25 Degree Ahmed Body aerodynamic study

8/9/2019 25 Degree Ahmed Body aerodynamic study

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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


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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


J.-L. Aider (&)

PMMH Laboratory, UMR 7636, CNRS, ESPCI ParisTech,

10 rue Vauquelin, 75231 Paris, France


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)

Exp Fluids (2012) 52:11691185 1173

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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)


reduction (in %) for the continuous slot at slant top edge end

configuration atU0 = 20 m s-1 (ReL = 1.4 9 10

6)


reduction (in %) for the winglets at roof end configuration atU0 = 20 m s

-1 (ReL = 1.4 9 106)


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




(in %) on the middle of the rear

slant, for the four flow control

configurations




(in %) on the middle of the rear

end, for the four flow control

configurations

Exp Fluids (2012) 52:11691185 1181

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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

1182 Exp Fluids (2012) 52:11691185

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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

Exp Fluids (2012) 52:11691185 1183

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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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25 Degree Ahmed Body aerodynamic study