A wind tunnel test was conducted to compare the aerodynamic and propulsive performance of a slot distributed propulsion model resembling a "jet-wing" concept to a conventional circular nozzle arrangement. Testing found that the slot configuration resulted in an average 27% increase in drag and detrimental effects on lift compared to the nozzle. As a result of these initial performance penalties, the propulsive efficiency of the slot was 2.4% lower than the nozzle. However, the slot was able to partially fill the aircraft's wake and showed potential benefits, though not enough to overcome the initial drag penalty. The results suggest the jet-wing concept may be inferior to a simpler nozzle system but indicate wake-filling as an area for further study regarding possible

1. 1
Practical Testing of the Slot Distributed
Propulsion Concept
Alex Loveday supervised by Prof. George Carter
Abstract—A wind tunnel test was performed to directly com-
pare a slot distributed propulsion model akin to the ‘jet-wing’ to
a circular nozzle arrangement. The aim was to resolve disparity
between the results of a recent 3D computational study and
previous 2D studies.
Through a momentum theorem approach, a general theory of
the targeted wake-filling effect was derived and then compared
to experimental results. Testing examined consequences of im-
plementing a jet slot, and showed an average 27% drag increase
in the jet-off case. A detrimental effect on lift was also observed.
Due to this initial deficit, the propulsive efficiency of the slot was
found to be 2.4% less than the nozzle at steady flight conditions.
Slot propulsive efficiency and aerodynamic improvements were
measured and showed to be linked to wake-filling. The principal
of wake-filling was directly observed and shown to offer potential
benefits. However, the slot was unable to overcome the initial
drag penalty it suffers, suggesting the ‘jet-wing’ concept to be
inferior to a simpler nozzle arrangement. Lift characteristic
results proved inconclusive, suggesting an area of further work
to investigate potential high lift benefits.
Index Terms—Distributed Propulsion, Electric Aircraft, Jet-
Flap, Jet-Wing, Slot Trailing Edge Jet, Wake-Filling, Wind
Tunnel Testing.
NOMENCLATURE
˙m Mass flow rate.
ρ Density.
σ Dimensionless resul-
tant force.
Fx Axial resultant force.
Ae Exit area.
b Wingspan length.
bJ Jet height.
bW Wake height.
c Chord length.
CD Drag coefficient
CL Lift coefficient.
CP 0 Total pressure coeffi-
cient
f Fuel to air ratio.
nP Propulsive efficiency.
pe Exit pressure.
p∞ Free-stream pressure.
T Thrust.
UJ Jet velocity.
UW Wake velocity
U∞ Free-stream velocity.
I. INTRODUCTION
In 2011, ‘Flightpath 2050 - Europe’s vision for Aviation’
[1], a report produced by the European Commission, set
ambitious targets for the aviation industry to meet by 2050,
including a 70% reduction in CO2 emissions and a 90% re-
duction in NOx production. At the United Nations Framework
Convention on Climate Change (UNFCCC) talks in 2008, the
International Air Transport Association signed an agreement to
work towards halving CO2 emissions by 2050 and to stabilise
its production within aviation by 2020 [2].
With aviation expected to contribute $1 trillion to world
GDP by 2026 and its employment by three and a half times
more than any other industry, [3], its continuation is not in
question.
As illustrated by the reversible Joule-Brayton cycle, the
ideal thermal efficiency of a gas-turbine is limited by its peak
combustion temperature [4]. Ignoring material limitations, this
cannot be continuously increased without producing mono-
nitrogen oxides, formed in volumes increasing exponentially
with the combustor inlet air temperature [5]. If a 50 % reduc-
tion in CO2 production requires the engines’ fuel efficiency
to double (assuming an inverse proportionality between the
two) then it is clearly not realistic to find this margin through
improvements in the gas-turbine alone.
Despite the future of aviation fuel availability being in
question [6] and EU imposed emission penalties of heavy
CO2 producers [7], the aviation industry must prosper through
continual developments in not only an aircraft’s airframe and
aerodynamics, but importantly within its propulsive system.
The scale of the challenges laid out has led to aircraft
propulsion being rethought from its foundations upwards. Con-
sidering electrically powered aircraft suggests some inherent
benefits; through using batteries, potential energy imparted to
liquid fuel during take-off and ascent is not lost as a result
of exhausting its mass at altitude, as occurs in turbo-fans.
Lithium-Air battery technology has now advanced to achieve
a competitive energy density as a replacement for jet fuel -
43.2MJ/kg [9] compared to 46.4MJ/kg gained from kerosene
[10] - and could therefore facilitate the targets set out by the
European Commission and UNFCCC. An electrical system
can also be divided into multiple smaller units, a series of
identical motors each with individual fans and power inputs
realising several benefits including:
• improving propulsive efficiency through an increased
effective bypass ratio [8];
• eliminating the need for superconductive cabling to a
single large power output device;
• reduced maintenance effort thanks to smaller, easily re-
placeable modules;
• reduced structural demands through avoiding a single
total weight load at the centre of thrust; and
• improved redundancy margin for an engine out scenario
It is clear that an electrical system such as this can begin
to realise the benefits of distributed propulsion, providing
powered flight within fixed-wing aircraft about which thrust
forces and airflows are distributed around the vehicle.
This study focuses on a slotted form of distributed propul-
sion known as the ‘jet wing’. Coined by Küchemann and illus-
trated in Figure 1, it seeks to create a synergistic integration of
an aircraft’s propulsion system and airframe, thus improving
the propulsive efficiency of the vehicle as a whole [11].

