An Experimental Study of the Nose Landing Gear Noise

with Emphasis on the Steering Actuators, Torque Link and

Tow Hook

by

Ozan Coskun

A thesis submitted in conformity with the requirements

for the degree of Master of Applied Science

Graduate Department of Institute for Aerospace Studies

University of Toronto

© Copyright 2019 by Ozan Coskun

ii

Abstract

An Experimental Study of the Nose Landing Gear Noise with Emphasis on the Steering

Actuators, Torque Link and Tow Hook

Ozan Coskun

Master of Applied Science

Graduate Department of Institute for Aerospace Studies

University of Toronto

2019

Aeroacoustic and Particle Image Velocimetry (PIV) tests were conducted on the simplified model

of a nose landing gear that permitted independent additions of the tow hook, torque link and

steering systems. Tow hook was found to suppress low-frequency overhead noise originating from

the axle-strut region. Tests involving torque links mounted upstream and downstream of the strut

with varying arm separation distances showed that upstream mounting induces: i) prominent low-

frequency noise in the sideline direction, and ii) higher broadband noise in the overhead and

sideline directions when the torque link arms are close. Increasing the separation between the

torque link arms decreased the broadband noise contribution of the upstream mounted torque links

to levels of those mounted downstream. The steering actuator orientation is shown to be crucial

for the acoustic response of the gear. Steering actuators placed nearly perpendicular to the

freestream flow were found to generate the quietest responses.

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Acknowledgments

I wish to thank my supervisor Professor Alis Ekmekci for her leadership and support. I am

thankful to her for believing in me, in my research and mentoring me with her knowledge.

I would like to express my deepest appreciation to Philip McCarthy for his selfless dedication to

teach and support me whenever I needed. His extensive knowledge and experience in the field

was vital to the timely completion of my research. I cannot thank him enough.

I would like to extend my gratitude to Ehsan Vaziri, who spent hours on teaching me how to

conduct Particle Image Velocimetry experiments and staying with me while operating the laser. I

also wish to thank Alec Xu for providing me with laughter and an extra set of hands during some

of my experiments. I would also like to thank Bronwyn for her emotional support in this journey.

Last, but certainly not the least, I would like to thank my dad for being my inspiration to pursue

aerospace engineering, my mom for being the inspiration to pursue research, and my sister for

being the fourth member of our family.

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Table of Contents

1. Introduction and Literature Review ........................................................................... 1

1.1. Motivation and Background ..................................................................................... 1

1.2. Objectives ................................................................................................................. 3

2. Experimental Setup ...................................................................................................... 4

2.1. Test Facility .............................................................................................................. 4

2.2. Test Model ................................................................................................................ 5

2.3. Instrumentation, Data Acquisition and Data Analysis ............................................. 8

2.3.1. Acoustic Measurements ................................................................................. 8

2.3.2. Flow Visualization Measurements ................................................................. 11

3. Aeroacoustic Analyis and Discussion .......................................................................... 13

3.1. Empty Anechoic Wind Tunnel ................................................................................. 13

3.2. Baseline Nose Landing Gear .................................................................................... 14

3.3. Tow Hook ................................................................................................................. 18

3.4. Torque Link .............................................................................................................. 19

3.5. Steering System ........................................................................................................ 29

3.5.1. Flat-Tipped Steering Actuators ...................................................................... 30

3.5.2. Round-Tipped Steering Actuators ................................................................. 36

3.6. Fully Dressed Landing Gear with Downstream Torque Link .................................. 41

3.7. Mach Scaling ............................................................................................................ 45

4. Flow Visualization Results and Discussion ................................................................ 48

4.1. Flat-Tipped Steering Actuators ................................................................................ 48

4.2. Round-Tipped Steering Actuators ............................................................................ 52

4.3. Effect of the Flow Field on Acoustics ...................................................................... 55

5. Conclusion and Future Work ...................................................................................... 57

5.1. Conclusion ................................................................................................................ 57

5.2. Future Work.............................................................................................................. 59

Bibliography ......................................................................................................................... 61

Appendix A ........................................................................................................................... 64

A.1 Reynolds Number ..................................................................................................... 64

A.2 Spectral Analyses ...................................................................................................... 65

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List of Figures

1.1 Commercial aircraft noise over years. Image taken from [2]. ..................................... 1

2.1 Panoramic view of the anechoic wind tunnel with the model of a generic nose

landing gear installed in the test section. Flow is from left to right. ............................ 4

