Base Flaps and Oscillatory Perturbations to
Decrease Base Drag
Tsun-Ya Hsu, Mustapha Hammache, and Fred Browand
Aerospace & Mechanical Engineering, University of Southern California, Los
Angeles, CA 90089-1191
Abstract
The objective of this investigation is to study possible means for reducing the base
drag of a tractor-trailer. The experiments are conducted in the Dryden wind tunnel
at the USC Ground Vehicle Aerodynamics Laboratory. A roughly 1/15 scale
model resembling a trailer is utilized for the study. The model is fitted with a
shaped nose-piece to ensure attached flow over the forward portion of the model.
The model is equipped with a force balance to measure drag. In addition base
pressures are measured, and hot-wire wake surveys are conducted downstream
from the model base. The Reynolds numbers (based on the square-root of the
model cross-sectional area), range from 0.1 x 106 to 0.4 x 106.
Drag reduction is effected by means of flaps attached along the edges of the
model base, and inclined inward to decrease the size of the downstream wake. In
addition, an oscillatory perturbation is applied at the flap origin in an attempt to
maintain attached flow for larger angles of flap inclination.
The present study has found that a simple, passive base-flap deflection—no
forcing whatsoever—produces significant drag saving. The maximum drag
reduction is 0.06 – 0.08 at an angle of 9-10 degrees. The magnitude of the saving
is in accord with both early and recent measurements in other laboratories.
The present results also show that oscillatory momentum addition has little
effect on drag reduction unless the net oscillatory momentum flux coefficient is
equal or greater than 0.1%. Increasing the oscillatory momentum perturbation to a
coefficient value of 0.3% produces drag savings at angles greater than 9-10
degrees, but has very little effect upon the maximum saving at 9-10 degrees.
To further study this anomalous behavior, follow-on experiments are planned to
investigate a larger range of forcing amplitudes, and a variety forcing-function duty
cycles. In addition, Digital Particle Image Velocimetry will be used to capture the
detailed flow-field in the vicinity of the flap.
304
T.-Y. Hsu, M. Hammache, and F. Browand
Introduction
The objective of this research is to provide guidance for the reduction of base drag
of a typical tractor-trailer. Roughly speaking, the total drag of a tractor-trailer can
be broken down into a contribution from the tractor and undercarriage, a
contribution from the trailer base and a much smaller contribution resulting from
skin friction along the (long) trailer sides. As truck tractors become more
streamlined, and other sources of drag are minimized, the base drag of the trailer
will remain and assume even more importance.
As a hypothetical example, assume a tractor-trailer with a drag coefficient CD ª
D
, where D is the measured
0.7. The drag coefficient is defined as C D =
1
2
rU A
2
drag; r is the air density; U is the wind tunnel speed, and A is the area of the
cross-sectional body perpendicular to the flow. In this case, the contribution from
the trailer base might be Cbased ª 0.14, which is 20 percent of the total drag.
Reducing trailer base drag by 50 percent would reduce total drag by approximately
10 percent. Now imagine continued improvements to the tractor that decrease the
total drag to CD ª 0.5. The base drag coefficient might actually increase somewhat
because the thinner boundary layer separating from the end of the trailer is more
efficient in entraining air from the base region. More efficient entrainment may
increase the base drag to a value Cbased ª 0.20. In this circumstance, trailer base
drag becomes approximately 40 percent of the total drag, and reducing trailer base
drag by 50 percent now represents a total drag reduction of 20 percent. The
observation that base drag and fore-body drag are inversely related was put forth by
Hoerner [5] based upon observations on aircraft shapes having blunt bases. The
argument has also been suggested to hold for heavy trucks (see Diebler & Smith in
this proceeding, and Saltzman & Meyer [10]).
There are two general approaches for the reduction of base drag. One approach,
we would term passive control, is to alter the geometry of the base region is some
way. The boat-tail-plate attachment studied by Lanser et al., [8], and illustrated in
Fig. 1, represents one possible alteration of base geometry. Measurements on a
full-scale tractor-trailer in the 80x120 Foot Wind Tunnel at NASA Ames indicate
total drag reductions of the order of 10 percent, and good performance at angles of
yaw (Lanser et al., [8]). Other more recent investigations of the boat-tail-plate
attachment include Khalighi et al., [6], and Storms et al., [11]. Storms et al., show
that total drag decreases from 0.263 without the boat-tail-plates to 0.215 with the
boat-tail-plates for a clean tractor-trailer geometry without wheels—the GTS
model. The total drag reduction is 18 percent.
