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8/9/2019 25 Degree Ahmed Body aerodynamic study
1/19
13
Experiments in Fluids
Experimental Methods and their
Applications to Fluid Flow
ISSN 0723-4864
Volume 52
Number 5
Exp Fluids (2012) 52:1169-1185
DOI 10.1007/s00348-011-1245-5
Drag reduction on the 25 slant anglehmed reference body using pulsed jets
Pierric Joseph, Xavier Amandolse &
Jean-Luc Aider
8/9/2019 25 Degree Ahmed Body aerodynamic study
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13
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8/9/2019 25 Degree Ahmed Body aerodynamic study
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RE S E A RCH A RT I CL E
Drag reduction on the 25
slant angle Ahmed referencebody using pulsed jets
Pierric Joseph Xavier Amandolese
Jean-Luc Aider
Received: 11 June 2011 / Revised: 13 November 2011 / Accepted: 29 November 2011 / Published online: 15 December 2011
Springer-Verlag 2011
Abstract This paper highlights steady and unsteady
measurements and flow control results obtained on anAhmed model with slant angle of 25 in wind tunnel. On
this high-drag configuration characterized by a large sep-
aration bubble along with energetic streamwise vortices,
time-averaged and time-dependent results without control
are first presented. The influence of rear-end periodic
forcing on the drag coefficient is then investigated using
electrically operated magnetic valves in an open-loop
control scheme. Four distinct configurations of flow control
have been tested: rectangular pulsed jets aligned with the
spanwise direction or in winglets configuration on the roof
end and rectangular jets or a large open slot at the top of the
rear slant. For each configuration, the influence of the
forcing parameters (non-dimensional frequency, injected
momentum) on the drag coefficient has been studied, along
with their impact on the static pressure on both the rear
slant and vertical base of the model. Depending on the type
and location of pulsed jets actuation, the maximum drag
reduction is obtained for increasing injected momentum or
well-defined optimal pulsation frequencies.
1 Introduction
Current environmental and economic issues lead automo-
tive manufacturers to search for innovative solutions to
reduce vehicles fuel consumption. One way is to reduce
aerodynamic drag, which is responsible for the largest
part of the fuel consumption for speed above 80 km h-1
(Hucho1998). Like bluff-body, automotive drag is mainly
governed by massive separation on the rear part: for a
typical family car, pressure drag on this area can reach near
a third of the total aerodynamic drag (Barnard 1996).
In order to simplify the study of automotive near wake,
and thus to understand aerodynamic drag generation in the
rear part of a vehicle, Ahmed et al. (1984) introduced a
simplified geometry (Fig.1a). Despite the fact that this
geometry is close to its thirtieth anniversary, it is still lar-
gely used by scientific community as an automotive ref-
erence model to work on complex three-dimensional wake
flow and its control, using numerical methods (Krajnovic
and Davidson 2005a, b; Fares 2006) and experimental
techniques (Beaudoin et al.2004; Thacker2010; Gillieron
2010).
The flow topology of the Ahmed body, and thus its
aerodynamic drag, is greatly dependent of the slant angle.
As this angle evolves from 0 to 90, the near wake of the
Ahmed bluff-body changes drastically. From 12 to 15,
the flow is typical of a rear blunt with a flow separation on
rear edges, generating mainly transverse vorticity: this first
type of separation is sometimes considered as quasi-two-
dimensional (Hucho 1998). From 15 up to 30, the near
wake is highly three-dimensional with partial separation on
the slant surface along with strong conical streamwise
vortices coming from the slant side edges and a ring-
shaped structure lying on the base surface. Beyond 30, the
separation can also be considered as quasi-two-dimensional
P. Joseph
Institut AeroTechnique (IAT), CNAM, 15 rue Marat,78210 Saint Cyr lEcole, France
e-mail: [email protected]
X. Amandolese
Aerodynamics Department, CNAM, 15 rue Marat,
78210 Saint Cyr lEcole, France
e-mail: [email protected]
J.-L. Aider (&)
PMMH Laboratory, UMR 7636, CNRS, ESPCI ParisTech,
10 rue Vauquelin, 75231 Paris, France
e-mail: [email protected]
1 3
Exp Fluids (2012) 52:11691185
DOI 10.1007/s00348-011-1245-5
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due to the massive separation from the top of the rear slant.
Figure1b is a schematic view of the mean flow topology
for the three-dimensional situation (Vino et al. 2005).
This three-dimensional complex wake exists with a 25
slant angle and presents high-drag coefficient, which makes
it a good test case for drag reduction study.
After years of drag reduction using shape optimization,
this technique shows its limit regarding design constraints
of the automotive industry. This trend causes flow control
techniques to be more and more studied, with the Ahmed
body as a benchmark.
A lot of successful studies can be found in the literatureusing passive strategies: for example, Fourrieet al. (2011)
obtained 9% of drag reduction using a classical automotive
style deflector, while Beaudoin and Aider (2008) reached
an impressive 25% reduction with several flaps located on
the edges of the rear end of a 30 configuration. An
approach using vortex generators to produce coherent
streaks that increase or decrease the separation bubble was
also carried out by Aider et al. (2009) and Pujals et al.
(2010) leading, respectively, to a 12 and 10% drag reduc-
tion. However, apart from the active vortex generators
proposed by Aider et al. (2009), all these passive tech-
niques introduce quite unsightly appendages on the body,which is in contradiction with design constraints.
