities are
changed into the positive ones; but as the flow rates continuously increase, the control effects have no distinct
improvement.
(a) Section 1
(b) Section 11
(c) Section 12
(d) Section 2
Figure 11. Mean velocity field of the test model under the flow rate of 95 L/min
To investigate the effects of the suction flow control for the wake along the axial direction, two measured PIV
planes are not enough, so another two planes are added between section 1 and section 2, named section11 and
section12 along the direction of section 1 to section 2. The PIV measurements are performed on these four planes at
the flow rates of 0, 20, 40 and 95 L/min for the test model.
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Figure 11 shows the mean velocity fields on four sections for the test model. The results indicate that the values
of the T. K. E. are greatly reduced and the turbulence intensities of the wake are decreased under the suction flow
control on which section. But only for the section 2, the velocities in the wake have the same direction with the main
flow and the velocities close to the centreline in the near wake of other three sections have the opposite direction
with the main flow. Furthermore, the separation point of section 2 has been moved backward compared with those
of other three sections. It can explain that the point with the minimum mean pressure coefficient moves to next point
as shown in figure 3(a). So, combining the results of the distribution of the mean pressure coefficient, it can be
deduced that the separation point may also move backward along the flow direction, i.e. the flow separation is
delayed on the section 2.
The velocity profiles behind 1.5 D from the center of four sections of the test model have been extracted as
shown in figure 12. Same with the analysis of figure 11, only for section 2, there are no negative velocities in the
wake. Under the suction flow control, the absolutes of negative velocities increase with increasing flow rate for
other three sections; besides the peak of negative velocities is less than that without suction flow control for section
12. Comparing the velocity profiles, the control effects are gradually improved along the direction of section 1 to
section 2.
(a) Section 1
(b) Section 11
(c) Section 12
(d) Section 2
Figure 12. Velocity profiles in the wake of the test model
IV. Conclusions
In this study, the suction flow control for vortex shedding of a circular cylinder based on a multi-point suction
type was investigated combining a PIV measurement system and a pressure measurement through a wind tunnel test.
The following conclusions were obtained from this study:
1. The angle of suction hole of 90.0 degrees which is close to the flow separation point can achieve good control
effects: reducing the average drag and the lift fluctuation of the test model, changing the alternating vortex shedding
into a symmetric mode and delaying the flow separation on the suction section.
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2. As a multi-point suction type, the control effects of the average drag and the lift fluctuation reduction are
improved from the section without suction to the suction section. The flow separation delaying is only achieved on
the suction section.
Acknowledgments
This research was funded by the National Natural Sciences Foundation of China (NSFC) (51008093 and
51161120359). The authors also want to thank Mr. Bill Richard and Mr. Kai Zhang of Iowa State University for his
help in manufacturing the test models.
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