2.2 Components of the fully-dressed landing gear model as seen from upstream and

downstream. ................................................................................................................. 6

2.3 Sketches of the inter-wheel region showing the axle, strut, torque link and tow

hook components attached to one of the wheels. The other wheel is removed for

demonstration purposes in the drawings. ..................................................................... 6

2.4 Drawings of the landing gear components and the assembly. All dimensions are

given in mm. ................................................................................................................. 7

2.5 Cartesian coordinate system used in this study is shown on the landing gear model

mounted in the anechoic wind tunnel. .......................................................................... 8

2.6 The layout of the microphones in the linear arrays. ..................................................... 8

2.7 Microphone locations of the phased array system. ...................................................... 9

2.8 View facing the wind tunnel inlet with the landing gear and pitot-static system. ....... 10

2.9 Microphone attached to the pistonphone for sensitivity calibration. ........................... 10

2.10 Particle Image Velocimetry Setup................................................................................ 12

3.1 Narrowband noise level spectra of the AWT at U∞ = 60 m/s with the empty testbed

(background noise), baseline landing gear, and baseline landing gear with the

background noise deleted. ............................................................................................ 14

3.2 View of the baseline noise landing gear from downstream. ........................................ 15

3.3 Sound pressure levels (SPLs) as a function of frequency for the baseline landing

gear configuration from the (a) overhead (θ = 0°) and (b) sideline (φ = -90°)

microphone positions. Black arrows indicate frequencies at which the

beamforming maps are generated. ............................................................................... 15

3.4 Noise source maps for the baseline nose landing gear at low to high frequencies.

Flow is from left to right. ............................................................................................. 17

3.5 Comparison of the narrowband noise spectra between landing gear configurations

with and without tow hook for the (a) overhead (θ = 0°) and (b) sideline (φ = -90°)

microphone positions. .................................................................................................. 19

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3.6 The definition of the distance L, which is measured from the bottom of the axle to

the bottom of the oleo................................................................................................... 19

3.7 Overhead and sideline spectra for landing gear models with the torque link

installed downstream of the strut, the torque link installed upstream of the strut

and baseline gear. Arrows indicate frequencies at which beamforming maps are

generated. ..................................................................................................................... 21

3.8 Low-frequency beamforming maps of the landing gear equipped with (a) a torque

link downstream of the strut and (b) a torque link upstream of the strut. The

separation in torque link arms is L = 0.9D. Flow is from left to right. ........................ 24

3.9 Mid- and high-frequency beamforming maps of the landing gear with downstream

and upstream torque links with an arm separation length of L = 0.9D. Flow is from

left to right. ................................................................................................................... 25

3.10 Beamforming map at 6 kHz for the landing gear equipped with the upstream

torque link that has a torque link arm separation length of L = 1.1D. Flow is from

left to right. ................................................................................................................... 26

3.11 Flyover and Sideline directivities of landing gear with and without torque links. ...... 28

3.12 The OASPL directivity map in the flyover direction and the SPL spectra of the

landing gear between θ = -30° to +30° positions with the upstream torque link

having L = 0.9D (shown in green) and without the torque link (shown in black). ...... 28

3.13 Schematic defining the angle α between the axis of the steering actuator and the

freestream direction. Here, the flow direction is from top to bottom........................... 29

3.14 Frontal and isometric views of the landing gear with steering actuators installed at

α = 60°. ......................................................................................................................... 30

3.15 Overhead noise spectra of the landing gear with and without steering systems.

Black arrows indicate frequencies at which the beamforming maps are generated. ............. 31

3.16 Sideline noise spectra of the landing gear with and without the steering systems.

Black arrows indicate frequencies at which the beamforming maps are generated. ............. 32

3.17 Sideline noise directivities of landing gear with and without steering systems at

two selected frequency ranges...................................................................................... 33

3.18 Low frequency beamforming maps of the landing gear installed with steering

actuator having α = 40°. Flow is from left to right. ..................................................... 34

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3.19 Mid-frequency beamforming maps of the landing gear installed with steering

actuator having α = 40°. Flow is from left to right. ..................................................... 35

3.20 High frequency beamforming map of the landing gear installed with steering

actuator having α = 40°. Flow is from left to right. ..................................................... 35