This fact makes active control by jet or suction very
attractive in automotive industry. This kind of control is
nearly invisible and can be adapted to changes in flow
conditions. Various successful active control studies have
been conducted using the Ahmed reference body with 25
slant angle. Roumeas et al. (2008) used steady aspiration on
the top of the slant: he obtained numerically a drag reduc-
tion of 17% and noticed experimentally a suppression of the
separation area (Roumeas2006). Experimental studies were
carried out by Leclerc (2008) with synthetic jets (zero net
mass flux) at the top slant edge area (8.5% reduction) and by
Krentel et al. (2010) with pulsed jets at the bottom slant
edge (5.7% reduction). Both Krajnovic et al. (2009) and
Lehugeur (2009) made numerical simulations of the same
case: the former obtained a little bit more than 7% reduction
using steady blowing and suction at the slant top edge (and
also studied several other blowing locations and jet types),
while the latter used steady blowing to force the bursting of
longitudinal coherent structures, leading to a 6% drag
reduction. Brunn et al. (2008) have experimented anadvanced control strategy by targeting simultaneously
particular structures with different actuator types (steady
and periodic), including closed-loop features. Beaudoin
et al. (2008) used as well feedback control by extremum
seeking on a rounded Ahmed Body. Periodic forcing was
also successfully used by Pastoor et al. (2008) on a more
simplified body. They managed to reduce drag by 15% by
synchronizing upper and lower vortex shedding with a
synthetic jet control system in closed loop.
According to those studies, it seems that periodic forcing is
a promising way of controlling the flow structures on the 25
Ahmed body. A good understanding of this kind of control isachieved on academic geometry, like backward-facing step
(see Tihon et al. 2010). Optimal frequencies are clearly
identified among the base flow instabilities, as theeffect of the
forcing amplitude. Meanwhile, on three-dimensional com-
plex flow like theAhmed body wake, influenceof each forcing
parameters (injected momentum, non-dimensionalfrequency,
spanwise modulation, etc.) is still under discussion.
In the present paper, we will focus on the suppression
of the rear slant recirculation bubble, without acting on
Fig. 1 aSide view and front view of full-size Ahmed body with 25slant angle.b Schematic view of rear flow topology for slant angle between
12.5 and 30, from experimental study of Vino et al. (2005)
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longitudinal structures. According to Thacker et al. (2009),
expected drag reduction is then about 10%.
The experimental setup will be described in a first part.
In the second part, main time-averaged and time-dependent
results obtained on the base reference Ahmed model will
be highlighted. The control parametric study using various
configurations of electrically operated valves in an open-
loop control scheme will then be presented. The linkbetween the drag reduction and the modification of the
mean static pressure distribution over the slant and rear
vertical base of the model is clearly demonstrated.
2 Experimental setup
Experiments were carried out in the 5 m 9 3 m test sec-
tion of the S4 wind tunnel at the Institut AeroTechnique
(France), using a 1.044-m length Ahmed model mounted
over a raised floor (see Fig. 2).
Due to the large cross-section of the wind tunnel com-pared with the model size, no blockage corrections are
necessary in the present study (blockage ratio B = 0.7%).
2.1 The Ahmed reference body
The generic car model used here is the one originally
described in Ahmed et al. (1984). In the present study, we
focus on the 25 slanted rear end in order to deal with the
high-drag configuration characterized by a large separation
bubble over the slanted surface along with highly energetic
streamwise vortices created along the slant side edges
(see Fig. 1b). The dimensions and the overall shape of the
model are given in Fig. 1a. The main dimensions of the
model are L = 1.044 m in length, H = 0.288 m in height
andl = 0.389 m in width. The height of the 25 rear slant
is h = 0.094 m, and the height of the rear base is
Hv = 0.194 m.
2.2 Wind tunnel and reference incoming flow
Experiments were carried out for flow velocities ranging
from 20 to 40 m s-1. The Reynolds number,ReL = U0L/t,
based on the overall lengthLof the model, ranges between1.4 9 106 and 2.8 9 106. The turbulence level in the S4
wind tunnel is less than 1.2% over this velocity range.
In order to reduce the influence of the natural boundary
layer growing on the wind tunnel ground, the Ahmed body
is fixed over a raised floor, 0.115 m above wind tunnel
ground. Other dimensions are given in Fig. 3.
Special attention has been given to the raised floor
leading edge in order to avoid any boundary layer sepa-
ration of the incoming flow. It has been designed based on
a NACA 0018 airfoil (Fig. 3) to avoid the increase in static
pressure that could generate a massive boundary layer
separation of the overall wind tunnel boundary layer.This modification significantly decreases the amount
of perturbations coming to the model and reduces the
incoming boundary layer thickness.
2.2.1 Upstream boundary layer
The incoming boundary layer profile has been measured
0.7 m downstream from the raised floor leading edge, i.e.,
DX/L = -0.29 upstream from the Ahmed model. The
mean velocity profile is reported in Fig. 4for a wind tunnel
velocity U0 = 20 m s-1. The boundary layer thickness d
is calculated according to the 99% criterion. The shape
parameter HBL has been calculated using the classical
definition of the boundary layer displacement and
momentum thickness. The boundary layer thickness d is
about 25 mm, and the shape parameter HBL is close to
1.25. The incoming boundary layer is then fully turbulent,
Fig. 2 View of the experimental setup in the 5 m 9 3 m test section
of the S4 wind tunnel
Fig. 3 Schematic description of the experimental setup (Ahmed
model on the raised floor)
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which is confirmed by the 1/8th power law that fits well the
experimental measurements.
2.3 Experimental measurements
2.3.1 Aerodynamic balance
Time-averaged forces were measured using a six-compo-
nent strain gauge balance mounted under the raised floor.
The balance is located in a dedicated rounded compartment
to avoid stream-induced perturbations on the force mea-
surements (Fig. 5).
Calibration was made out of wind tunnel using standard
procedure, and calibration checks were also conducted in
situ. The maximum error in drag measurements associated
with repeatability and hysteresis was found to be approx-imately 0.5%.
