Reduction of Noise Emission of Suboptimal Operating ... - IAG

Reduction of Noise Emission of Suboptimal Operating Propellers
A. Hirner, Th. Lutz and E. Krämer
University of Stuttgart
Stuttgart, D-70569
Germany
Abstract
Propellers designed for cruising flight often suffer from an increased noise radiation if they are
operating under off-design conditions such as take-off and approach as well as reverse thrust
where partially high flow velocities and separations on the propeller blades can occur. This is
especially true for the airship Zeppelin NT where the propellers have to cover a very large pitch
angle range when used for manoeuvering on the ground. In cooperation with the project partners
Zeppelin Luftschifftechnik GmbH and Steinbeis Transferzentrum a combined theoretical and experimental
design approach is established which comprises modified vortex lattice methods for
preliminary design as well as CFD calculations for the analysis of the aerodynamic and aeroacoustic
behaviour of the propeller designs. To validate the theoretical results additional aerodynamic
and aeroacoustic tests with a propeller test bench are conducted. Besides load measurements the
detection of characteristic noise sources on a propeller blade is done by scanning the propeller
blade with an acoustic mirror. The project aims at the development of a time-efficient process
chain to reduce turn-around times in the design phase of new propellers. This comprises preliminary
design codes, subsequent highly automated propeller CAD design and grid generation for
the CFD calculations as well as structural analysis and milling of selected propeller designs.
Nomenclature
λ [mm] Wave length
f [Hz] Frequency
w 0
[mm] Half-width of the diffraction
pattern, u ∞ = 0
D [m] Diameter
G [dB] Effective gain factor
I M [ W m
] Sound intensity in the maximum
2
of the diffraction pattern
I Ref [ W m
] Sound intensity measured by
2
reference microphone
L M [dB] Sound pressure level in the
maximum of the diffraction pattern
L Ref [dB] Sound pressure level measured by
reference microphone
1 Introduction
The appearance of the aircraft turbofan came along
with a decreasing interest in developing new propeller
driven aircraft, especially for long-range routes and
the application of the propeller was mainly restricted
to small airplanes. With energy ressources becoming
shorter many leading aerospace companies are
currently investigating the potential of highly-loaded,
many bladed propellers operating at more and more
increasing flight mach numbers. Compared to turbofan
designs significant fuel savings are possible at the
expense of a higher noise radiation of propeller driven
aircraft. This effect is undesirable both in the cabin
in regard to passenger comfort as well as in the surrounding
area of airports. Current propeller research
is focused on understanding the noise generating mechanisms
and developing highly efficient, noise reduced
propeller designs. For modern turbo-prop aircrafts these
efforts usually result in many bladed propellers with
more or less swept blades. To gather an overview on
the basic design strategies references [1] and [2] may
be consulted.
The operation of the propeller in reverse mode is usually
restricted to a short-time period and therefore does
not play a major role in the design criteria. In contrast
the current propellers at the Zeppelin NT are operating
under partly extremly unfavourable conditions
with an extensive noise radiation when the airship is in
approach or manoeuvering on the ground. Within the
current project a highly automated process chain for

time-efficient design of noise reduced propellers is developed
which aims at the design of a new stern thruster
for the Zeppelin NT under particular consideration
of off-design conditions. To get new insights in the
noise generating mechanisms of the current 3-bladed
propeller both numerical and experimental studies are
conducted with the full-sale and a model propeller. In
addition a second propeller with a flexible mounting of
the blades at the hub is examined. This configuration
is said to have noise reducing effects as the blades can
oscillate to a certain amount if the blade encounters
unsteady inflow caused by atmospheric turbulence or
inclined inflow. This effect will be studied by operating
the propeller in inclined flight condition both with flexible
and fixed mounting of the blades.
2 Process chain for highly
automated propeller design
As for commercial aircraft there is a large need in increasing
efficiency as well as reducing noise emission
of propeller driven general aviation aircraft. This effort
arises due to a lower noise acceptance of the population
resulting in stricter noise regulations of local
airfields. The current state of the art propeller design
of small propeller manufacturers is usually limited to
preliminary design methods and broadly based on experience.
This approach cannot fully account for the
complex three dimensional flow field about a propeller.