2. M.ENG RESEARCH PROJECT APRIL 2015 2
Fig. 1: Illustration of Küchemann’s ‘jet-wing’ concept [11].
Multiple two-dimensional (2D) computational fluid dynam-
ics (CFD) studies on jet wing cross sections/self propelled
aerofoils exist, showing propulsive efficiency benefits to be
possible in this computational representation. The work fo-
cuses on truncating the trailing edge of an aerofoil to facilitate
an exhaust of some finite width, and then determining both
the effects on the aerofoils performance and the propulsive
efficiency of the aerofoil as a self propelled body.
A recent three-dimensional (3D) CFD model of a jet-wing
implemented within a blended wing body (BWB) aircraft has
produced results in direct disagreement with the previously
accepted literature [15]. This contrary view is supported by
only a single study, but its 3D computation allows a more
real-life approach to ascertaining whether a slotted propulsion
system can produce a propulsive efficiency benefit for an
aircraft.
While the latest computational study is yet to have the
validity and accuracy of its results proven, it has raised
the question whether the jet-wing concept is as tenable as
concluded by 2D studies.
It has therefore become of value to resolve this disparity
through practical wind tunnel testing. This test seeks to ascer-
tain whether a slot distributed propulsion system can provide
a propulsive efficiency benefit when implemented in a manner
akin to the Küchemann jet wing. Real life limitations were
applied to the model, reflecting important parameters similar
to the vehicles observed in computational studies, and allowing
arguments pertinent to the concept’s applicability in a full
aircraft design to be formed.
II. THEORY
A. Distributed Propulsion
Distributed propulsion systems provide powered flight
within fixed-wing aviation vehicles about which thrust forces
and airflows are distributed around the vehicle, with the aim
of increasing the propulsive efficiency of the system.
The concept operates by dividing up the thrust load among
several propulsive units and can be realised through a variety
of scenarios. In its broadest view however the applications can
be divided into two major categories consisting of ‘leader’ and
‘follower’ arrangements [12].
1) The Leader method uses multiple engines with all units
contributing directly to the flight thrust. This is a pro-
posed method for some electrically propelled aircraft
using relatively small fans mounted in/upon the main
wing of the aircraft.
2) The Follower method uses a main unit to generate power
whilst there is at least one propulsion system acting as
secondary unit. A proposed follower arrangement is the
Rolls-Royce / EADS system, seen in Figure 2.
Fig. 2: Propulsion system of the Rolls-Royce and EADS ‘Advanced Hybrid
Distributed Propulsion Concept’ [8]. A large power generating gas turbine can
be seen in the rear tailplane, with multiple electric fans converting the power
to thrust via a battery system. Note that the fans can also act as secondary
generators during descent.
Distributed propulsion is well suited to electric aircraft as
each propulsive unit can draw upon a separate power line,
eliminating the cost of larger central cabling approaching a
power rating in the tens of megawatts (MW). Furthermore the
batteries can be located close to the electric motors, reducing
transmission losses and cable weight.
B. Propulsive Theory
As previously mentioned, distributed propulsion can im-
prove propulsive efficiency by increasing the effective bypass
ratio. In quantitative terms, it increases the mass flow through
the propulsive system.
The Froude Propulsive Efficiency, ηP , is defined as the ratio
of useful propulsive power to the rate of kinetic energy added
to the flow [13].
ηP =
TU∞
˙m[(1 + f)(U2
J /2) − (U2
∞/2)]
(II.1)
where: T = thrust, U∞ = freestream velocity, f =fuel to air
ratio, UJ = exhaust jet velocity and ˙m = mass flow rate.
The thrust that is produced by an engine in freestream flow
is given by [13]
T = ˙m[(1 + f)UJ − U∞] + (pe − p∞)Ae, (II.2)
Ae = exhaust exit area, and pe and p∞ are the exhaust exit
and ambient pressures respectively.
If the exhaust pressure is assumed to be equal to the ambient
pressure, and the fuel mass flow is neglible (or zero for an
electical system) then the ratio in Equation II.1 can be written
as
2

3. M.ENG RESEARCH PROJECT APRIL 2015 3
ηP =
˙mU∞(UJ − U∞)
1
2 ˙m(U2
J − U2
∞)
(II.3)
which simplifies to
ηP =
2
1 + UJ
U∞
(II.4)
Therefore the propulsive efficiency depends only on the ratio
of the exhaust jet and freestream velocities. For a constant
flight-speed, a propulsive efficiency improvement requires a
decrease in jet velocity. To achieve this without compromising
the thrust magnitude, Equation II.2 shows that an increased
mass flow is required.
The mass flow can be linked to the jet velocity and geometry
by continuity:
˙m = ρAeUJ . (II.5)
It can be seen that when targeting a propulsive efficiency
improvement, obtaining a higher mass flow with a decreasing
exhaust velocity requires an increase in exhaust exit area.
C. The Jet-Wing
Küchemann’s Jet wing (Figure 1) was a design first pro-
posed by Dietrich Küchemann in which the exhaust from
embedded engines is ducted out of the trailing edge of the
aircraft wing [11]. The design aims to employ a synergistic
integration of the aircraft’s propulsion and structural compo-
nents, whereby the wake of the airframe in the free-stream flow
is ‘filled in’ by the engine exhaust. Less net kinetic energy is
left in the flow behind the vehicle, increasing the system’s
propulsive efficiency.
This system is therefore expected to offer improvements
over a conventional circular nozzle independent of the air-
frame.
Several studies have analysed this concept in 2D using
CFD. The work focuses on truncating the trailing edge of an
aerofoil to facilitate an exhaust of some finite width, and then
determining the system’s performance as a self propelled body.