The second approach is to attempt flow control by means of an active forcing
(such as an oscillating flap, or a blowing slot), meant to alter the boundary layer
properties—usually to avoid an unwanted separation. The blowing device
discussed by Englar in this proceeding and in [3], falls within this second category.
Our approach encompasses both a modification of the base geometry by means
of the addition of flat-panel flaps, and an additional active control by means of an
oscillatory mass flow perturbation within the boundary layer meant to delay flow
separation over the surface of the flaps. The (flat) flaps are attached to the trailer
Base Flaps and Oscillatory Perturbations to Decrease Base Drag
305
base along the trailer base edges, and are inclined to the free stream to close the
wake more efficiently, as in Fig. 2. They are attractive, because they can be folded
flat against the base when not in use.
Fig. 1 Boat-tail-plate attachment at a truck base (from NASA Ames [11])
Fig. 2 Base flaps attached to model truck base, recent test at NASA Ames
The addition of passive flaps to the base of a realistic tractor-trailer model was
first studied by Cooper [2]. Cooper shows that the drag coefficient for a typical
straight-sided truck decreases from 0.78 without flaps to 0.72 with flaps. The
optimum flap angle is within the range 10-15 degrees. Further, Cooper
demonstrates that the effectiveness of the drag reduction increases with nondimensional flap length, l, defined as l = Lf/sqrt(A), where Lf is the flap length and
A is the cross-sectional area of a truck. He points out that most of the drag
reduction is accomplished for flap length less than a value of approximately, l =
0.18.
With regard to active flow control, Nishri and Wygnanski [9] reveal that an
oscillatory jet introducing momentum flux with a net zero mass flow is more
effective than a steady blowing jet in delaying flow separation over a flapped
306
T.-Y. Hsu, M. Hammache, and F. Browand
airfoil. They are concerned primarily with increasing lift by means of oscillatory
suction/blowing. Their oscillatory jet, generated by a loudspeaker, is located at the
origin of the flap as illustrated in Fig. 3 from Nishri and Wygnanski. They show
that the effectiveness in delaying flow separation is determined by the location of
the jet, the frequency of the induced oscillation, the net momentum flux coefficient,
and the shape size of the slot. The frequency of the induced oscillation, f, is nondimensionalized by the flap length and free-stream velocity, U∞. It is defined as
f* Lf
F+ =
. The net oscillatory momentum flux coefficient, Cμ, is defined as
U•
Uj 2
g
*(
) , where g is the slot height, i.e., the gap between the flap and
Lf
U•
the side wall, and Uj is the amplitude of the oscillatory jet fluctuation. Favorable
delays in separation are obtained for slot heights of the order of 1/3-1/2 of the
incoming boundary layer displacement thickness, and for Cμ values less than 0.1%.
The most effective non-dimensional frequencies appear to lie in the range F+ =
0.3—1.5.
We wish to carry over these previous insights and technologies, and to apply
them for the purpose of base drag reduction. Specifically, we wish to address the
following questions.
• For the flat-panel base flap geometry applied to our model, what is the optimum
flap angle to achieve drag reduction?
• Will the oscillatory jet delay the separation of the flow, resulting in additional
drag reduction?
• If the oscillatory jet has the positive effect of drag reduction, what will be the
magnitude of the drag reduction?
• What are the critical parameters for the application of active flow control?
Cm = 2 *
Fig. 3 Sketch of the flap (from Nishri & Wygnanski [9])
Base Flaps and Oscillatory Perturbations to Decrease Base Drag
307
Experimental Apparatus
Wind Tunnel
The USC Dryden closed-circuit wind tunnel facility consists of a settling chamber,
contraction, test section, diffuser and axial fan section. The contraction represents
an area change of 10:1. The test section—6.1 meters in length—has an octagonal
cross-section. The dimension between any two parallel sides of the octagon is 1.37
meters. For the ground vehicle studies, a porous ground-plane 1.37 m in width and
5.8 m in length is installed in the test section. A small amount of suction is applied
over the surface of the ground-plane to maintain a thin boundary layer. The
maximum free stream velocity in the wind tunnel is approximately 30 m/s.