Therefore within the current project detailed numerical
studies combined with experimental work are
conducted to develop new design criteria for noise reduced,
efficient propellers with special attention to offdesign
conditions. To ensure short turn-around times
of a propeller design there is a need for a highly automated
process chain comprising both numerical as
well as experimental studies. In the future parts of this
process chain can be supplied to local propeller manufacturers
as service work. The components of this
procedure according to Figure 1 will be presented in
the following.
Preliminary Design
XROTOR
CAD, Structural Analysis
CATIA
Experiment
Thrust, Torque, Visualisation,
Acoustics
New Noise
Reduced Design
Evaluation of Design
Numerical Flow Simulations
FLOWer / TAU
2.1 Preliminary design
For preliminary design studies the propeller code
XROTOR from Drela [3] is used. This code allows
the interactive design and analysis of free-tip as well
as ducted propellers. The propeller blade is represented
by the classical lifting-line theory whereas the aerodynamic
characteristics of the blade sections can
be modelled by specifying arbitrary airfoil aerodynamic
properties. To account for compressibility effects
a Prandtl-Glauert correction is applied. The induced
velocities are calculated by numerically solving for the
exact potential flowfield about the helical vortex field.
This approach is therefore valid for all advance ratios
and blade numbers compared to classical theories
which assume a small advance ratio and many blades.
A first-order correction for the change in the pitch of
the trailing sheets is applied to account for the propeller’s
self-induction which is important for highlyloaded
propellers. To examine more general propeller
geometries with non-radial lifting lines (raked or swept
blades) an additional discrete vortex wake formulation
is implemented to calculate the induced velocities.
Discrete line vortices on the rigid helicoidal wake
surface trailing from the lifting line into the farfield
downstream are used to calculate the velocitiy on the
propeller lifting line. Noise level contour plots can be
calculated on a rectangular grid based on a method
developped by Succi [4],[5].
This code is used for time-efficient investigation of the
aerodynamic and aeroacoustic behaviour of different
blade geometries both under design and off-design conditions.
This approach delivers important information
concerning necessary lift coefficients of the airfoils
as well as Reynolds- and Mach number distribution
along the blade at an early stage in the design process.
The rotation of the propeller blades and the resulting
centrifugal and Coriolis effects strongly influence the
boundary layer development on the blade especially
near the hub. Due to these effects unstalled flow can
be achieved at very high angle of attack and stall itself
occurs more gently. To account for these effects numerous
approaches were derived in literature which either
consider rotational effects by correcting calculated 2Dairfoil
polars [6], [7] or directly implement a modified
boundary layer model into the airfoil code [8].
Most of these 2D-corrections were derived for slow turning
wind turbines. Within the current project rotational
effects on the airfoil characteristics will be considered
by evaluating both CFD and experimental data
gathered with a propeller test bench (see Figure8).
Grid Generation
GRIDGEN
Figure 1: Flowchart of process chain
2.2 CAD design and structural analysis
The propeller geometry output from the XROTOR code
is imported by the 3D-CAD-CAM software package
CATIA by means of a macro. This procedure ensu-

es a time-efficient generation of a watertight propeller
surface which can be used as a basis for further
grid generation. As the subsequent grid generation for
the numerical flow simulations is script based as well,
it’s advantageous when each propeller geometry has
the same layout to be uniquely identified by the mesh
generating software independent of the detailed blade
geometry. The propeller surface is exported in IGESformat
which is the basis for mesh generation as well
as for milling the moulds of selected propeller designs
for further experimental studies. In addition CATIA is
used for the structural analysis of the propeller blades,
propeller shaft and clamping hub.
2.3 Grid generation
The computational mesh for the numerical flow simulations
is based on the IGES surface data of the propeller
blades and generated with the meshing software
GRIDGEN from POINTWISE [9]. For time-efficient
grid generation of multiple different blade geometries
and blade numbers this procedure was realised within
a script based approach. For first preliminary numerical
studies including interference effects of the propeller
blades with a motor cowling an unstructered
approach was realised first. In addition, a completely
structered procedure is persecuted. As no wall function
is applied in the CFD calculations a y + -value of
about 1 is anticipated for the first cell layer. In the
grid 30 cell layers are extruded from the blade surface
to represent the boundary layer on the blades. In addition
the cell distribution on the propeller blades and
the farfield and its extent are easily adaptable within
the script. This procedure is necessary as the numerical
flow simulations are the basis for the calculation
of noise levels both in the vicinity of the blade as well
as in the near farfield by evaluating the pressure fluctuations.