3.21 Steering actuators equipped with rounded caps viewed from downstream. ................ 36

3.22 Overhead noise level spectra of landing gear equipped with rounded and flat

steering actuators compared to the spectrum of the baseline gear. .............................. 37

3.23 Sideline noise level spectra of landing gear equipped with rounded and flat

steering actuators compared to the spectrum of the baseline gear. .............................. 39

3.24 Overhead calculated (black) and measured (red) noise level spectra of the fully

dressed landing gear. .................................................................................................... 43

3.25 Sideline calculated (black) and measured (red) noise level spectra of the fully

dressed landing gear. .................................................................................................... 44

3.26 Beamforming maps of landing gear equipped with steering system, torque link,

and fully dressed landing gear. ..................................................................................... 45

3.27 Overhead spectra of the fully dressed landing gear model with steering actuators

at α = 10° and a torque link separation of L = 0.9D scaled up to the various powers

of the Mach number with a reference freestream velocity of U∞ = 68 m/s. ................. 46

3.28 Sideline spectra of the fully dressed lading gear model with steering actuators at

α = 10° and a torque link separation of L = 0.9D scaled up to the various powers

of the Mach number with a reference freestream velocity of U∞ = 68 m/s. ................. 47

4.1 Field of View of the PIV setup..................................................................................... 48

4.2 Instantaneous skipped velocity field superimposed on the instantaneous vorticity

contour for the flat-tipped steering actuator. Flow from left to right. .......................... 50

4.3 Time averaged skipped velocity field superimposed on the time averaged vorticity

contour for the flat-tipped steering actuator. Flow from left to right. .......................... 51

4.4 Instantaneous skipped velocity field superimposed on the instantaneous vorticity

contour for the round-tipped steering actuator. Flow from left to right. ...................... 53

4.5 Time averaged skipped velocity field superimposed on the time averaged vorticity

contour for the round-tipped steering actuator. Flow from left to right. ...................... 54

4.6 OASPLs of the gear with steering actuators in the 200 – 20,000 Hz range ................. 55

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Nomenclature D = Wheel diameter

f = Frequency

L = The separation of the torque link arms measured by the distance from the bottom of

the axle to the bottom of the oleo.

M = Mach number

Re = Reynolds number

x = Coordinate in streamwise direction (positive points downstream)

y = Coordinate in transverse direction in line with the centerline of the axle (positive

defined according to the right hand rule)

z = Coordinate in spanwise direction, in line with the strut centerline (positive points away

from the flat plate)

α = Angle of incidence for the steering actuator angle (with respect to the freestream flow)

Δt = user-defined time interval between two consecutive images

θ = Polar angle of the flyover microphones (with respect to the positive z axis)

φ = Azimuthal angle of the sideline microphones (with respect to the positive x axis)

1

Chapter 1

Introduction and Literature Review

1.1 Motivation and Background

The accelerating worldwide expansion in air transportation over the years has come with noise

pollution as a major negative side effect. A series of increasingly stringent noise standards

introduced by aviation authorities around the globe has ultimately provided a huge impetus for

research efforts that aimed to mitigate aircraft-induced noise. Over the past several decades,

efforts in aircraft noise reduction have primarily focused on minimizing engine noise because

originally it dwarfed all other aircraft noise sources. With the development of high-bypass-ratio

turbofan engines and improved acoustic liners in the nacelle, the engine noise has been reduced

so much on modern airliners (as shown in Figure 1.1) that today the non-propulsive components

constitute the majority of the total aircraft noise emission during approach and landing. Landing

gear is one of the largest contributors to this noise [1].

Figure 1.1 Commercial aircraft noise over years. Image taken from [2].

2

Landing gear noise is essentially broadband in nature [3]. Additional prominent tones are

occasionally observed due to the excitation of cavity modes in the landing gear wheel wells [4, 5]

and the hollow pins that connect different subcomponents of the gear structure [3]. When detected,

the pin-hole-related tonal noise can easily be eliminated by adopting pin-hole caps [6]. Turbulent

flow separating from the bluff components of landing gears and the interaction of the resulting

turbulent wakes with the downstream gear components are the two mechanisms generally

attributed to broadband noise production of landing gears [3]. Current generation landing gears

have very complex designs, consisting of a multitude of small- and large-size components. The

local turbulent and mean flow characteristics differ for each one of these gear elements due to the

occurrence of complex flow interactions between them. Moreover, the noise radiation directivities

vary for each component due to their dissimilar shape and orientation. As a result, the overall

noise characteristics for landing gears are very complex, spanning a wide range of frequencies

including the frequencies for which human hearing is most sensitive [3]. Generally, small landing

gear components are found to be responsible for high-frequency broadband noise levels, whereas

bulky gear components, such as tires and large struts are related to low-frequency noise [3].