2.3.2 Wall-pressure measurements
Steady wall-pressure measurements were carried with 121
pressure taps located inhomogeneously mainly on the
slanted surface, as well as on the roof end and vertical rear
base. Because of the body symmetry, only a half of the
model was equipped. Symmetry of the flow was previously
checked with the help of surface oil flow visualizations on
the entire slant surface. All the taps were plugged in a
Scanivalve pressure scanner. Precision of this system is
usually 0.03% of the full scale.
On the other half of the model, six piezoelectric micro-
sensors were implanted. These sensors allow both steady
and unsteady measurements in areas where characteristic
structures of the Ahmed wake are expected. These sensors
have a typical sensitivity of 1 Pa. The location of all these
pressure taps is shown in Fig. 6.
2.3.3 Surface oil flow visualizations
Surface oil flow visualizations were conducted using a
mixture of silicone oil, dodecan, titanium dioxide and oleic
acid. This mixture was applied with paintbrushes on the
slanted surface and allows visualization of friction lines
when the model is exposed to air flow. Friction lines give
information about mean flow topology (Gillieron2000).
2.3.4 Near-wake total pressure loss measurements
Wake measurements were carried out with the help of a
two-dimensional motorized explorer, allowing measure-ments in a transversal plane with hot-wire or Kiel probe.
The last one could be used for unsteady total pressure
measurements thanks to the embedded piezoelectric sensor.
2.3.5 Description of the pulsed jets control device
Pulsed jets are obtained using eight electromagnetic binary
valves feeding a rectangular chamber before flowing
through a removable perforated plate (Fig. 7).
Fig. 4 Evolution of the boundary layer on the raised floor upwind
the Ahmed model for U0 = 20 m s-1 (ReL = 1.4 9 10
6) at
DX/L = -0.29
Fig. 5 Schematic description of the experimental setup (aerodynamic
forces measurement system)
Fig. 6 Pressure taps distribution on the rear end of the model, the
open circles correspond to the 6 piezoelectric sensors
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Thanks to this setup, the jets geometry and configura-
tions can be easily changed from a long continuous slot to a
set of winglet-type jets just by changing the removable
plate. Electromagnetic valves (Matrix Ltd.,) are built as a
magnetic circuit closed by a steel spring tongue, which can
take one of two stable positions. A short, low-energy
electric impulse applied to the coil can change the spring
position to the opposite one, thereby clearing or closing theoutput opening. The valve controller is stimulated by a sine
wave generator furnishing a wavy train of variable fre-
quency (Fj) in the range of 5300 Hz. The level of the
pressure impulse, and then the jet velocityUj, can be varied
by simply changing the pressure supply level (Pj).
To ensure the spatial homogeneity of the jet speed along
the actuation slot, small calibrated balls (2 mm diameters)
are set in the chamber between the valves exits and the
perforated plate mounted on the wall of the model (Fig. 7).
The porous layer distributes the air flow from each valves
exhaust to the entire surface of the perforated plate.
A typical time history of the jet velocity Uj(t) measuredwith a hot-wire 1 mm above a jet exhaust is shown in
Fig.8. Due to the valve technology, the jet velocity is
periodic but not sinusoidal. Indeed, the signal is closer to a
square wave signal but exhibits a significant overshoot and
associated rebounds. This overshoot is characteristic of this
type of valves, which induces brutal pressure release
immediately after the valve opening.
In the present paper, the pulsed jets are characterized by
their mean velocity Uj and main frequency Fj.
One can notice that this signal exhibits some additional
fluctuations. It is not clear whether it can have an influence
on the flow control experiment. This point is complex and
still under consideration.
2.3.6 Description of the flow control configurations
Three different perforated plates have been used in order to
test the influence of different types of pulsed perturbations
(Fig. 9): discontinuous slot, continuous slot andwinglets. Each plate can be used on two different locations:
roof end 100 mm upstream the slant edge (X/L = -0.1),
and slant top edge 15 mm (X/L = 0.01) downstream the
slant upper edge (Fig. 9). Blowing sections are detailed in
blue, with dimensions in millimeters. The choice of these
configurations is justified in Sect.4.
3 Characterization of the base flow
This first step is obviously to study the natural flow aroundthe body. This knowledge will be helpful to compare to
previous studies and will help in understanding the flow
control mechanisms.
3.1 Drag coefficient
Drag force is expressed by its drag coefficient, with the
following expression:
CX FX
12qSU20
1
where FX is the drag force measured by the aerodynamicbalance,q is the air density (corrected with the atmospheric
pressure and wind tunnel ambient temperature), S is the
model cross-section (excluding struts) and U0 is the free
stream velocity.
This drag coefficient was measured for several Reynolds
numbers ReL corresponding to U0 from 20 to 40 m s-1.
Results are shown in Fig. 10.
A significant Reynolds effect is observed as the drag
coefficient decreases with increasing Reynolds number
(from CX = 0.335 at ReL = 1.4 9 106 to CX = 0.312 at
ReL = 2.7 9 106). However, these results are consistent
with other studies like Aider et al. (2009) and Roumeas(2006) in the same Reynolds number range.
3.2 Steady wall-pressure distributions on the rear end
and associated surface oil flow visualizations
As the flow is symmetric, steady wall-pressure measure-
ments were made over the half of the slant surface allowing
pressure coefficient mappings on the entire slant surface.
Pressure coefficient Cp is expressed as:
Fig. 7 Schematic description of the pulsed jets device using perfo-
rated plates system to change the jets geometry
Fig. 8 Typical time history of jet velocity Uj(t) (Fj = 200 Hz,
measured at 1 mm above a jet exhaust)
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Cp p p0
12qU20
2
where p is the local static pressure and p0 is a reference
static pressure measured upwind the model, in the undis-
turbed flow.