As no special CAA method is used it must
be granted that no erroneous sound levels result due
to inappropriate discretization of the farfield.
2.4 Numerical flow simulations
The numerical flow simulations are conducted with the
flow solvers TAU [10] and FLOWer, both developed by
the DLR.
TAU is a three-dimensional, parallel, hybrid, multigrid
code. It implements a finite volume scheme for solving
the Reynolds-averaged Navier-Stokes (RANS) equations.
The code uses explicit time stepping, the multistep
Runge-Kutta scheme and optionally an implicit
time stepping with a LU-time scheme. For accelerating
the convergence to steady state a local time-stepping
concept, different residual smoothing algorithms and
a geometrical multigrid method are implemented.
At the beginning the main focus of the project is to
examine different propeller configurations in regard to
noise emission and aerodynamic behaviour. In combination
with the preliminary studies with XROTOR
these numerical flow simulations offer detailed output
for the selection or design of adapted propeller airfoil
sections operating without separation under offdesign
conditions of the propeller. At this stage propeller
operation is restricted to axial inflow. Therefore
time-consuming unsteady calculations and the application
of the chimera technique are not necessary as a
quasi-steady approach is adequate. To consider interference
effects the cowling is modelled as an axisymmetric
body. To prevent a rotating boundary layer on
the cowling EULER boundary conditions are applied
to this surface whereas the propeller blades are calculated
fully viscous.
2.5 Experimental investigations
The design loop is closed by experimental studies conducted
with a propeller test bench where selected propeller
designs can be further tested. This facility offers
the possibility to thoroughly investigate the predicted
performance from the previous flow simulations
and serves simultaneously to validate the used codes.
This is especially important as the project is aimed
at the reduction of noise under off-design conditions
when partially high flow velocities and flow separation
on the propeller blades have to be expected.
Propeller test bench
The propeller test bench was constructed and built by
Steinbeis-Transferzentrum [11]. The propeller is driven
by a compact air-cooled servo motor. Propeller thrust
is measured with a load cell mounted in-between the
rear side of the motor and a rib in the motor cowling.
The motor is mounted rotatable about its length axis
so that propeller torque can be measured with a second
lateral mounted load cell (Figure 2). With an additional
turnable rig two prandtl sensors can be positioned
near the propeller plane, one sensor being positioned
ahead and the second one behind the propeller plane.
Evaluation of the static and dynamic pressure at these
two locations enables the determination of ciculation
at a specific radial position on the blade and therefore
offers additional information on possible separation in
addition to paint coatings. Within the current project
the propeller diameters to be examined are about 1m
und the rotational speed about 3250U/min. This rotational
speed at the test bench results from the chosen
scale factor between the current reference propeller of
the Zeppelin NT and the model propeller to guarantee
the same blade tip mach number of 0.5 between original
and model propeller which is essential to conserve
aeroacoustic similarity. Reynolds number effects between
both propeller scales can be determined by inve-

stigating the original propeller at a second test bench
with identical propulsion units from the Zeppelin NT.
velocities of about 7 m/s at the outside location of the
test bench.
Acoustic measurements
Figure 2: Propeller test bench for model propellers
Test facility Gust Wind Tunnel
The experimental studies are conducted at the Gust
Wind Tunnel of the institute. This wind tunnel was
concipated to simulate gusts on wind turbines by suddenly
opening and closing sails at the inlet of the tunnel.
The tunnel itself has a diameter of 6.3m and the
length of the test section is about 8m with a maximum
flow velocity of 18m/s which corresponds to the effective
inflow velocity at the stern propeller of the Zeppelin
NT. For acoustic measurements the test bench is
operated outside the test section to avoid disturbances
from the tunnel walls. For the Zeppelin NT noise levels
are highest when the airship is manoeuvering on the
ground and the propeller operates under nearly static
thrust conditions. The inflow area of the wind tunnel
was extended with a mobile frame construction with
nets of varying porosity fastened between the spars
(see Figure 8). Because of the missing contraction of
the wind tunnel itself this construction guarantees a
more laminar and steady flow in the test section and
the sensitivity of the tunnel concerning atmospheric
turbulence is reduced. On the other hand, the canalization
of the flow due to this device still offers flow
Acoustic noise level contour maps dependent on flow
veloctity and blade angle of the propeller will be captured
by ground microphones. This approach delivers
the total sound pressure level at a specific location but
offers no indication of the noise source distribution on
the blades. To clarify sound generating mechanisms
and to localize the dominating noise sources along a
propeller blade especially under off-design conditions
an acoustic mirror with elliptical shape was designed
which serves as a highly directional microphone system.