Computational methods often cannot predict the acoustic response accurately, especially the

high-frequency response that is caused by small-scale parts on the landing gear. On the contrary,

the lower frequency acoustic response resulting from the large-scale vortex shedding behind the

wheels, strut, oleo and torque links as well as the tonal response from bay cavities can be predicted

computationally, however, experimental validations are still needed. This need for experimental

validation has spurred a number of high-profile test campaigns around the world involving

independent collaborations of NASA, DLR and NRC with aircraft companies Gulfstream, Airbus

and Bombardier respectively as well as several universities [7-12]. Vast majority of the existing

studies often omitted several geometric details in their test model, missing many of the

complicated interactions present within real-life landing gears. However, a component-by-

component assessment of landing gear noise is necessary to understand the causes of broadband

and tonal noises, as well as the interaction noise between landing gear components.

Noise source localization maps obtained by Ref. [11] via beamforming on different models of

nose landing gear comprised of a torque link placed downstream of the main strut highlighted the

torque arm as a dominant noise contributor. Other studies, such as the computational results of

Ref. [13] and the experimental results of Ref. [14], supported this observation. This is believed to

be due to the interaction effects between the torque link and the strut assembly [11]. Interestingly,

3

Long et al. showed that having the torque link assembly downstream of the main strut is less noisy

than having it at the upstream [15]. Windiate et al. also showed that if the angle between the two

torque link arms reaches 130°, then the torque link and the strut act as one bluff body, and suppress

the vortex shedding from the torque link-strut pair [16]. Bennett et al. conducted a component

removal assessment on a full-scale nose landing gear, where the influence of removing one

component at a time on noise output was characterized [17]. Although in that test campaign, only

the steering actuators installed perpendicular to the freestream flow direction were considered,

results suggested that the landing gear wheels suppress vortex shedding from steering actuators

perpendicular to the flow.

1.2 Objectives

Redesigning the shape and/or position of a gear component can change the local flow

conditions for adjacent landing gear components, and thereby, alter their noise contributions. This

makes the development of noise reduction concepts associated with landing gears a remarkably

challenging problem. It is not uncommon to see low-noise concepts developed with a purely

heuristic intuition to lead to even much noisier designs. Development of effective noise reduction

strategies requires a thorough understanding of the underlying noise generation mechanisms and

how these are affected by the design variations. To this end, this thesis aims to identify the far-

field noise contributions of landing gear components, focusing specifically on torque links, tow

hooks and steering system components and their variations. The model used is a simplified version

of a generic nose landing gear, but has a modular design that permits independent addition,

subtraction as well as variation of the landing gear components for full appreciation of their

acoustic impacts.

4

Chapter 2

Experimental Setup

2.1 Test Facility

Acoustic investigations were executed in the Anechoic Wind Tunnel (AWT) located at the

University of Toronto Institute for Aerospace Studies (UTIAS). This is an open-jet, open-loop

wind tunnel, capable of delivering a range of flow speeds (up to 75 m/s) through its square nozzle

of 0.60 m × 0.60 m. The speeds achievable in this facility fall within the range of real commercial

aircraft approach speeds. The jet exits the test section through a circular flow collector located 2.4

m downstream of the square inlet nozzle. The open jet is enclosed within a chamber that measures

6 m long, 6 m wide and 3 m high. Its walls are covered by polyurethane foam wedges (0.55 m

deep), making the chamber anechoic down to a lower frequency limit of 170 Hz. A panoramic

photo of the wind tunnel provided in Figure 2.1 depicts the nozzle inlet, the foam-treated anechoic

chamber and the circular flow collector along with the landing gear test model installed in the test

section. Details about the full characterization of the aerodynamic and aeroacoustics properties

for this wind tunnel can be found in Ref. [18].

Figure 2.1 Panoramic view of the anechoic wind tunnel with the model of a nose landing

gear configuration installed in the test section. Flow is from left to right.