In order to highlight the link between flow structures and
pressure distribution, pressure coefficient mappings are
associated with surface oil flow visualizations on Fig. 11.
From left to right, Reynolds number increases from
1.4 9 106 (20 m s-
1) to 2.8 9 106 (40 m s-
1). XS is localaxis, i.e., the x-axis projection along the slant.
For eachReynolds number, one can observe a low-pressure
area on the top of the slant, followed by a gradual pressure
recovery at the bottom of this surface. Comparing with cor-
responding flow visualizations, the pressure contours match
well with the recirculation bubble (circled in red), which is
responsible for the low-pressure distribution in this area.
On both sides of the slant, other low-pressure areas are
visible. They are located under the longitudinal vortices
(Fig.1b), which also have an important role on the low-pressure repartition and thus on the drag.
A significant Reynolds effect can be observed on the
mean pressure distribution (Fig. 11). Indeed, the recircu-
lation area significantly decreases with the Reynolds
number. The upper slant low-pressure area then gets
smaller, and pressure recovery occurs sooner. Even though
the pressure coefficient becomes smaller on the top of the
slant, its reduction cannot balance the earlier pressure
recovery leading to an overall mean pressure value on the
slant surface, which increases with the Reynolds number.
This tendency is consistent with the drag reduction
observed on Fig.10in the same range of Reynolds number.The highest drag configuration that exhibits the largest
separation bubble (for ReL = 1.4 9 106) has been chosen
to carry out the flow control experiments.
3.3 Near-wake total pressure loss measurements
Base flow topology has been investigated using time-
averaged total pressure loss measurements in the near
wake. Results are presented as total pressure loss coeffi-
cient defined as:
Cpi 1
pT p012qU20 3
where pT is the total pressure measured in the wake. This
coefficient value is zero in the undisturbed flow (i.e., no
pressure loss) and gradually increases as total pressure in
the wake decreases due to pressure losses associated with
mixing processes (shear layer, recirculation areas,
vortices ).
Results are reported on Fig.12 for a cross-section
located at DX = 0.144 m behind the model (i.e., a relative
Fig. 9 Perforated plates
geometry and blowing locations
Fig. 10 Evolution of the drag coefficient with the Reynolds number
ReL (without control)
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distance DX/H = 0.5), for U0 = 20 m s-1 (ReL = 1.4 9
106). The structure of the wake is classic: the conical
streamwise vortex signature (with a core located near
Z1 = 220 mm andY1 = 85 mm (Y1,Z1) being a local axis
system associated with the plan-wake), the base ring-
shaped structure and the mixing region associated with the
flow separation over the slant surface.
3.4 Velocity profiles
Various velocity profiles have been measured in several
locations to complete the characterization of the mean base
flow.
3.4.1 Boundary layer over the model roof
Figure13 shows the boundary layer profile on the model
roof atX = -0.1 m (X/L& -0.1) upstream from the slant
edge. The boundary layer thickness isd & 24 mm, and the
shape parameter HBL & 1.21. The incoming boundary
layer is then fully turbulent, which is confirmed by the
1/8th power law that fits well the experimental measure-
ments.
3.4.2 Shear layer
The recirculation bubble is separated from the so-called
external flow region by a shear layer characterized by a free
stream velocity (close to the wind tunnel velocity) and a
Fig. 11 Influence of the Reynolds number on the pressure coefficient distribution on the rear slant and associated surface oil flow visualizations
(without control)
Fig. 12 Total pressure loss coefficient distribution in the near wake
of the model in theDX/H = 0.5 cross-section (without control) in the
(Y1, Z1) local axis system associated with the plan-wake
Fig. 13 Boundary layer mean velocity profile on the model roof for
U0 = 20 m s-1 (ReL = 1.4 9 10
6) at X/L& -0.1
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quasi-zero velocity area in the recirculation region. This is
a region of intense mixing characterized by strong velocity
gradients and turbulence intensities. Velocity profiles have
been measured in the vertical symmetry plane (Y = 0) of
the Ahmed model at two distinct axial positions from the
slant edge: X/h & 0.1 and X/h & 0.5 (with h, the height
of the slant face). Results are reported on Fig.14 for
U0 = 20 m s-1.The mean velocity profiles are well fitted by a hyper-
bolic-tangent velocity profile. The model used is expressed
as (Ho and Huerre1984):
u z U 1 R tanh z z0
2h
h i 4
where U Umax Umin=2 is the average velocity,
R DU=2Uis the velocity ratio, DU Umax Umin is thetotal shear,z0the mean vertical position of the shear layer,
i.e., the position of the inflexion point, and h is the
momentum thickness of the shear layer.
Values of those parameters for both the positionsX/h & 0.1 and X/h & 0.5 are reported in Table 1, along
with the associated Reynolds numbers based on the
momentum thickness and average velocity Reh Uh=m.
3.5 Unsteady measurements
The Ahmed model exhibits intense unsteady three-dimen-
sional wake. According to Thacker (2010), this unsteadi-
ness is mainly concentrated in the shear flow region over
the slant surface and in the near-wake region. In the shear
flow region, the unsteady flow features can be associated
with both an absolute and convective instability of the
shear layer (Cherry et al. 1984; Kiya and Sasaki 1985;
Thacker 2010). The former being associated with a
flapping of the shear layer and the latter at a natural
KelvinHelmholtz instability of the shear layer (Aider
et al. 2007). On the other hand, the unsteady character-
istic of the near-wake flow region is mainly the conse-
quence of an unstable organization due to the flow
separation on both the upper and lower edges of the rearvertical base linked to the ring-shaped structure observed
in the wake.