The characteristic of an elliptic mirror is to focus
sound waves emanating from a volume element
of a source distribution at the far focus upon the near
focus point where a microphone is located. Due to
constructive superposition of sound waves at the microphone
a gain is achievable dependent on the shape
of the mirror. Provided that there is no significant flow
between sound source and mirror both spatial resolution
and gain factor of the mirror system are mainly
limited by diffraction of the sound waves at the edge
of the mirror.
As mentioned before the acoustic measurements will
be conducted outside the test section of the wind tunnel.
In this case the test bench is mounted on an additional
framework to keep the propeller axis centered
in the tunnel axis. The acoustic mirror is positioned
ahead or behind the propeller and can be moved parallel
to the propeller rotating plane with a linear drive
unit. The focus point is adjusted scarcely above the
blade surface of a horizontally levelled blade. By moving
the mirror parallel to the propeller plane from one
blade tip to the other, the noise distribution of sources
emanating from the vicinity of the blade surface can be
captured. By positioning the mirror ahead or behind
the propeller noise sources emanating rather from the
upper or lower side can be detected (Figure 3). As the
distance of the focus point to the blade surface slightly
varies dependent on blade angle and airfoil thickness
distribution it has to be studied if a significant difference
between both positions can be detected at all.
In the design of the mirror some geometric restrictions
were given by the dimensions of the wind tunnel. The
diameter of the mirror should be kept small enough
to avoid any blocking effects on the propeller flow but
large enough to capture low frequencies in the range
of the blade passing frequency. The focal length of the
mirror is influenced by the desire of positioning the
mirror both in front and behind the propeller plane in
the inflow section of the tunnel (Figure 3).

Figure 3: Alignment of acoustic mirror
Theoretical relations were derived both for the gain
and the spatial resolution of an acoustic mirror by
Grosche [12]. The effective gain factor G according to
equation (1) describes the enhancement of the sound
intensity caused by the focusing effect of the elliptic
mirror at the mirror microphone compared to its free
field value measured by a reference microphone at the
same distance from the sound source. The ability of the
acoustic mirror to enhance sound intensities at the far
focus and therefore discriminate background noise makes
this system a highly directional sound microphone
system with the ability to resolve sound sources in all
directions normal to the mirror axis.
( )
IM
G = L M − L Ref = 10 log
I Ref
)
(1)
= 10 log
(0.5 D4
λ 2 B 2
Whereas the focusing effect decreases relatively slowly
in the direction of the mirror axis the maximum
in vertical and lateral direction is much more sharper.
Using the theory of diffraction of waves by a circular
aperture provides an approximate relation for the halfwidth
w 0 of the diffraction pattern of a point source.
According to Figure 4 w 0 is the distance from the maximum
within which the sound intensity decreases by
50%.
(2) w 0 = 0.54λ S D
In consideration of all parameters the final diameter
D of the acoustic mirror was set to 2.45m and the
focal length S to 4.5m with the distance of the near
focus B chosen to 1.06m.
The calibration of the acoustic mirror was performed
using a Kenwood KA 3080 R preamplifier and a broadband
loudspeaker vifa 10 BGS 119/8 which provides a
balanced frequency response. Both for the mirror and
reference microphone 1/2” Free-field Microphone Type
4190 from Brüel & Kjæer were used. For lack of
Figure 4: Principle of acoustic mirror [12]
a special acoustic room various problems due to reflections
and diffraction emerged during the beginning
of the calibration process. For this reason the principle
of the impulse response measurement of the software
package ARTA [13] was utilized to generate a
swept sine impulse of about 5 ms short enough to avoid
ground reflections at the microphones. The corresponding
ARTA measurement setup for single channel
measurement setup is shown in Figure 5. Due to the
single channel setup the response of mirror and reference
microphone were not measured simultaneously
but in two subsequent measurements.