5

2.2 Test Model

For the present test campaign, a 30%-scale, simplified model of a generic nose landing gear

was constructed and attached upside down on a rigid flat plate, which can be considered to

simulate the aircraft fuselage. The fully-dressed model consisted of two wheels, an axle

connecting the wheels, a main strut, steering actuators with brackets, an oleo, a torque link and a

tow hook, as shown in Figure 2.2. The landing gear model had a modular design that permitted

its steering actuators, torque link and tow hook components to be independently added, taken out

or varied. This helped attain information about the contribution of these three components to noise

and allowed acoustic evaluations to be made for various configurations of the model.

The landing gear wheels of D = 0.152 m diameter were constructed from acetal, with no interior

or exterior cavities. They had a smooth surface with three tire treads on their circumference. The

strut was made of 0.024 m-diameter steel electrical conduit tube that was cut to length and sanded

smooth. The axle, oleo, tow hook and torque link arms were all 3D printed using polylactic acid

and were sanded smooth to remove printing surface imperfections. The steering actuators were

machined from aluminum and attached to the steering brackets with 3D printed lock pins.

Brackets holding the steering actuators were constructed using laser-cut acrylic sheets that were

bonded to machined acrylic steering columns.

Two types of axle were constructed, which enabled the torque links to be positioned either

downstream or upstream of the strut as shown in Figure 2.3. Both axles enabled an offset of 23

mm between the centerlines of the strut and the wheel, with the centreline of the strut positioned

upstream of the centerline of the wheel. A total of 10 steering brackets were constructed, enabling

the angle of incidence of the steering actuators to be varied in 10° increments from 𝛼 = 0° to 90°,

with respect to the freestream direction, where 0° defines a steering actuator aligned in the

freestream direction with its free tip pointing upstream. The dimensions of the steering brackets

are shown in Figure 2.4 (a). The steering bracket dimensions for each 𝛼 configuration were varied

by defining the center of rotation as the pinhole “A” as shown in Figure 2.4 (a), and sweeping the

pinhole “B” in a constant radius of 42.7 mm around A. Two variations were considered to the free

tip of the steering actuators: flat and rounded. The latter was achieved through 3D printed exact

half spheres. For the tow hook, which is located upstream of the axle when present, landing gear

configurations with and without this component were tested. Dimensions of the tow hook can be

seen in Figure 2.4 (b). The inclination of the torque link arms was varied by changing the distance

6

between the two ends of the fixed-length torque link arms and their dimensions of the torque link

arms are shown in Figure 2.4 (c). The dimensional characteristics of the baseline landing gear

model are shown in Figure 2.4 (d) for one configuration of the landing gear model.

(a) View from upstream (b) View from downstream

Figure 2.2 Components of the fully-dressed landing gear model as seen from upstream and

downstream.

(a) configuration with the torque link

downstream of the strut.

Figure 2.3 Sketches of the inter-wheel region showing the axle, strut, torque link and tow

hook components attached to one of the wheels. The other wheel is removed for

demonstration purposes in the drawings.

(b) configuration with the torque link

upstream of the strut.

7

(a) Steering bracket

(c) Torque link arms

(i) lower arm

(ii) upper arm

Figure 2.4 Drawings of the landing gear components and the assembly. All dimensions are

given in mm.

The photograph provided in Figure 2.5 depicts one configuration of the landing gear model

installed on the flat plate in the anechoic wind tunnel, along with the Cartesian coordinate system

used. The origin of this coordinate system is located at the half-distance between the wheels,

(b) Tow Hook

(d) Baseline landing gear

8

cutting the centerline of the strut, and sits on the same horizontal plane as the centerline of the

axle. The positive x direction points downstream, the y coordinate is the transverse direction

perpendicular to the centerline of the strut, with its positive direction defined according to the

right hand rule, and the z coordinate is in line with the centerline of the strut, positive pointing

away from the flat plate.

2.3 Instrumentation, Data Acquisition and Analysis

2.3.1 Acoustic Measurements

In the present study, far-field noise levels and the directivity information were quantified

through linear microphone array measurements, and locations of significant noise sources on

different configurations of the landing gear model were identified using the data acquired through

a phased microphone array.

For linear microphone array measurements, two setups were employed, one in the flyover

direction and the other in the sideline direction with respect to the model. Seven PCB 130F20

microphones were placed along a straight line in 10° increments on each of the flyover and

sideline arrays. The flyover microphone array setup was in the y-z plane and involved

microphones spanning over the range of polar angles from 𝜃 = -30° to +30° (with respect to the

positive z axis). The sideline setup was positioned in the x-y plane and had microphones positioned

over an azimuthal angular range from 𝜑 = -60° to -120° (with respect to the positive x axis). Figure

Figure 2.5 Cartesian coordinate system

used in this study is shown on the

landing gear model mounted in the

anechoic wind tunnel.