Schetz et al. [14] performed a variance of the trailing edge slot
hight, showing that a propulsive efficiency benefit is possible
with an increasing slot height.
There is a detrimental affect on the aerofoil as a result of
truncating its trailing edge - in the simple case of the modified
aerofoil in a free-stream flow with no propulsive jet a drag
penalty is observed . The established work shows that this is
eliminated in the jet-on case, and the system begins to benefit
from wake filling to reduce the aerofoil drag co-efficient
and improve propulsive efficiency compared to an unmodified
system. Successful wake-filling is qualitatively demonstrated
through examination of the wake/exhaust velocity profile pre-
dicted by the computational model [14][15][16].
Schetz et al. and Kim et al. developed the use of the slotted
exhaust as a jet-flap by vectoring the thrust downwards. Both
observe increases of lift coefficient of up to 200% at the cost
of increasing drag [14][16].
The latest study by Hudson agrees with the jet-off 2D
analysis; a previously unnoticed lift reduction is also seen.
Although evidence of a degree of wake-filling is observed, the
jet-on drag shows no improvement over the baseline aerofoil,
and an additional reduction in lift is measured due to upward
deflection of the suction side flow [15].
This is in direct disagreement with Schetz et al., who
showed the improved wake profile to produce favourable drag
characteristic. Additionally, Schetz et al. found the aerofoil
pressure distribution to be unaffected by the jet, allowing a
constant lift coefficient.
A 3D study conducted by Hudson targets a more real-world
implementation of the theory with some limitations applied.
The study shows a 2.6% reduction in lift and 5.7% increase
in drag for the no jet case, and a 11.3% reduction in lift and
3.8% increase in drag for the jet-on case [15]. The practical
limitations applied limit the quality of wake-filling achieved,
and again the effect of the jet is shown to be non localised.
This investigation of the jet-wing argues it does not outperform
a conventional circular nozzle and airframe system.
D. Wake Filling
The suggested improvements in propulsive efficiency arise
from the idea that the exhaust jet from the trailing edge fills in
the wing wake, an approach already implemented in ships and
submarines. Less kinetic energy is left in the flow behind the
airframe, improving propulsive efficiency. Alternatively, wake
filling can be viewed as reducing momentum loss, decreasing
the drag upon a system at a constant airspeed. These are the
benefits that the jet-wing seeks to obtain.
Ko et al. derived a mathematical formulation of the prin-
cipal [18]. The vehicle drag and engine thrust are no longer
considered separate quantities balanced at a later stage; the
net force on the whole system is derived for a separate jet
and jet-wing case. Ko et al. solved the steady flight case (ie
zero resultant force) for the jet velocity to show the difference
in propulsive efficiency. Their result shows an increasing
propulsive efficiency benefit with increasing height of the jet
relative to the wake. For the jet wing, if the jet height is
equal to the wake height then the wake is perfectly filled,
ie Uj = U∞ and an ideal propulsive efficiency is achieved.
A wind tunnel test would conduct sweeps of flight param-
eters including thrust levels and airspeeds; it would not be
practical to target only the steady flight case. Furthermore,
many potential benefits of the jet-wing may be realised in the
high thrust case, such as noise reduction or utilising a jet-flap
for short take-off capability. To allow examination of potential
efficiency gains at high or low thrust, the model of Ko et al.
will be expanded to include a net resultant force.
The problem is set up in Figure 3, as per Ko et al.’s
work on the steady flight case, by considering a 2D control
volume around each system. For simplicity the wake is given
a square profile, and the downstream pressure assumed to be
undisturbed from the upstream ambient pressure.
From the momentum theorem and conservation of mass, the
resultant force vector, F, is given by Equation II.6 [18].
Fx = −
S
(p − p∞)dS −
S
ρq(U∞ + q)dS (II.6)
3

4. M.ENG RESEARCH PROJECT APRIL 2015 4
(a) The seperate jet case
(b) The jet-wing case
Fig. 3: Control surface around both propulsor configurations [18].
Here, p = pressure at the boundaries, q = velocity perturba-
tion from U∞, with componenets {u, v, w}T
, ρ = density and
S = control surface.
Following the assumptions stated in the problem set-up, this
simplifies to
Fx = −
h
−h
ρu(U∞ + u)dy . (II.7)
First we consider the separate jet case. Performing the
integration in Equation II.7 around the control surface in
Figure 3a gives
Fx = ρ[(U∞ − UW )UW bW − (UJ − U∞)UJ bJ ] , (II.8)
and after retaining the resultant force Fx and re-arranging
we obtain
1
2
Fx
1
2 ρU2
∞bJ
=
bJ
bW
−1
UW
U∞
1 −
UW
U∞
−
UJ
U∞
2
+
UJ
U∞
(II.9)
Fx is a force per unit span of the body; thus the quantity
on the left of Equation II.9 is dimensionless, in keeping with
the right hand side. We can therefore define the parameter
σ =
Fx
1
2 ρU2
∞bJ
, (II.10)
to allow a non-dimensional examination of the effect of a
net resultant force on wake filling. This parameter captures
the ratio of the created thrust force to the force due to the
free-stream dynamic pressure acting over the same jet height.