Truck Model
A bluff-body model consisting of a parallelepiped having a rounded nose is used to
study base drag reduction, Fig. 4(a). At this point, no attempt is made to model an
actual tractor-trailer. The width of the model is 0.197 m. The ratio of width:
height: length for the parallelepiped is [1:1.4:4]. It is mounted above the ground
plane a distance 50.8 mm. To this parallelepiped is attached a rounded nose of
length 0.934 m, as shown in blue in Fig. 4(a). The nose is carefully shaped to
avoid flow separation. The cross-sectional area of the model, A = 0.0535 m2, is
used as the reference area.
Model drag, side force and yawing moment are measured by means of a strain
gage sandwich designed specifically for this purpose. The model is mounted
directly to the upper surface of the sandwich. The lower surface of the sandwich is
anchored below the surface of the ground plane by means of the two protruding
pins seen in Fig. 4(a). Also seen within the model is a Scanivalve pressure
sequencer. There are fourteen pressure ports on each side of the model, and sixteen
pressure ports on the model base. Each of these ports is connected to the
Scanivalve and sequentially sent to a Baratron pressure transducer located outside
the wind tunnel.
The model incorporates flat-panel flaps along each edge of the base. In Fig.
4(a), the flaps are shown fabricated from acrylic plastic flat stock. To provide a
cleaner design and more reliable flap-slot geometry, we now construct the flaps
from 0.89 mm brass flat stock, as in Fig. 4(b). The flaps are soldered along the
four common edges to form a box-like structure attached to a plate mounted
directly to the base of the model. A total of eight independent sets of flaps at flap
angles 0, 4, 8, 12, 16.5, 20.5, 25 and 29 degrees are tested. Flaps can be doubled in
length by taping on the appropriate flap extension, as the second photograph in Fig.
4(b) illustrates.
A detailed picture of the model base is shown in Fig. 5. A schematic drawing
looking from the top is shown to the left. A commercial hi-fi loudspeaker made by
Vifa (frequency range from 35 to 4000 Hz) is used to produce the oscillatory
forcing. Motion of the speaker diaphragm in an otherwise sealed enclosure forces
air out through the slot gaps, g = 1mm, along all four edges of the base. Two flap
lengths, Lf = 5.4 and 10.48 cm, are studied. When normalized with the
308
T.-Y. Hsu, M. Hammache, and F. Browand
characteristic scale, ÷A = 0.231 m, the non-dimensional flap lengths become,
respectively, l = 0.233 and 0.453).
Hot Film Anemometer
In order to measure the amplitude of the oscillatory jet, a hot-film anemometer is
adopted (see Blackwelder 1981 [1] for details of anemometers). A TSI Model
1210-10 general-purpose probe is used for single component velocity
measurements. The probe is calibrated by varying the wind tunnel speed as
measured by a Pitot-static tube connected to a second Baratron.
(a)
(b)
Fig. 4 USC model equipped with rounded nose and base flaps
Base Flaps and Oscillatory Perturbations to Decrease Base Drag
309
Fig. 5 Details at the base
Experimental Method
Experimental Conditions
All data sets are recorded at wind tunnel speeds between 13 and 25 m/s. Reynolds
numbers, based on the square-root of the model cross-sectional area are,
respectively 0.11 x 106 and 0.36 x 106. A sine wave is chosen as the forcing
function. The forcing frequency is zero or within the range 40—600 Hz. The
corresponding non-dimensional frequencies are, F+ = 0, 0.17—3.93. The
oscillatory momentum flux coefficient, Cμ, ranges from 0 (no forcing) to 0.3%.
Data Acquisition and Data Processing
Signals from the force balance, the Baratron, and the hot-wire probe are digitized
using Labview. The nominal sampling rate for the force measurements in the wind
tunnel is 100-1000 Hz for a total sampling time of 5-8 seconds for each data point.
The sampling rate for the pressure measurements is similar. The averaging time of
5-8 seconds is sufficient to obtain accurate mean values. Previous observations of
wind tunnel performance (Zabat et al., [12]) have shown that virtually all of the
wind tunnel turbulence (including low frequency unsteadiness) lies above 0.2 Hz.
Typically, 5-second averages of the free stream velocity lie within .0015*U• of the
long-time mean. Hot-wire signals are normally sampled at about 20 samples per
cycle for a suitable number of cycles.
The rms amplitude of the unsteady velocity in the vicinity of the slot gap is
measured outside the wind tunnel by placing the hot-wire a distance of about one
gap width downstream from the slot. This measurement is needed to determine the
magnitude of the applied forcing (the value of Cμ). Observations at several
positions along the length of the slot verify that the influence of the corner is
restricted about two slot widths. The slot flow is two-dimensional within ± 8%.