Figure 2.6 Top view of the layout of

the microphones in the linear

arrays.

9

2.6 illustrates the location of the linear array microphones used in the flyover and sideline

directions relative to the coordinate system. The closest microphone to the model in both arrays

was at approximately 1 m distance from the origin of the coordinate system. The exact distance

of each microphone was measured with a laser distance measurer with an accuracy of ±1 mm and

the sound pressure level amplitudes for each microphone were corrected to an equivalent radial

distance of 1 m from the origin.

The phased microphone array consisted of a total of forty-five PCB 130F20 microphones,

mounted on a fibre board with approximate dimensions of 1 m × 1 m × 0.013m. The sensing

element of each microphone was placed 9 mm in front of the surface of the board, and the array

board was placed at the side of the landing gear in the x-z plane, offset 1.5 m from the coordinate

origin. The array microphones were distributed along 5 logarithmic spirals based on the “equal

area aperture” layout developed by Ref. [19] due to its favorable performance at reducing

sidelobes [20]. In this design, the array area is divided into equal area annuli, which are further

subdivided into equal area regions, and the microphones are placed at the centre of these regions.

Figure 2.7 shows the microphone locations in the phased array design used in this investigation.

Figure 2.7 Microphone locations of the phased array system.

Prior to data acquisition, once on each day of experiments, the wind tunnel speed was calibrated

with a pitot-static system mounted upstream of the landing gear model, which can be seen in

Figure 2.8. The wind tunnel voltage input was varied between 0 to 10 V and the corresponding

linear relationship between voltage and freestream velocity was acquired. Following this, each

10

microphone in the arrays was calibrated with a GRAS 42AA pistonphone as seen in Figure 2.9,

and their sensitivities were recorded.

All experiments discussed in this thesis with the exception of Mach scaling tests discussed in

Section 3.7 were conducted at the freestream velocity of U∞ = 60 m/s, which corresponded to the

Mach number of M = 0.175 and the Reynolds number of Re = 600,000 based on the wheel

diameter (D = 0.15 m). Mach scaling tests were conducted at freestream velocities between 54 –

68 m/s, with 2 m/s increments, which corresponded to the Mach numbers of M = 0.157 – 0.198

and Reynolds numbers of Re = 560,000 – 700,000 based on the wheel diameter. The microphone

signals were recorded using National Instrument PXI-4497 data acquisition (DAQ) cards at a rate

of 65,536 (216) Hz for 20 s for the linear array and at a rate of 32,768 (215) Hz for 64 s for the

phased array experiments. Through the use of a National Instruments PXIe-1082 chassis, the data

acquisition system was capable of simultaneous acquisition of all the microphones in the linear

array or the phased array systems. The pressure data, which is in time domain in its raw form was

then converted into the frequency domain for the noise level spectra to be displayed. The sound

pressure level (SPL) and the overall sound pressure level (OASPL), which is the integration of

the SPL over a defined frequency range, were generated from the acquired data. In this study, the

SPLs are presented in the form of narrow-band spectra with a frequency resolution of 4 Hz. As

the wind tunnel is no longer anechoic below 170 Hz, when calculating OASPLs, the frequency

Figure 2.8: View facing the wind tunnel inlet

with the landing gear and pitot-static system.

Figure 2.9: Microphone attached to the

pistonphone for sensitivity calibration.

11

integration band limits were set to 200 Hz to 20,000 Hz. Phased microphone array data were

analyzed using the functional beamforming algorithm with an order of 50 [21]. In the presented

phased array data, the microphone locations were corrected for the refraction effect of the shear

layer of the wind tunnel open jet, following the method proposed by Ref. [22]. Shear layer

corrections were not applied to the linear array data, hence, the sideline angles, 𝜑, presented in

this paper refer to the geometric positions of the microphones. Uncertainty regarding the

instrumentation and data acquisition are shown in Appendix A.