Substituting Equation II.10 into Equation II.9 and re-
arranging to standard quadratic form:
0=
UJ
U∞
2
−
UJ
U∞
−
bJ
bW
−1
UW
U∞
1 −
UW
U∞
−
1
2
σ
(II.11)
Solving for UJ /U∞:
UJ
U∞
=
1
2
±
1
2
1 + 4
bJ
bW
−1
UW
U∞
1 −
UW
U∞
−
1
2
σ
(II.12)
If we take UJ /U∞ ≥ 1, the positive solution is
applicable. For a real solution, we can also state that
2 bJ
bW
−1
UW
U∞
1 − UW
U∞
+ 1
2 ≥ σ. As σ is negative in the
thrust direction and the left hand side is always positive, there
is no limit to the model for the high thrust (T 0 =⇒ Fx <
0) case. When there is net drag, the limitation stated above
applies. Given the independence of the jet and wake, UW /U∞
can be taken as a constant for a particular body. From Equation
II.12, we can subsequently see that to decrease the jet speed
UJ , a large jet height, bJ , is desirable, although for a given
thrust to optimise the jet speed would require bJ /bW → ∞. It
can also be seen that an increased net thrust (ie σ decreasing
or becoming more negative) results in a higher jet speed.
Taking the positive solution of Equation II.12 and substitut-
ing for UJ /U∞ in Equation II.4 gives the propulsive efficiency
of the separate jet case with a resultant force:
ηP =
2
3
2 + 1
2 1 + 4 bJ
bW
−1
UW
U∞
1 − UW
U∞
− 1
2 σ
(II.13)
Now consider the jet-wing case seen in Figure 3b. Applying
Equation II.7 to this case gives
Fx =−ρ[bW (U2
W −U∞UW )+bJ (U2
J −U2
W +UW U∞−U∞UJ )]
(II.14)
and after retaining the resultant force Fx and re-arranging
we obtain
1
2
−Fx
1
2 ρU2
∞bJ
=
UJ
U∞
2
−
UJ
U∞
+
UW
U∞
UW
U∞
− 1
bJ
bW
−1
− 1
(II.15)
Again we arrive at a non-dimensional resultant force term.
Substituting for this with Equation II.10 and re-arranging to
standard quadratic form:
0=
UJ
U∞
2
−
UJ
U∞
−
UW
U∞
1 −
UW
U∞
bJ
bW
−1
− 1 −
1
2
σ
(II.16)
4

5. M.ENG RESEARCH PROJECT APRIL 2015 5
Solving for UJ /U∞:
UJ
U∞
=
1
2
±
1
2
1 + 4
UW
U∞
1 −
UW
U∞
bJ
bW
−1
− 1 −
1
2
σ
(II.17)
As before, we take UJ /U∞ ≥ 1 so the positive solution is
relevant. Again, the model’s solution is versatile in all high
thrust cases, where Fx < 0. We can bound the lower force
level of the model: 2 UW
U∞
1 − UW
U∞
bJ
bW
−1
− 1 + 1
2 ≥ σ.
Considering Equation II.17, for a given thrust level it
is possible to minimise UJ by arranging bJ /bW = 1, an
observation also made by Ko et al. [18]. With the addition of
interaction between the jet and wake, achieving UW /U∞ = 1
will also optimise UJ . Both of these objectives correspond to
perfect wake filling.
Substituting the positive solution of Equation II.17 into
Equation II.4 gives the propulsive efficiency of the jet-wing
with a resultant force:
ηP =
2
3
2 + 1
2 1 + 4 UW
U∞
1 − UW
U∞
bJ
bW
−1
− 1 − 1
2 σ
(II.18)
It is worth noting that in the special case of steady flight
(Fx = 0 ∴ σ = 0) the analytical expressions of Equations
II.13 and II.18 match those focused on the steady flight case
by Ko et al.. The effect of a net force on the efficiency margin
introduced by the jet wing is shown in Figure 4. Although there
is still a difference, the propulsive efficiency benefit of the jet-
wing reduces asymptotically to approximately a 1% increase
at high thrust.
The σ positive region is not relevant to any self-propelled
body as it represents a deceleration. However, it may be
applicable to passive implementations of wake-filling, such
as trailing edge ducts fed from leading edge or pressure
surface intakes. Subsequently, to examine exhaust velocities
UJ /U∞ < 0.5, it may be worth expanding this analysis to the
negative solutions of Equations II.12 and II.17.
III. EXPERIMENTAL DESIGN
A model was designed to directly compare a slot distributed
propulsion system, akin to the jet-wing, to a simpler and
more conventional nozzle system. It used interchangeable
components, one utilising a slot and the other containing an
embedded circular nozzle. To facilitate the trailing edge jet an
embedded engine was used. For simplicity the test section was
designed as a single aerofoil profile extruded to a sufficient
aspect ratio.
A. Model Design
Extruded polystyrene was chosen for the bulk body material
so that the internal geometry of the model could be freely
crafted. The material also offers good strength and stiffness
when laminated. Model aircraft electric ducted fans (EDF)
were found to be most appropriate to power the exhaust
Fig. 4: Propulsive efficiency with varying non-dimensional net force, σ. Here,
UW /U∞ = 0.7 and bJ /bW = 0.13 as measured for the model in Section
IV-B. Note that in the positive region, σ > 0, the body is no longer self-
propelled.
jet. Due to their target use in hobby remote-controlled (RC)
aircraft they are extremely compact for the power output
they provide, and are well suited for embedding within a
model. Their 3-phase brushless motors draw large currents,
normally supplied for short periods by lithium-polymer (LiPo)
batteries. To achieve good test functionality, heavy duty lead-
acid batteries were used instead. Their capacity allowed for
up to 54 minutes of run time at full EDF power output.
The early design tackled several conceptual decisions:
• wing chord;
• profile thickness;
• choice of aerofoil;
• wing span;
• airspeed;
• number of EDFs; and
• diameter and power of
EDFs.
A parametric model was set up to balance these quantities
by modelling the performance of the aerofoil and the EDF.