According to Thacker (2010), a significant level of
velocity fluctuations can also be measured in both the two
steady streamwise vortical structures, but mainly due to an
interaction with the unsteadiness of the shear layer, and
thus only at a significant distance from the core of the
vortices.
3.5.1 Unsteady organization of the shear layer
Unsteady velocity measurements have been performed inthe shear layer for both the positions X/h & 0.1 and
X/h & 0.5. Power spectral densities associated with
velocities measured at the inflexion point of both shear
layer profiles (see on Fig. 14) are reported on Fig. 15.
Results are shown in a non-dimensional form introducing
the Strouhal number Sth = fh/U0 (the reduced frequency
basedon the slant height),for two wind tunnelvelocitiesin
order to highlight specific unsteady organization that
could be characterized by a constant value of Strouhal
number.
At the position X/h & 0.1, the non-dimensional spec-
trums exhibit a significant low-frequency organization
Fig. 14 Shear layer velocity
profiles, with:a mean velocity
profiles andb root mean square
velocity profiles, at two axial
positions from the slant edge:
X/h & 0.1 andX/h & 0.5, for
U0 = 20 m s-1
(ReL = 1.4 9 106)
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characterized by a Strouhal number Sth & 0.1 for
U0 = 20 m s-1 and Sth & 0.14 for U0 = 30 m s
-1, with
frequency values of, respectively, f& 20 Hz and
f& 45 Hz. According to Kiya and Sasaki (1985), a con-
stant Strouhal number can be associated with this low-
frequency organization due to the flapping of the shear
layer, with an appropriate definition of the reduced fre-
quency based on the length of the recirculation bubble Lr.
Results are reported in Table2, where Lr has been esti-
mated from the surface oil flow visualizations (see on
Fig.11).
At the position X/h & 0.5, the non-dimensional spec-
trums exhibit the same low-frequency organization along
with a significant increase in the energy fluctuations in a
higher frequency range between Sth & 0.5 (f& 100 Hz
for U0 = 20 m s-1 and f& 150 Hz for U0 = 30 m s
-1)
and Sth & 2 (f& 450 Hz for U0 = 20 m s-1 and
f& 650 Hz for U0 = 30 m s-1). Taking the shear layer
relevant parameters, i.e., the momentum thickness h and
the average velocity U measured at X/h & 0.5 (see
Table1), those high-frequency fluctuations occur between
Sth & 0.033 andSth & 0.13. In the light of the work of Ho
and Huerre (1984), those fluctuations can then be associ-
ated with the roll-up of the shear layer due to the Kelvin
Helmholtz instability mechanism.
3.5.2 Unsteady organization of the near wake
Unsteady velocity measurements have been also carried
out on several locations in the near wake, exhibiting a
strong unsteady organization characterized by a constant
Strouhal numberStHv & 0.31 (based on the rear-end ver-
tical height Hv = 0.194 m).
Results are reported on Fig. 16 for a point near the
bottom of the rear end where the organization is particu-
larly strong (Z/H = -1). Indeed, the non-dimensional
spectrums (based on the power spectral densities of mea-
sured velocities) exhibit strong and narrow peaks for a
reduced frequency StHv & 0.31, indicating a very orga-
nized phenomenon. Taking another definition of the
reduced frequency based on the square root of the model
cross-section A =HS, one find StA = fA/U0 & 0.53,which is in accordance with the results of Vino et al. ( 2005)
and Thacker (2010).
Table 1 Parameters of the
hyperbolic-tangent velocity
profiles used atX/h & 0.1
and X/h & 0.5
X/h Umax (m s-1) Umin (m s
-1) U R Z0 (mm) h (mm) Reh
0.1 21 0 10.5 1 -2.75 1.2 700
0.5 20.5 3.5 12 0.71 -12 3.75 3,000
Fig. 15 Non-dimensionalpower spectral densities of the
velocity at the inflection point of
the shear layer at a X/h & 0.1
andb X/h & 0.5 for
U0 = 20 m s-1 and 30 m s-1
(ReL = 1.4 9 106 and
ReL = 2.1 9 106)
Table 2 Parameters of the low-frequency organization of the shear
layer forX/h & 0.1
U(m s-1) f(Hz) Sth Lr (m) StLr
20 &20 &0.1 &0.17 &0.1730 &45 &0.14 &0.13 &0.195
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4 Flow control experiments
Among the control parameters, one can identify physical
parameters associated with the jet (mean and maximum jet
velocities, pulsation frequency, duty cycle, signal form,
etc.) and geometric parameters (shape of the cross-section
of the nozzle, number and spatial organization of the jets,
location of the jets over the model, jets angles, etc.) In the
present work, control experiments were realized with four
different geometric configurations:
Discontinuous slot, corresponding to rectangular jets
aligned along the spanwise direction, over the slant
upper edge and roof end.
Continuous open slot close to the slant upper edge.
Winglets jets over the roof end.
Each of them corresponds to different flow control
strategies.
The discontinuous slot and winglets configurations over
the roof end correspond to jet vortex generators. The
objective is to test, with two geometrical configurations,
the impact of longitudinal vorticity injection upstream the
separation point. The idea is to use pairs of streamwise
counter-rotating vortices induced by the jets in cross-flow
(Cortelezzi and Karagozian 2001) to modify the property
of the boundary layer and postpone the separation of the
boundary layer (Duriez et al. 2006,2008a,b). In the case of
pulsed jets, one can expect both a modification of the mean
flow and, as a consequence, of the shear layer, together
with an effect of the pulsation frequency injected in the
shear layer.
The continuous slot at slant top edge was intended to
quantify the effect of transversal vorticity injection near the
separation point. This configuration is inspired by the work
of Leclerc (2008) who showed that it was possible to
decrease the drag with synthetic jets. In this case, the
injection is homogenous along the spanwise direction so
that the shear layer is perturbed by a time-periodic span-
wise vorticity sheet. In this case, no streamwise vorticity is
injected.