310
T.-Y. Hsu, M. Hammache, and F. Browand
The rms fluctuation amplitude that determines the value of Cμ is measured at
ambient pressure on a bench test, as is common practice [4], [7], [9]. The particular
level of ambient pressure has no influence on the measurement of rms fluctuation
since all surfaces are exposed to the same uniform pressure. In the wind tunnel,
however, the situation is slightly different. One side of the driving speaker is open
to the slot gap while the other side is exposed to a static pressure equivalent to a
pressure along the underside of the body. The speaker must be driven against this
pressure difference. In our case these pressures are both close to the wind tunnel
static pressure, and no correction to Cμ has been made. A more thorough in situ
calibration of forcing is planned.
Fig. 6 shows the large variation in rms amplitudes that are a consequence of the
dynamic response characteristic of the speaker plus enclosure at the different
applied frequencies. The largest rms amplitudes occur in two periods—one at
about 50 Hz and the other centered at about 140 Hz.
The hot-wire is also used to determine boundary layer thickness at the trailing
edge of the model. The boundary layer there is turbulent with a displacement
thickness of approximately 2 mm at a wind tunnel speed of 16 m/s. The hot-wire is
used to measure the downstream wake of the model truck for a variety of flap
angles. For this purpose, the hot-wire is positioned at the mid-height of the model
in a cross stream plane 76.2 mm behind the flaps.
Amplitude (V)
10
9
rms vel
(m/s)
8
5.5
5
4.5
4
3.5
3
2.5
2
1.5
1
0.5
0
7
6
5
4
3
0
100
Frequency (Hz)
200
300
Fig. 6 Strength of the oscillatory jet for 1 mm gap with sine wave forcing
Base Flaps and Oscillatory Perturbations to Decrease Base Drag
311
Results and Discussion
Passive Control
We first examine the addition of flat-panel base flaps in the absence of the
oscillatory jet forcing. The effect of flap angle upon drag reduction is shown in
Fig. 7. The horizontal axis represents the flap angle in degrees, and the vertical
axis represents the change in drag coefficient, DCD = CDnoflaps - CDflaps,. A positive
value of DCD corresponds to a reduction in drag. Data is shown here from three
widely different experiments. The two curves in red are from our wind tunnel
experiment on a model having an aerodynamically clean fore-body at Reynolds
numbers of 0.23 and 0.36x106. The experimental reliability is indicated by the
scatter of data points representing multiple trials.
The blue curve is an earlier mentioned result from Cooper [2]. Although the
flaps are similar, the model truck geometries are different. His flaps are attached to
the base of a truck model along the vertical edges and the top edge. The bottom
edge has an attached flap, but this flap is not deflected. The result shown
represents flaps on a model truck having straight sides, wheels and a trailer with
rounded nose tested at a Reynolds number of Re÷A ≈ 1.5x106.
The data shown in black is from a recent experiment performed in the NASA
Ames 12-foot pressurized wind tunnel at a Reynolds number of 6x106, which is
comparable to a Reynolds number for a full-scale truck at highway speeds. In this
case, the model is a generic truck shape having a gap between tractor and trailer,
cab extenders and simple axles with wheels. Multiple data points again give an
indication of the measurement reliability.
In spite of the different model geometries, all of the DCD curves have a roughly
similar shape. Drag savings first increase and then decrease with increasing flap
angle. Thus there is an optimum angle for maximum drag saving. The maximum
saving is in the range DCD ª .06 - .08 for all three experiments. Overall, the three
data sets suggest a robust drag saving that is neither particularly dependent upon
the details of truck shape nor upon the Reynolds number of operation. However,
the angle for maximum saving does seem to depend upon Reynolds number. The
optimum angle seems to increase steadily from 9-10 degrees at 0.3x106 to perhaps
20 degrees at 6x106.
The drag reduction resulting from base flaps can also be observed in the wake
profile as a reduction in the wake momentum deficit. Wake velocity profiles are
shown in Fig. 8 for flap angles of 0 and 12 degrees. The horizontal axis represents
the span direction, non-dimensionalized by the square-root of the model crosssectional area. The model centerline is at the zero span position. The vertical axis
represents the non-dimensional velocities. The blue curve indicates the wake
profile for flap angle equal to zero. It can be seen that the wake width is the order
of the model width. The red curve shows the wake profile for flap angle equal to
12 degrees. The wake width is now considerably narrowed. The small velocity
increase at mid-span may actually represent a recirculation (backflow) along the
center-plane. The single hot-wire used here is not capable of determining flow
direction and may thus “demodulate” the recirculation.