2.3.2 Flow Visualization Measurements

Particle Image Velocimetry (PIV) is a quantitative, optical flow visualization technique used

to obtain instantaneous velocity of the plane of interest in a flow field. It is executed by seeding

the flow with tracer particles that follow the flow without altering it and illuminating the plane of

interest with a laser sheet. Images of the plane of interest are captured with a camera in pairs, with

each image in the pair separated by a user-defined time interval known as Δt. A PIV processing

software then can use the two consecutive images to obtain the instantaneous velocity vector field

by calculating the distance in which the tracer particles have moved between the two frames.

Using the velocity vector field extracted from the image pairs, vorticity, turbulent kinetic energy,

and other flow characteristics can be computed.

In the present investigation, a double-pulse Nd:YAG laser was used as the source of particle

illumination on the horizontal plane, cutting through the midplane of the steering actuators as seen

in Figure 2.10 (a). The plate that the landing gear was attached had an optically transparent section

made out of acrylic which enabled the bottom-mounted camera to capture images of the exterior

of the steering actuators. A photo of this setup can be seen in Figure 2.10 (b). The camera used in

these experiments was a TSI PowerView 2MP Plus. The dimensions of the field of view (FOV)

were 98 × 72 mm with a resolution of 16 pixels/mm. A total of 1,000 image pairs were captured

during each experimental run with a data capture rate of 14.5 Hz, which corresponds to the

maximum operating rate of the laser source. The optimal Δt was determined to be 6 μs for U∞ =

60 m/s through trial and error. The synchronization of the camera shutter and the laser pulse was

ensured via the use of a TSI LaserPulse 610035 synchronizer. An oil-based ViCount Compact

1300 smoke generator was used as the source for the tracer particles. The smoke was pumped

with a centrifugal pump to the inlet of the anechoic wind tunnel. The height of the laser sheet was

adjusted via placing a burn paper at the height corresponding to the flow region of interest and

12

moving the laser vertically until the paper was burnt. Data were recorded and the initial vector

fields were calculated by INSIGHT 4GTM, supplied by TSI Inc. Vector fields were later analyzed

on an in-house software to ensure the integrity of the vectors produced. Insufficient tracer particles

and highly reflective surfaces can result in erroneous vectors in their surrounding areas. By

defining lower and upper limits to the components of velocity, the erroneous vectors were

removed and the vectors for empty grids were calculated by linearly interpolating vectors from

surrounding grids.

(a) Drawing of the setup (b) Photo of the setup

Figure 2.10 Particle Image Velocimetry Setup

13

Chapter 3

Aeroacoustic Analysis and Discussion

This chapter compares far-field acoustic results of different landing gear model configurations,

which involve individual addition and/or variation of the torque link, tow hook and steering

actuator components. Far-field noise spectra, noise radiation directivity and noise source maps

are discussed. Far-field noise spectra are assessed at two microphone positions: directly overhead

of the model (at a flyover angle of θ = 0°) and to the side of the model (at a sideline angle of φ =

-90°). Based on these results, the noise source maps are computed for frequencies where important

spectral features are observed in the noise spectra. The OASPL mapping of the noise directivity

is constructed from noise records of both the flyover and sideline linear microphone arrays.

3.1 Empty Anechoic Wind Tunnel

Overhead and sideline microphone measurements were conducted in the AWT with the test

fixtures, cables and the empty test bed at freestream velocity of U∞ = 60 m/s and the noise level

spectra of this configuration can be found in Figure 3.1. The objective was to assess the

background noise levels of the AWT beyond its anechoic cut-off frequency of 170 Hz and to

determine the significance of the background noise to the test configurations discussed later in

this Chapter. In Figure 3.1, three spectra can be seen: AWT background noise, baseline landing

gear, and baseline landing gear with the AWT background noise deleted. Focusing on (a) for the

overhead spectra, omission of the background noise induces a 2 dB noise reduction at frequencies

above 600 Hz. However, the background noise becomes more prominent in the lower frequencies,

contributing up to 12 dB of noise at 200 Hz to the measured acoustic response of the landing gear.

Sideline contribution of the AWT background noise is similar, as seen in (b), however less intense.

Deleting the background noise decreases the measured noise levels by 2 dB at frequencies above

350 Hz, and up to 6 dB at lower frequencies. Despite the significant contribution of the AWT

background noise to the overall measured noise of the landing gear at low frequencies, the

14

background noise is not omitted in any of the data shared in Chapter 3 for two reasons. First, the

main point of interest is the relative noise levels and spectra of each landing gear configuration

and that deleting the background noise would not add any meaning to the comparison since it is

constant for all the configurations. Second, the AWT background noise may be significantly

altered in the presence of the landing gear models, as these two noise sources can be interacting

with one another. Hence, a linear subtraction of sound pressure levels would not yield accurate

results when discussing the landing gear acoustic spectra.