Figure 5: Single channel measurement setup for acoustical
measurements with ARTA software
[13]
The resulting gain factor of the elliptical mirror is
shown in Figure 6 and the spatial resolution in Figure
7. The effective gain factor shows a good agreement
compared to the theoretical relation according to eq.
(1) derived by Grosche [12]. In the relevant frequency
range a gain of about 10 to 40dB is attainable.
Above frequencies of about 10kHz a decrease in the
measured gain factor of the mirror can be detected
which is caused by the non-omnidirectional directivity
of the loudspeaker. These effects have to be corrected

y measuring the directivity pattern of the loudspeaker.
The measured spatial resolution of the mirror widely
agrees with the linear trend of equation (2) but
shows slightly better values.
40
Reference Microphone
Acoustic Mirror Microphone
Theoretical Gain
Measured Gain
20
SPL [dB]
0
-20
Figure 8: Measurement setup
-40
-60
10 3 10 4
f [Hz]
Figure 9 shows a closer view of the described servo
controlled rig to measure static and dynamic pressures
in front and behind the propeller plane.
Figure 6: Effective gain factor of the elliptic mirror as
function of sound frequency
100
Theoretical values
Measured values
80
w 0
[mm]
60
40
20
20 40 60 80 100
λ [mm]
Figure 7: Width of diffraction pattern
Measurement setup
Figure 8 shows a typical measurement setup for acoustic
measurements in front of the wind tunnel. The
acoustic mirror is positioned in front of the propeller
with the far focus adjusted in the vicinity of a horizontally
levelled blade. The mirror can be moved parallel
to the propeller plane using a linear drive unit. In addition
a ground microphone is placed on a metal plate in
front of the propeller. To registrate possible atmospheric
turbulences a sonic anemometer is located besides
the propeller to measure local wind speeds.
Figure 9: Rig for measuring static and dynamic pressures
at the propeller plane

3 Results
In the following two characteristic operating conditions
of the reference propeller will be presented to illustrate
general effects of aerodynamic phenomenons
and noise radiation of the propeller. The measurements
were conducted outside the wind tunnel with
the measurement setup according to Figure 8. The
propeller was operating under static thrust conditions
(u ∞ = 0). Some preliminary results from the aeroacoustic
measurements are given in noise plots where
the sound pressure level is shown dependent of the
sound frequency and the dimensionless radial position
on the blade focused by the mirror. Positive values of
the radial position mean a position where the blade is
moving downward and negative values a position where
the blade is moving upward. In the analysis of the
acoustic data the doppler effect caused by the rotation
of the blades and the inclined adjustment of the mirror
is not yet considered. This effect leads to a shift
of the fequency to lower values for the downward moving
blade and a shift to larger values for the upward
moving blade while sound pressure levels remain unchanged.
The increment of the acoustic mirror while
scanning the propeller blades was set to 50mm which
corresponds to a dimensionless propeller radius increment
of 0.1. The sample rate of the mirror microphone
was 44kHz. The acoustic data was postprocessed
with a FFT analysis with a bandwidth of 3Hz. The
sound pressure levels of the acoustic mirror represent
the measured values at the mirror microphone. If the
sound pressure level in the vicinity of the blade surface
is of interest, these values have to be corrected according
to the distance law under consideration of the
focal length of the mirror according to Figure 3 and
4. The ground microphome was covered with a wind
shield which has not been calibrated yet. All in all the
following noise pressure values are intended to illustrate
general effects between two characteristic test cases
and not absolute values at this stage of the project.
3.1 Operating point with positive
propeller thrust
The first testcase demonstrates an operating point
when the propeller is generating positive thrust. According
to the CFD calculations and paint coatings
there are no separations on the blades. Figure 10 illustrates
the flow field calculated by TAU in the vicinity
of the propeller under static conditions. The resulting
vortex about 4 propeller radii downstream and the resuction
of air lead to a divergent wake.