The discontinuous slot at the same location was used to
experiment the effect of the reduction of injected
momentum quantity and spanwise modulation. In this case,
the shear layer is no longer perturbed by a spanwise vortex
but rather by a set of streamwise vortices spaced along the
spanwise direction, even if their location downstream the
separation make the comparison with jets in cross-flow
more difficult.
Dimensionless quantities are used for jets speeds and
jets frequencies, with the classical definition for momen-
tum coefficient Cl and dimensionless frequency Stj (Sj is
the perforated plates blowing surface for the considered
control configurations):
Cl qSjU
2j
1=2qSU20
5
Stj Fjh
U06
For each geometric configuration, two physical para-
meters, jets speed and jets frequency, were varied. For
every parameter, drag coefficients and pressure coefficients
were measured and plotted as iso-contours in the (Cl,Stj)
space. White areas in the contour plots correspond to
parameters that have not been measured mainly because ofelectric power limitations. In the following, all data
were obtained at ReL = 1.4 9 106 (corresponding to
U0 = 20 m s-1).
4.1 Influence of the forcing parameters on drag
Drag coefficient without and with control are, respectively,
noted CX0 and CXC. Figures17, 18, 19 and 20 show the
drag coefficient variations DCX = (CX0 - CXC)/CX0 as a
function of the momentum coefficient Cl and dimension-
less frequency Stj for the four geometric configurations. In
the following, the space parameter is mapped with incre-
ments dFj = 20 Hz or dFj = 40 Hz (i.e., dStj & 0.1 or
dStj & 0.2) depending on the tested configuration. In the
same way, Cl variations were obtained by changing Pjwith dPj = 0.5 bar and dPj = 1 bar, corresponding to
dCl = 0.3 9 10-3 or dCl = 0.6 9 10-3. DCX[ 0 cor-
responds to drag reduction in percentage.
One can observe that 8% drag reduction is reached for
various configurations and that the corresponding physical
parameters depend strongly on the geometric ones.
Fig. 16 Power spectral
densities of the velocity in the
near wake forU0 = 20 m s-1
and 30 m s-1
(ReL = 1.4 9 106 and
ReL = 2.1 9 106)
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The discontinuous slot at roof end (Fig. 17) leads to drag
reductions for almost all the frequency range tested with
two optimal Strouhal number at Stj = 0.62 andStj = 1.03.
This drag reduction seems to be optimal for a relatively
narrow range of momentum coefficients 3 9 10-3\Cl\
3.5 9 10-3. For this configuration, both the amount of
injected momentum quantity and pulsation frequency play
an important role, even if the influence of the pulsation
frequency is weaker.
On the opposite, the continuous slot configuration at
slant top edge (Fig.18) leads to a significant drag reduction
for a particular frequency range 0.3\ Stj\0.6 with
local optimal areas for Stj & 0.28 and Cl & 2.2 9 10-3,
Stj & 0.35 and 0.4 9 10-3\Cl\0.9 9 10-3 and
Stj & 0.56 andCl & 0.3 9 10-3. The influence of the jet
velocity (and then of the injected momentum) seems here
to be linked to the jet pulsation frequency that plays a
major role in the drag reduction mechanism.
Two important points should be noticed. First, the
evolution of the drag reduction in the (Cl, Stj) space
parameters is completely different for the two configura-
tions. It confirms that the actuations (and control strategy)are completely different. Second, the maximum drag
reduction is the same (about 8%) but for much smaller Cl
(0.4 9 10-3 instead of 3 9 10-3) with the continuous slot
at slant edge.
It is not possible to link the optimal frequency for the
drag reduction to natural frequencies measured in the shear
layer region. Nevertheless, one can notice that some of
them are close to the KelvinHelmholtz frequency
Sth & 0.5 measured at X/h & 0.5 in the shear layer.
For the two last configurations (winglets at roof end on
Fig.19and discontinuous slot at slant top edge on Fig. 20),
the drag reduction is maximal for given points in the spaceparameter: Stj & 1.1 and Cl C 1.2 9 10
-3 for the wing-
lets at roof end configuration and 1 9 10-3\Cl
\ 2.2 9 10-3 and Stj & 0.55 or Stj C 0.9 for the discon-
tinuous slot at slant top edge configuration. In these cases,
both jet velocity and pulsation frequency seem to be
important to optimize the drag reduction. In the case of the
discontinuous slot at slant top edge, the optimal frequency
is also close to the natural frequency measured in the shear
layer.
Fig. 17 Influence of the control parameters (Stj and Cl) on the drag
reduction (in %) for the discontinuous slot at roof end configuration at
U0 = 20 m s-1 (ReL = 1.4 9 10
6)
Fig. 18 Influence of the control parameters (Stj and Cl) on the drag
reduction (in %) for the continuous slot at slant top edge end
configuration atU0 = 20 m s-1 (ReL = 1.4 9 10
6)
Fig. 19 Influence of the control parameters (Stj and Cl) on the drag
reduction (in %) for the winglets at roof end configuration atU0 = 20 m s
-1 (ReL = 1.4 9 106)
Fig. 20 Influence of the control parameters (Stj and Cl) on the drag
reduction (in %) for the discontinuous slot at slant top edge
configuration atU0 = 20 m s-1 (ReL = 1.4 9 10
6)
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By comparing the Figs. 17and 19, and Figs. 18 and 20,
one can observe that changing the jet exhaust geometry for
a fixed forcing location dramatically changes the control
behavior. Changing discontinuous slot for winglets on the
roof position suppresses the local optimum of drag reduc-
tion identified at Stj = 0.62, and exchanging continuous
slot by discontinuous slot also suppresses several optimal
frequencies like Stj = 0.35.It is also interesting to notice that the most efficient
frequency is about two times the natural frequency of the
shear layer when the perturbation is located upstream of the
separation, while it is close to the natural frequency when
the perturbation is close and downstream of the separation.