Further considerations included:
• a large enough chord to allow for sufficient lead in length
for the exhaust duct;
• finding a suitable slot width and area per EDF to prevent
high duct losses;
• having sufficient thickness in the aerofoil profile to embed
the EDF;
• obtaining a high enough aspect ratio to prevent finite wing
effects dominating;
• designing for a wind speed that would allow for good
performance from both the wing section and the EDF;
• using a large enough fan to produce enough thrust without
resorting to multiple fans unless otherwise needed due to
cost; and
• choosing an aerofoil appropriate to the above considera-
tions and implementation of the slot exhaust.
These dependencies were summarised into the design struc-
ture matrix (DSM) shown in Table I, which was used to guide
the design approach.
5

6. M.ENG RESEARCH PROJECT APRIL 2015 6
TABLE I: Optimised DSM for the model parameters, highlighting clustered
modules. Interactions below the leading diagonal represent feedback loops,
which required an iterative design approach.
The chosen solution uses a NACA 4424 aerofoil with a
0.5m chord and a 1.25m wingspan. A single 70mm 1900W
EDF was selected, drawing a maximum of 80A through a
100A electronic speed controller (ESC). A servo-tester was
used to control the ESC; it controlled the pulse width used
by the power electronics which acted as a measurable throttle
setting.
Schetz et al. noted that a thick aerofoil is preferable, as
a smaller percentage chord truncation is required to give the
desired slot height [14]. This also suited the embedded EDF.
To help with duct design, the slot exhaust area was made
equal to the fan exit area. Hudson argues practical limitations
when creating an exhaust slot along the trailing edge of an
aircraft [15], and concludes a 40% coverage to be reasonable.
In line with this conclusion, the slot was made 0.5m wide in
the centre of a 1.25m test section. The slot height is 7.7mm,
ie bJ /c=1.54%, sufficient to allow wake filling [14]. Including
material height above and below the slot opening, this required
a 6.64% chord truncation of the aerofoil. This is replaced by a
trailing edge pivoting mechanism, which allows the slot height
to be precisely set and the slot to be closed to an improved
edge.
The slot component required a duct to shape the cylindrical
flow from the fan to the narrow slot at the trailing edge. This
was optimised using CFD in SolidWorks Flow Simulation. As
the optimisation was only comparative to the previous design
iteration, a mesh refinement study was not required. Figure 5
illustrates the final duct profile.
The ratio of slot span to lead in length was maximised to
give greater freedom in other parts of the design. The optimi-
sation targeted a uniform exit velocity profile and minimum
dynamic pressure loss. The comparative baseline nozzle duct
consisted of a cylindrical thrust tube with constant area to the
fan exit. The completed wind tunnel model can be seen in
Figure 6.
B. Experimental Set-up and Method
The test was performed in the Durham University 2m Wind
Tunnel. The model was suspended inverted (to prevent a
zero weight zone [19]) by a computer controlled overhead
strut, via a 6-axis 600N rated force balance. This was mated
externally via a rig and strong points in the model to provide
sufficient airflow clearance. The model position was then
adjusted to provide good pitch range and minimise ground
Fig. 5: Flowlines in the X-Y profile of the optimised duct. To further control
the flow the duct shape was also optimised in the X-Z plane.
Fig. 6: The completed tunnel test model, with two interchangeable trailing
edge components. Note the embedded fan in the leading edge.
effects. Although the model was within one chord height of
the ground and the fan power lines introduced a possible force
bridge, the comparative approach of the test allowed for some
external influences as long as they were constant between the
slot and circular nozzle. A 5-hole probe on a three-axis arm
was used to perform velocity and pressure traverses.
Force balance readings were logged at 1024Hz for 16
seconds to restrict errors to within ±2.5% and ±1.27% in
the x (thrust) and z (lift) directions respectively, to a 97.5%
confidence. Probe results were logged at the same frequency
for 4 seconds to give a ±2.03% error to the same confidence.
The logging durations were a trade-off between accuracy
and model functionality; ideally longer readings would have
been taken but the total length of probe traverses and pitch
sweeps was limited by battery capacity and the EDF motor
rating. All component control and data logging routines were
performed using the Durham University Software for Wind
Tunnels (DSW).
The tests performed fall into several categories:
1) Tare tests: Measurement of balance and rig weight,
balance and rig weight with aerodynamic forces, and
balance, rig and model weight for both slot and nozzle
models.
2) Engine-out tests: Incidence sweeps of the nozzle and
slot models, and with the slot closed - EDF off.
6

7. M.ENG RESEARCH PROJECT APRIL 2015 7
3) Slot/nozzle performance: Static thrust (wind off) throt-
tle sweeps of slot and nozzle models.
4) Propulsive performance: Wind on thrust sweeps and
constant thrust wind-speed sweeps for nozzle and slot
models.
5) Engine-on aerofoil performance: Incidence sweeps
at fixed thrust and wind-speed for nozzle and slot.
Repeated at maximum thrust.
6) Probe traverse: 5-hole probe traverses along span of
trailing edge, and axial traverses examining the wake.
Engine on and off.
C. Data Analysis
The post-processing routine applied calibration files to mea-
surements from each test. Additionally, the aerodynamic forces
on the rig and balance were removed non-dimensionally -
this allowed for differences between general tests and the tare
test which measured the tare forces. The combined weight
of the balance, rig and model was also accounted for. These
operations were performed using the DSW.
The circular nozzle jet was assumed to have no effect on
the aerofoil characteristics; therefore its performance coeffi-
cients could be measured separately in an engine-off test and
assumed constant. Due to wake interactions in the slotted jet
model, the coefficients will vary with the introduction of the
EDF exhaust.