Figure 11 shows a three-dimensional plot of the noise
spectrum for this condition. The locations of dominant
noise sources can be clearly seen as peaks. Figure 12
shows the corresponding projection of this plot. Three
main locations of higher noise levels are apparent. As
the focusing effect in the direction of the mirror axis
decreases slowly the noise radiation of the servo motor
can still be clearly seen in the hub region. A second
region with higher noise levels is located between 50
to 80% of the propeller blade. This corresponds to the
region where propeller circulation is highest and most
of the thrust is generated for normal operating conditions.
A third peak is found at the blade tips.
Figure 10: Flow field about reference propeller generating
positive thrust, u ∞ = 0
SPL [dB]
70
60
50
40
30
20
-1 -0.5 0 0.5 1
r/R [-]
0
5000
10000
15000
20000
f [Hz]
SPL [dB]
65
60
55
50
45
40
35
30
25
Figure 11: Sound pressure level 1/6-octave, positive
thrust
It’s striking that the noise peaks are not exactly symmetric
to the location of the hub. The symmetry plane
is slightly shifted to the side of the upward moving blade.
Although the acoustic tests were done under static
conditions bearing Figure 3 and 8 in mind propeller
induced flow velocities in the inside of the acoustic

Figure 12: Sound pressure level 1/6-octave, positive
thrust
Figure 14: Flow field of reference propeller generating
reverse thrust, u ∞ = 0
SPL [dB]
70
60
50
40
30
20
for this operating point and the airfoils do not deliver
high negative lift coefficients. As the propeller was originally
designed for forward flight conditions the lift
coefficient curve is inflated at the inboard sections of
the propeller blade. Changing the blade angle to lower
pitch angles causes a nearly vertical shift of the lift
coefficient curve. Under reverse thrust conditions this
fact is responsible that at the inboard sections of the
blade still positive lift coefficients prevail. According to
Figure 14 this effect causes a recirculation area at the
inboard sections of the blade in front of the propeller.
10
0
0 1 2 3 4 5
f [kHz]
Figure 13: Noise spectrum of ground microphone, positive
thrust
mirror can be responsible for an apparent shift of the
noise sources.
Figure 13 shows the noise spectrum of the ground microphone
located in front of the propeller. Tonal noise
components are dominating as far as about 5kHz
and the harmonics of the blade passing fequency of
162.5Hz can be clearly seen.
3.2 Operating point with reverse propeller
thrust
Manoeuvering on the ground the Zeppelin NT also
uses the variable pitch stern thruster to generate reverse
thrust although the used propeller was not designed
Figure 15: Sound pressure level 1/6-octave, reverse
thrust
Compared to the previous operation case the sound
pressure levels for reverse thrust are much higher (Fi-

gure 15) with the maxima shifted to the tip region.
This effect is caused by separations on the lower side
of the propeller blade in this region. Again the maxima
are not symmetric to the hub region. For this operating
point the symmetry axis is shifted to the side of
the downward moving blade.
The noise spectrum of the ground microphone supports
the assumption of separations in the tip region
as broadband noise components dominate and tonal
noise components are restricted to the first two harmonics
of the blade passing frequency.
SPL [dB]
100
80
60
40
20
f 0
f 1
0 1 2 3 4 5
f [kHz]
Figure 16: Noise spectrum of ground microphone, reverse
thrust
4 Conclusion and Outlook
The development of a process chain for highly automated
propeller design along with first experimental
results conducted with a propeller test bench were presented.
The aeroacoustic experiments show significant
separations in the tip region of the propeller blade if
the propeller operates in reverse mode. One reason for
the current asymmetry in regard to the hub in the
acoustic plots are caused by propeller induced flow
velocities in the acoustic mirror. This effect will be
pretty hard to account for as the flow field is strongly
dependent on the operating point of the propeller and
the flow field of the wind tunnel. From the beginning
the use of the acoustic mirror was intended to show
relative differences between different blade angles or
propeller designs rather than absolute noise levels. Nonetheless
this effect will be studied furthermore with
the intention to find appropriate corrections. Another
aspect in regard to the asymmetry between downward
and upward moving blade is caused by the directivity
of the trailing edge noise and convective amplification
which has been observed in acoustic array measurements
on a wind turbine as well [14]. These effects have
to be considered in the further analysis of the acoustic
data. In additon, further tests have to be done concerning
the repeatability of the noise measurements
especially under static thrust condition when the flow
about the blade tips is quite sensitive to external disturbances.
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