Table3 summarizes the better drag reductions and
corresponding parameters for each configuration. The
continuous slot at slant edge is clearly the most efficient
configuration with a much smaller optimal Cl. From the
industrial point of view, it is also important to notice that
most of the perturbations lead to significant drag reduction
so that it is possible to choose the right flow controlstrategies depending on the location where it can be inte-
grated in the vehicle.
4.2 Influence of the forcing parameters on local
pressure
In order to explain the drag reductions, mean local static
pressure has been monitored at various locations using
pressure sensors. Only a few sensors have been used during
control tests, so those measurements can only give general
trends. Static pressure results are presented in the same waythan previous drag reduction measurements: for each pair
of parameters, the corresponding local static pressure var-
iation DCp = (Cp0 - CpC)/Cp0 is reported. Pressure coef-
ficients without and with control are, respectively, noted
Cp0 and CpC.
4.2.1 Wall pressure at slant upper edge:
Figure21 presents slant upper edge local pressure varia-
tions for the same parameters and configurations as the one
presented on Figs.17,18,19 and 20 for the drag. One can
observe strong similarity with the drag variation for thediscontinuous slot configurations at slant top edge (by
comparing Fig. 21a with Fig. 17) and roof end (by com-
paring Fig.21d with Fig. 20): graphics are nearly identical.
This means that, for both those configurations, the drag
reduction is strongly connected with pressure recovery in
this particular area.
For the winglets configuration, pressure variation results
are also quite similar to the drag variation results but with
lower pressure variations. For the continuous slot config-
uration at slant top edge, similarity can only be detected
for two local areas Stj & 0.6/Cl & 0.3 9 10-3 and
Stj & 1/Cl & 0.5 9 10-3 but with small pressure variations.
Table 3 Better drag coefficient reductions for each configuration and
associated parameters
Pulsed jets configuration DCX (%) Stj Cl (10-3)
Discontinuous slotroof 7.8 1.03 3.1
Continuous slotslant edge 7.5 0.35 0.3
Wingletsroof 6.9 1.13 1.8
Discontinuous slotslant edge 6.3 1.22 1.7
Fig. 21 Influence of the control
parameters (Stj and Cl) on the
pressure coefficient evolution
(in %) on the top of the rear
slant, for the four flow control
configurations
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4.2.2 Wall pressure in the middle of the rear slant
Figure22describes local pressure variations at the centerof the slant. Here again, comparison with the drag reduc-
tion results highlights very similar trends and relation
between drag and wall-pressure variations. The two dis-
continuous slot configurations (Fig. 22a, d) and the wing-
lets configuration (Fig.22b) show areas with pressure
benefits larger than those observed on drag (3040%). The
pulsed blowing through continuous slot configuration
(Fig.22c) also shows areas of pressure variations of almost
15% for frequency close to Stj & 0.6.
These measurements confirm that the discontinuous slot
at roof end reduces the drag through pressure recovery over
the rear slant. As expected, drag reductions with the otherconfigurations also induce an increase in the pressure dis-
tribution over the rear slant.
4.2.3 Wall pressure in the middle of the rear end
Rear blunt local pressure variations are plotted in Fig. 23.
Here again, comparison with drag reduction mappings
in Figs. 17, 18, 19 and 20 enables to locate where a par-
ticular configuration produces benefits. Except for the
Fig. 22 Influence of the control
parameters (Stj and Cl) on the
pressure coefficient evolution
(in %) on the middle of the rear
slant, for the four flow control
configurations
Fig. 23 Influence of the control
parameters (Stj and Cl) on the
pressure coefficient evolution
(in %) on the middle of the rear
end, for the four flow control
configurations
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discontinuous slot at roof end configuration (Fig. 23a), all
other three configurations (Fig.23b, c, d) show similar
trends than the drag reduction results. For those specific
configurations, the drag reduction can then be linked with
the pressure recovery over the rear-end surface of the body.
4.2.4 Conclusion on local pressure variation results
In light of previous results, it is now possible to clarify the
influence of each particular control strategies on the nearwake of the Ahmed body:
The pulsed blowing through the discontinuous slot at
roof end has a strong influence on the wall pressure in
the top and middle of the slant. This suggests an action
mainly on the recirculation bubble.
The continuous slot at slant top edge seems mainly to
modify the wall pressure on the middle of the slant and
on the rear blunt part of the model, suggesting a
modification of the ring-shaped vortical structure.
The pulsed blowing through winglet-type jets acts on
the wall pressure over the slant with a strongest effectnot only in the middle of the slant but also in the rear
blunt part of the model. One can hypothesize a
modification of the shear layer starting from the top
of the slant and, as a consequence, a stronger interac-
tion between the shear layer and the torodal recircu-
lation on the rear part of the model.
As for the winglets configuration, the discontinuous slot
at roof end modifies the mean pressure value at each
location (top and middle of the slant and rear vertical
surface), suggesting a complex interaction between the
main structures of the wake.
In order to highlight previous results, the near-wake
modification has been investigated for the best control
strategy, i.e., the pulsed blowing through the discontinuous
slot at roof end (Fig. 17), and compared with the base flow
without control.
4.3 Near-wake modification by pulsed blowing
Time-averaged total pressure loss coefficient mappings
in a vertical cross-section located at a relative distance
DX/H = 0.5 behind the model for ReL = 1.4 9 106 are
presented on Fig. 24, for both the base flow and the con-
trolled flow. The controlled flow result is the one associated
with the discontinuous slot at roof end, with the better set
of parameters defined in Table 3.