The static thrusts of the circular nozzle and slot were
measured, allowing the performance of the slot to be found as
a function of the nozzle’s in a matrix of ratios for each throttle
setting. In analysing each test, the nozzle model’s aerodynamic
coefficients were subtracted from the nozzle readings, leaving
the propulsive forces for the nozzle. From this, the propulsive
force of the slot could be found and subtracted from the slot
readings to leave the slot aerofoil effects for examination. By
this method the circular nozzle is treated as a baseline. The
flowchart in Figure 7 illustrates this process.
Fig. 7: Illustration of the process used to find force components with wake
interactions present.
IV. RESULTS AND DISCUSSION
A. The Engine-Out Scenario
Figure 8 shows jet-off aerofoil performance. Drag coeffi-
cient is consistently increased at all incidences compared to
the nozzle baseline, varying from a 26% penalty at −2◦
, to
24% at 14◦
. A maximum increase of 31% is seen at 3◦
,
and the average increase is 27%. This is a a result of the
imperfect trailing edge created by truncating the aerofoil;
flow separation at the trailing edge produces high energy
dissipation in resulting vortices, increasing momentum transfer
and therefore drag. Furthermore, using tuft flow examination
features on the trailing edge pivot mechanism were noted
to cause vortices at particular wind-speeds and incidences;
a subsequent performance penalty can be seen in Figure 8b,
where the slot’s Lift/Drag curve is impaired at an incidence
of approximately 2◦
.
(a) Graph showing lift and drag coefficients varying with incidence.
(b) Graph showing the variation of lift to drag ratio with incidence.
Fig. 8: Engine out aerofoil performance for the nozzle model, and for the slot
model with the slot both open and closed.
Figure 9 illustrates the wake behind the three scenarios.
It is apparent that a greater pressure loss occurs behind the
slot, corresponding to the increased drag. The pressure loss is
slightly reduced when the slot is closed; despite this improved
wake, Figure 8a shows that no significant or consistent drag
reduction results from closing the slot. It follows that the
geometry of the pivoting components is unable to completely
prevent flow detachment when approximating a trailing edge.
It also suggests the fixed pivoting mechanism components
accounted for a large part of the drag increase. Alternatively,
the benefits of the improved wake may be cancelled out by an
increased induced drag from larger flow deflection.
Figure 8a shows that the nozzle has a clear lift deficit
compared to the slot model, primarily because its suction
surface is significantly interrupted by the thrust tube over a
7

8. M.ENG RESEARCH PROJECT APRIL 2015 8
70mm span. Additionally, Figure 9 shows slightly less flow
deflection from the nozzle model. On the same note, the closed
slot produces a further increase in flow deflection. In Figure
8a we see that this produces a consistent lift improvement.
Fig. 9: Contour plots of jet-off wakes showing CP 0 for the nozzle (a), open
slot (b) and closed slot (c). Note that the nozzle wake was taken behind the
complete trailing edge, not the circular nozzle.
The presence of a thrust tube in the nozzle model prevents
a direct comparison between the two models to examine the
effect of chord truncation on lift. However, refining the slot’s
trailing edge with the adjusting mechanism gives an improved
wake, producing a better aerodynamic lift performance than
when truncated completely open. An average 3.4% increase
from the open slot is observed for the closed slot, peaking to
6.6% at zero incidence.
This observation is in agreement with the latest study
by Hudson, whose 2D analysis first noted a lift penalty in
association with aerofoil truncation in the jet-off case [15].
However, it is possible that manually closing the trailing edge
introduced an artificial camber if it did not follow the natural
taper of the aerofoil.
B. Propulsive Peroformance
To compare propulsive efficiencies the jet speeds had to
be found from force measurements. Assuming pe − p∞ = 0,
by substituting Equation II.5 for ˙m into Equation II.6 and
expanding we get
0 = ρAeU2
J − ρAeUJ U∞ − T . (IV.1)
This can then be solved for UJ /U∞:
UJ
U∞
=
1
2
+
1
2
1 +
4T
ρU2
∞Ae
. (IV.2)
After utilising the process described in Figure 7 to obtain
the slot and nozzle thrusts, the above equation yields exhaust
velocities from which the propulsive efficiency can be found.
Figure 10 shows the drag coefficient behaviour of the slot
model, found using the process in Figure 7. We can see that
drag co-efficient decreases with decreasing force parameter σ.
By Equation II.10, we see that this occurs with increasing
(more negative) net force, Fx, decreasing airspeed, U∞, or
a combination of both. As σ is dimensionless, the results of
varying either of these parameters can be compared.
Fig. 10: Drag coefficient of the slot model plotted against non-dimensional
net force (Equation II.10). The net force has been varied by sweeping the
thrust at a 15m/s wind-speed, and seperately sweeping the airspeed at full
throttle. At σ = 0, CD = 0.075.
As described in Section II-D, a more negative σ represents
a greater ratio of jet force to free-stream dynamic pressure
acting over the same area. The jet-stream has a greater energy
density compared to the airflow around it.
At the large positive value of σ = 3.05 (close to the engine-
out scenario examined previously) a drag penalty of 27.4%
is observed, consistent with the 27% increase measured in
Section IV-A. At the steady flight condition (σ = 0) an 18%
drag increase is still present. The overall deficit has reduced
by a third from the jet-off case; this exactly matches Hudson’s
3D computational model, who showed the drag increase varied
from 5.7% jet-off to 3.8% at steady flight, a change of a third
[15]. Following the assumption employed in this method that
the nozzle jet and wake are independent, the Nozzle drag is
constant. The slot only ‘breaks even’ from its initial deficit at
a large negative σ, when the slot is exhausting comparatively
high energy airflow into the wake. However, the downward
trend in Figure 10 does show drag reducing up to 48.9% as
significant wake-filling is achieved.