312
T.-Y. Hsu, M. Hammache, and F. Browand
Active Control
Now consider the addition of oscillatory blowing/suction along the edges of the
model base, coincident with the origin of the flap. The purpose of the oscillatory
blowing/suction is to attempt to maintain attached flow over the flap for larger flap
deflection angles, and thereby to further reduce drag. Fig. 9 summarizes the best of
our results. The drag coefficient difference, DC D, is now shown for two flap
lengths—with and without oscillatory blowing. The two blue curves represent
flaps of length 0.233 and 0.453 in the absence of forcing. The drag saving peaks at
about 9 degrees as in Fig.7, and doubling the flap length has no significant effect
on the result. The results in red, for a momentum coefficients of Cμ = 0.15% and
0.3%, demonstrate the effectiveness of oscillatory blowing/suction. There is a
modest additional drag saving at the higher flap angles, but the application of
forcing has very little effect in the vicinity of peak saving. Thus the peak saving is
almost unchanged. The cause of this unexpected result is still under investigating.
Examination of the forced and unforced results in the vicinity of 10 degrees
suggests that the momentum addition does delay the onset of separation in some
manner. However, we had anticipated that in the presence of forcing and delayed
separation, the drag saving would continue to increase beyond the unforced peak at
9-10 degrees. Rather, the demonstrated effect of forcing is to broaden the peak
region by extension to larger angles but not to increase the maximum value of drag
saving.
For the shorter flap length, the optimum forcing frequency definitely lies in the
range F+ ª 0.6 – 0.7. For the longer flap, the forcing frequency may be slightly
below optimum. Here the non-dimensional value F+ ª 0.33 is dictated by the need
to stay in the frequency band near 40 Hz to achieve the highest momentum
coefficient possible.
Our experiments suggest that there is a rather broad range of non-dimensional
forcing frequencies F+ ª 0.3 – 1.0 that will effect flow modification, but that
forcing amplitude is critical. We begin to see noticeable change only at momentum
coefficients of the order of 10-3 or greater (Cμ ≥ 0.1%). Nishri & Wygnanski [9]
show observable increases in lift over deflected flaps at momentum coefficients as
small as 0.01%. Evidently drag reduction for a three-dimensional body such as
ours is a much more subtle proposition.
Base Flaps and Oscillatory Perturbations to Decrease Base Drag
Fig. 7 Effect of base flap angle on total drag
Fig. 8 Effect of base flap angle on the momentum defect in the wake
313
314
T.-Y. Hsu, M. Hammache, and F. Browand
Fig. 9 Effect of oscillatory forcing on drag decrease
Conclusions
Present Study
The present study describes the addition of flat-panel flaps to all four sides of the
downstream end of a bluff-body model resembling a tractor-trailer. The nose of the
model is rounded to prevent unwanted upstream separation (and boundary layer
thickening). It is possible to form partial answers to the questions posed at the end
of the introduction.
• The flaps—when deflected to flap angles of 9-10 degrees—produce a total drag
reduction of DCD ª 0.075, as measured by a force balance within the model. The
drag reduction is broadly consistent with reductions measured earlier by Cooper
[2], and with more recent measurements acquired in the 12-foot wind tunnel at
NASA Ames.
• The amplitude of the oscillatory forcing applied at the origin of the flaps is most
important. Forcing has little effect at small values of the oscillatory momentum
coefficient, Cμ = 0.1%. Forcing at amplitudes of Cμ = 0.3%, does produce
modest additional drag reduction at larger flap angles, but does not alter the peak
drag reduction at 9-10 degrees.
• The forcing does appear to delay separation over the flap.
• Non-dimensional forcing frequencies in the broad range, F+=f*Lf/U∞ ª 0.3 – 1.0,
do appear to provide the greatest effect on the flow, although this is a tentative
conclusion.
Base Flaps and Oscillatory Perturbations to Decrease Base Drag
315
Modifications to Present Experiment
To date, the results from oscillatory forcing leave several questions unanswered.