One can notice several differences between the natural
and the controlled flows:
The total pressure loss coefficient area associated with
the slant recirculation bubble is clearly reduced, whichconfirms the effectiveness of the pulsed blowing on the
roof end boundary layer separation and then on the
recirculation bubble. It also confirms that the recircu-
lation bubble is reduced when the pulsed blowing
reduces the drag.
As expected (Aider et al.2009; Fourrieet al.2011), the
control system acts also on the longitudinal vortices: a
pressure loss drop happens at the edge of the structure,
while losses in the core seem to weakly increase. One
Fig. 24 Half planes of total pressure loss coefficient distributions in the near wake of the modela without control,b with control andc the Cpidifference between the two cases
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can hypothesize that this is an effect of the complex
interaction existing between separation bubble and
longitudinal structures: the reduction of the former
allows the later to develop with a smaller size but a
stronger intensity.
Pressure losses become also a bit more important in the
blunt recirculation. Here again, it is probably a
consequence of the interaction with the slant separa-
tion. Various authors (Roumeas et al.2008; Pujals et al.
2010) report that the cores of the blunt recirculation
move downstream when the slant recirculation issuppressed: one can made the assumption that blunt
recirculation cores come closer to the measurement
plane when control is activated and the recirculation
bubble suppressed, with the result of more apparent
pressure losses.
The standard deviations associated with the unsteady
total pressure coefficients are plotted on Fig. 25. Without
control, separation bubble exhibits strong pressure varia-
tions (Fig. 25a). In the controlled case, fluctuations only
remain on the longitudinal structures and, even if reduced,
on the underbody flow (Fig. 25b).
5 Conclusions
Time-averaged and time-dependent base flow around a
standard Ahmed body with 25 slant angle has been
characterized in wind tunnel. Mean flow and drag results
are in accordance with previous studies. A significant
Reynolds effect has been observed in both the drag
coefficient and mean pressure distribution on the rear end,
due to a reduction of the recirculation bubble.
Unsteady measurements in the rear-end flow reveal
three mechanisms that can be characterized by a constant
value of reduced frequency. The more organized, linked
with the ring-shaped structure observed in the wake, is
characterized by a Strouhal number (based on the rear
vertical height), StHv & 0.31.
In the shear flow region that separates the recirculation
bubble from the external flow region, two unsteady orga-
nizations have been highlighted. One is characterized by aStrouhal number (based on the slant height) Sth & 1.2 and
is associated with the natural KelvinHelmholtz instability
of the shear layer. The other is due to the flapping of the
shear layer and is characterized by a Strouhal number
(based on the length of the recirculation bubble)
StLr & 0.17 at a Reynolds number ReL = 1.4 9 106. At
this specific Reynolds number, the Ahmed model exhibits a
high-drag coefficient characterized by a large separation
bubble along with energetic streamwise vortices. This
Reynolds number has then been chosen to carry out flow
control experiments focused on slant recirculation, without
any attempt to control longitudinal structures.The influence of rear-end periodic forcing on the drag
coefficient has then been investigated using electrically
operated magnetic valves in an open-loop control scheme.
Four distinct configurations of flow control have been
tested: pulsed jets in a discontinuous slot or in winglets on
the roof end and in a discontinuous or continuous slot at the
top of the rear slant. For each configuration, the influence
of the forcing parameters (non-dimensional frequency,
injected momentum quantity) on the drag reduction has
Fig. 25 Half planes of total
pressure loss coefficient
fluctuations in the near wake of
the modela without control and
b with control
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been examined, along with their impact on the static pressure
on both the rear slant and vertical base of the model.
Maximum reductions between 6 and 8% have been
measured depending on the geometric and jet exhaust
configurations that show different sensitivity to the forcing
parameters.
Indeed, for the jets pulsing through discontinuous slots at
roof end, the amount of injectedmomentum quantity seems tobe the key parameter, and there is a weaker influence of the
pulsation frequency. On the opposite, the vortical sheet puls-
ing through continuous slot at slant top edge leads to a sig-
nificant drag reduction only for a particular frequency range.
For the two last configurations (jets pulsing through winglets
at roof end or through discontinuous slots at slant top edge),
bothjet velocity and pulsation frequency seemto be important
to optimize the drag reduction.
However, the influence of other parameters has to be
investigated. The control of the amplitude of the jet
velocity signal is probably an important feature, while the
study of the duty cycle may be a promising way in order toreduce the needed momentum quantity. In the present
study, an important overshoot has been observed, and it
would be interesting to highlight its effect on the flow
control.
One can also notice that when the perturbations are
close to the separation the blowing frequencies that pro-
duce the best results are close to the KelvinHelmholtz
instability frequency of the shear layer or between the
flapping frequency and the KelvinHelmholtz instability
frequency of the shear layer. This is in accordance with the
work of Sigurdson (1995) on the effect of a periodic
velocity perturbation on the separation bubble downstream
of the sharp-edged blunt face. On the contrary, the most
efficient frequencies when the flow is perturbed upstream
the separation are about two times the natural shedding
frequency. This is of course different from the work of
Sigurdson where there is no incoming boundary layer.
Meanwhile, further investigations need to be done to
highlight in our case the impact of the velocity perturbation
on the entrainment of flow and/or growth rate of the shear
layer and the impact on the reattachment length.
Acknowledgments This work was carried out in the framework of
the CARAVAJE project supported by the Agence pour le Devel-oppement Et la Matrise de lEnergie (ADEME). We thank the
Renault SA and PSA Peugeot-Citroen Aerodynamics Research teams
and the Plastic Omnium research team for fruitful discussions.
Technical support by the S4 Wind Tunnel team is also gratefully
acknowledged.
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