Although a value of σ should be achievable through any
combination of U∞ and Fx, the results of the wind-speed
sweep and throttle sweep do not align. Fx is defined as
the resultant force per unit span; to give this the measured
axial force was divided by the wingspan (1.25m). The whole
test section is therefore treated as homogeneous. However,
only 40% of the model is slotted and follows the behaviour
described derived in Section II-D; the inert model sections
will not perform independent of the parameters which non-
dimensionalise σ, and the behaviour is mixed as wind-speed
varies.
Applying Equations IV.2 and II.4 to the same experimental
data yields Figure 11, where the propulsive performances of
the slot and nozzle are compared.
8

9. M.ENG RESEARCH PROJECT APRIL 2015 9
Fig. 11: Graph of propulsive efficiency against non-dimensional net force for
the nozzle and slot models. Included for comparison is the nozzle with the
same drag penalty the slot incurs. The net force was varied through a wind-
speed sweep at constant throttle.
The slot’s propulsive performance suffers from the initial
drag penalty incurred from aerofoil truncation. To produce
the same resultant force at the same wind-speed more thrust
is required, so a higher jet speed is needed, reducing the
propulsive efficiency. At the steady flight condition (σ = 0)
the nozzle model’s propulsive efficiency is 73.2%, whereas the
slot achieves 70.8%.
As observed for the drag co-efficient, beneficial wake filling
occurs at more negative σ, where the influence of the slotted
jet is greater. To illustrate this, Figure 11 includes a plot for
the nozzle with its drag co-efficient artificially increased 27%,
corresponding with the slot’s default setback. We can see that
at large positive σ this brings it in line with the slot model.
As σ decreases, the slot benefits from wake-filling and starts
to improve over the offset baseline nozzle. As with drag, it
only equalises with the nozzle at very large negative σ.
Figure 4 incorporates measurements from the slot model
(UW /U∞ = 0.7, bJ /bW = 0.13, bJ = 0.0077) into the
theory derived in Section II-D for comparison against the
experimental results in Figure 11. It predicts more pronounced
effects of wake filling as σ increases. The experimental results
show the opposite however.
The theory assumes square wake profiles, which produce
the upper limit of efficiency gain possible; Ko et al showed
that assuming triangular profiles offers 34% of the efficiency
difference found by square profiles.
It is also assumed that the square profiles remain so, and
that no mixing occurs between wake velocity layers. Figure
12a shows the wake velocity profile at increasing distance
behind the slot trailing edge. It can be seen that the narrow
high velocity jet decelerates and spreads outwards further
downwind.
Figure 12 also shows the slot height to be too small; a large
momentum excess in a narrow area is produced by the jet, with
a significant wake momentum deficit remaining on either side
showing good wake filling is not achieved. Within a quarter
chord length however the jet has spread to occupy the whole
wake - Figures 12a and 12b show that UW /U∞ > 1 across
(a) Surface plot of the wake velocity profile. This shows the wake velocity,
U/U∞, across the vertical height, y/c, of the trailing edge slot, for increasing
distance downwind of the trailing edge, x/c. The non-dimensionalising length
c is the chord, and the surface is coloured by total pressure coefficient CP 0.
(b) Span-wise and axial contour plot of the wake showing total pressure
coefficient CP 0 and the traverse position in relation to the model.
Fig. 12: Wake investigation of the slot model - jet on.
the whole wake.
The jet-wake interactions seen in Figure 12 may explain
why wake-filling benefits were observed at large negative σ,
where the jet dominates due to its high relative energy density.
The narrow jet has a large enough velocity to dilute its energy
further downwind and replace the lost kinetic energy in the
wake. Figure 12 shows the total pressure of the jet dropping
while the lower total pressure wake on either side rises, cap-
turing the wake mixing as the jet transfers work through fluid
shear. This is not an ideal mechanism for replenishing the wake
kinetic energy however, as viscous energy dissipation increases
with the square of the velocity gradient [20]. Subsequently, in
Figure 12a CP 0 also decreases within the jet due to viscous
effects. At higher σ, the jet is too narrow and low energy to
achieve wake-filling.
The 1.54% of chord slot height is therefore too small to
produce wake filling benefits at high σ. Despite this, the
aerofoil truncation for the slot used in this test already causes
large drag penalties, which is responsible for compromising
9

10. M.ENG RESEARCH PROJECT APRIL 2015 10
the propulsive efficiency. Implementing a larger slot in pursuit
of improved wake filling would therefore likely only be
counter-productive. These consequences are more severe than
found by Schetz et al., whose finding of improving propulsive
efficiency with slot height led to the conclusion that drag
penalties were only of concern in an engine-out scenario [14].
The illustrations of their 2D analysis show the aerofoil wake
heights are small enough to be well filled by the slot. Figures
9 and 12 show that a greater wake height was found here
as a result of momentum transfer through boundary layers,
particularly along the suction surface. It may be that the
computational model employed by Schetz et al. was unable to
capture this boundary layer growth and the subsequent effect
on the wake.
C. Jet on Aerofoil Performance
Figure 13 shows aerodynamic performance of the slot model
aerofoil with the jet on. As predicted by the results of Section
IV-B, a slight drag reduction is found at zero incidence.
However, as incidence increases drag increases at a higher
rate than with the jet off.