We believe oscillatory forcing should lead to greater drag savings. This view is
supported by the more recent work of Kjellgren et al., [7], who investigated drag
changes for a two-dimensional V-22 model wing in hover mode. In this case, the
bluff-body is the wing itself with down-flow normal to the upper surface of the
wing, as would be achieved in hover mode (with tilt rotors above the fixed wing
providing the down-flow). Forcing at the leading and trailing edges near flap
junctures does delay separation to larger flap deflection angles, and appears to
provide DCD of as much as 40-50 percent. What is evident from this work is that
drag improvements may require larger forcing amplitudes—of the order of Cμ
≥0.01.
The present experiment will be modified in several important respects to allow
an expansion of the range of forcing parameters—particularly the forcing
amplitude. The present loudspeaker arrangement cannot achieve Cμ values beyond
0.3%, when all four sides are forced. A means for forcing at higher amplitudes is
currently being investigated, and will be incorporated into the redesign of the
model.
We plan to investigate forcing-signals other than sinusoidal. For example, a step
in speaker voltage produces a blowing impulse response or a suction impulse
response as sketched in Fig. 10(a). If either the suction or the blowing portion of
the cycle should provide more effective control (there is anecdotal support for the
importance of suction), one or the other can be emphasized by means of saw-tooth
signals as in case (b).
Finally, we plan DPIV studies (Digital Particle Image Velocimetry) of the flow
adjacent to the flap with and without forcing. The intent is to answer the question
regarding the state of the boundary layer and the degree of attached/unattached
flow present.
Fig. 10 New designs: (a) Example of equal blowing and suction response, (b) Saw-tooth
emphasizing either blowing response or suction response
Acknowledgement:
Support from Dr. Sidney Diamond, Chief of Technology Development, Heavy
Vehicle Systems, Office of FreedomCAR and Vehicle Technologies, is gratefully
acknowledged. Additional thanks to Dale Satran and to NASA Ames for providing
the recent experimental results for base flaps operating at full-scale Reynolds
numbers.
316
T.-Y. Hsu, M. Hammache, and F. Browand
References
1. R. Blackwelder, “Hot-Wire and Hot-Film Anemometers,” Methods of
Experimental Physics: Fluid Dynamics, Vol. 18 Part A, 1981.
2. K. Cooper, “The Effect of Front-Edge Rounding and Rear-Edge Shaping on the
Aerodynamic Drag of Bluff Vehicles in Ground Proximity,” SAE, 850288, 1985.
3. R. Englar, “Advanced Aerodynamic Devices to Improve the Performance,
Economics, Handling and Safety of Heavy Vehicles,” SAE, 2001-01-2072, 2001.
4. R. Grife, A. Darabi & I. Wygnanski, “Download Reduction on a Three
Dimensional V-22 Model Using Active Flow Control,” AIAA, 2002-3071, 2002.
5. S. Hoerner, Fluid Dynamic Drag, self-published, Library of Congress Card
Number 64-19666, Washington, D.C., 1965.
6. B. Khalighi, S. Zhang, C. Koromilas, S. R. Balkanyi, L. Bernal, G. Laccarino &
P. Moin, “Experimental and Computational Study of Unsteady Wake Flow
behind a Bluff with a Drag Reduction Device,” SAE, 2001-01-1042, 2001.
7. P. Kjellgren, D. Cerchie, L. Cullen & I. Wygnanski, “Active Flow Control on
Bluff Bodies with Distinct Separation Locations,” AIAA, 2002-3069, 2002.
8. W. Lanser, J. Ross & A. Kaufman, “Aerodynamic Performance of a Drag
Reduction Device on a Full-Scale Tractor/Trailer,” SAE, 912125, 1991.
9. B. Nishri & I. Wygnanski, “Effects of Periodic Excitation on Turbulent Flow
Separation from a Flap,” AIAA Journal, Vol. 36, No. 4, April 1998.
10.E. Saltzman & R. Meyer, Jr., “A Reassessment of Heavy-Duty Truck
Aerodynamic Design Features and Priorities,” NASA TP-1999-206574, 1999.
11.B. Storms, J. Ross, J. Heineck, S. Walker, D. Driver, & G. Zilliac, “A n
Experimental Study of the Ground Transportation System (GTS) Model in the
NASA Ames 7- by 10-Ft Wind Tunnel,” NASA TM, 2000.
12.M. Zabat, S. Frascaroli, F. Browand, “Drag Measurements on a Platoon of
Vehicles,” UCB-ITS-PRR-93-27, California PATH, University of California,
Berkeley, 1993.