Most noticeably, the lift-coefficient has increased by an
average difference of 0.24, ranging from 121% at −2◦
through
to 56% at 14◦
.
Fig. 13: Graph of lift and drag coefficients against incidence for the slot jet-on
aerofoil performance. For this test, σ ≈ −0.1 at zero incidence.
The jet may give recovery of the lift deficit from aerofoil
truncation observed in the jet-off case. 2D numerical analyses
reports the NACA 4424 aerofoil to have lift coefficients of up
to 1.3 at 15◦
[21], although no unmodified wing section was
tested to verify this as the model baseline. Taking account of
jet-on further lift penalties predicted by Hudson [15] and the
magnitude of the difference observed, this seems unlikely.
It is also possible the trailing edge mechanism gave the slot
jet a slight negative angle, giving it a positive component in
the lift direction subsequently measured as aerodynamic lift.
However this should have been captured and accounted for in
the static thrust ratio matrices described in Section III.
If a thrust vector was present but accounted for through
the data analysis technique, then Figure 13 may be displaying
practical effects of the ‘jet-flap’. Previously investigated com-
putationally by Kim et al. and Steiner et al., vectoring the slot
jet is reported to give large aerodynamic lift increases at the
cost of a modest drag increase by producing supercirculation
[16][14]. An examination of the slot exhaust wake and aerofoil
pressure distribution during this test would be required resolve
this.
D. Discussion of Model Suitability
Figure 14 shows the performance of the slot duct in
creating an even slot jet. Manufacturing flaws and surface
finish produce perturbations but there is no region of severe
exhaust disturbance that could affect the quality of wake filling
achieved.
Fig. 14: Plot of the exhaust axial velocity across the slot, as predicted by CFD
and measured by a 5-hole probe.
The model scale does not give equivalent Reynolds numbers
to conceptual commercial implementations of the jet-wing.
However, Schetz et al. demonstrated that the performance of
slotted propulsion is not Reynolds number dependant.
The model was finished in a thermo-softening laminate
which gave a good surface finish, but with some imperfections
such as small creases. These may have contributed to the
model wake, which was noted as more significant than found
computationally. However, due to boundary layer growth a
wing’s wake can be expected to be greater than the thickness
of its trailing edge, particular in thick aerofoils suited to the
jet-wing [22].
Although the EDF used in the model was factory balanced,
it was a source of substantial force noise. Force balance
readings were taken at at least twice the noise frequency to
prevent aliasing. For small drag components captured by the
z-axis strain gauge (which lies in the plane of the fan), noise
errors started to dominate; the small z-component of drag
was subsequently discounted where appropriate. The noise
magnitude was therefore a factor in restricting the testing to
a comparative nature, examining differences and trends rather
than absolute values.
V. CONCLUSIONS
A wind-tunnel model has been manufactured and tested to
directly compare a slot distributed propulsion concept to a
more conventional embedded wing thrust nozzle. The test was
10

11. M.ENG RESEARCH PROJECT APRIL 2015 11
of a comparative nature, with the conventional circular nozzle
treated as a baseline against which to measure the performance
of the slot.
Following examination of the results and comparison with
the theory, several conclusions can be made. The negative
effects of truncating an aerofoil to facilitate a jet-wing style
slot are well understood and verified to be more severe than
2D studies conclude. Practical limitations of implementing the
slot, such as maximum trailing edge coverage, also contribute
to this. A negative lift response in the engine-out case has
been measured and its cause examined.
Wake-filling has been observed through its influence on the
systems propulsive and aerodynamic performance, and more
directly through wake probe traverses. Propulsive improve-
ment trends were measured and linked to achieving wake-
filling. The circumstances in which maximum benefits were
found allowed the limitations of the derived wake-filling theory
to be examined. Direct examination of the influence of the jet
on the aerofoil wake suggests explanations for the nature of the
wake-filling benefits. The dimensionless force parameter, σ,
found through derivation of the high force wake-filling theory,
offers a physical interpretation of the effect of the jet on the
wake. Furthermore it allows the performance of wake-filling
to be evaluated across a range of operational circumstances.
Wake-filling was shown to improve propulsive efficiency but
these benefits were limited by the nature of its exploitation in
the jet-wing.
Measurement of jet-on aerodynamic performance of the slot
model showed a large lift improvement. If this result is val-
idated through further investigation, it could offer substantial
developments in aircraft design. The high lift to drag ratios
observed in this test offer the potential to greatly increase
wing loading, thus offering compensation for drag penalties
discussed previously. However, if this wing loading is jet-
dependent then an engine-out scenario may limit its use due
to lack of redundancy.
When compared to a simpler nozzle arrangement, the slot
model suggests that the jet-wing offers inferior propulsive
performance across all but the highest thrust ranges. Although
wake-filling benefits are achieved, the slot is unable to over-
come the initial drag penalty it suffers.
Given the magnitude of lift improvement noted in Figure
13, it is worth undertaking further testing to clarify the jet-on
aerofoil performance of the slot. As a part of this and utilising
the same model, slot exhaust vectoring as implemented by the
’jet-flap’ can be tested for benefits. Despite the limitations
of the jet wing, wake-filling has shown to offer synergistic
benefits; it is therefore recommended that a conceptual study
is undertaken to conceive and assess alternative propulsive
arrangements aimed at achieving wake-filling.
ACKNOWLEDGMENT
Thanks go to Prof. George Carter for his support and ‘go for
it’ attitude. The author would also like to thank Ken Ridgeway
for his enthusiasm for the project’s test model and the time
he invested in it. Additionally, many thanks to Josh Newbon
and Dr David Sims-Williams for their support during the test
week.
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