(a) Overhead (b) Sideline

Figure 3.1 Narrowband noise level spectra of the AWT at U∞ = 60 m/s with the empty testbed

(background noise), baseline landing gear, and baseline landing gear with the background

noise deleted.

3.2 Baseline Nose Landing Gear

The simplest landing gear model considered in this study consists of two wheels, an axle, a

main strut and an oleo and can be seen in Figure 3.2. This will be taken as the baseline

configuration against which the acoustic response of different model variations will be compared.

Exclusion of the torque link, tow hook and the steering system in this model helps deduce

information about some of the important contributors to the landing gear noise that may otherwise

be concealed or changed as a result of the noise shielding/reflecting effects or the new flow

interactions arising from the addition of those components.

The SPL spectra measured directly overhead and to the side of the baseline nose landing gear

model are shown in Figure 3.3 (a) and (b), respectively. In the overhead position, given in Figure

3.3 (a), the spectrum is largely dominated by noise below 1 kHz. In this range, two distinct regions

15

Figure 3.2 View of the baseline noise landing gear from downstream.

(a) Overhead (b) Sideline

Figure 3.3 Sound pressure levels (SPLs) as a function of frequency for the baseline landing

gear configuration from the (a) overhead (θ = 0°) and (b) sideline (φ = -90°) microphone

positions. Black arrows indicate frequencies at which the beamforming maps are generated.

of frequencies, where spectral peaks are detected, are of particular interest. These are 300-400 Hz

and 500-700 Hz. Both are too low in frequency to be observed by the present phased array setup

with adequate resolution for the identification of their origins. In the first range, between 300-400

Hz, a tonal excursion centred around 324 Hz is observed. The theoretically expected vortex

shedding frequencies for smooth cylinders with diameters equal to those of the strut and oleo are

527 Hz and 421 Hz, respectively. Although these values are only estimates calculated using the

results of Ref. [23], their difference from the 324 Hz peak suggests that the radiated noise in this

range is unlikely to be a result of the Karman vortex shedding from the strut or the oleo. This

hypothesis is further supported by separate experiments conducted between freestream velocities

16

of 54 – 68 m/s (results not shared here), where the spectral peak at around 324 Hz was observed

to be independent of both the variations in the freestream velocity as well as the absence or

presence of the oleo. Furthermore, in the sideline spectrum shown in Figure 3.3 (b), a frequency

hump that is broader but peaks at around the same frequency value is also present, which may be

related to the peak centred around the 324 Hz in the flyover direction. Although the origins of this

noise are not possible to isolate based on the present results, in the upcoming sections, further

discussions will be presented related to this tone based on how the SPL varies with the model

configuration variations. The second low frequency range where the noise spectrum of the

overhead microphone, given in Figure 3.3 (a), shows a further tonal peak is between 500-700 Hz.

In this range, the tonal peak is centered at around 648 Hz. The acoustic response in this frequency

excursion is highly directional as it is not present in the sideline spectrum given in Figure 3.3 (b).

Given that it is detected in the overhead direction but not in the sideline position, this peak may

potentially be generated in the inter-wheel region and radiated only in the flyover direction, while

being shielded by the wheels against sideline radiation. Although the source of this tone is not

conclusively apparent from the results presented so far and cannot be deduced from the phased

array data due to the limited resolution of the measurements at such low frequencies, the possible

origins of this noise source will also be discussed in the upcoming sections based on the relation

of this peak to the variations of the landing gear configuration. It is noted that the blade passing

frequency of the anechoic wind tunnel for the freestream velocity of 60 m/s, which is the speed

used in the results presented, is 210 Hz, and hence, unrelated to the two peaks discussed in

association with Figure 3.3.

In order to gain better insight into the observed notable features in SPL, sideline beamforming

maps are presented in Figure 3.4 at frequencies that correspond to the peaks and troughs of SPL

observed in the sideline spectrum (given in Figure 3.3 (b)) as well as at selected low frequencies

between 800 Hz to 1,000 Hz. Flow direction in those maps are from left to right. As indicated

earlier, at lower frequencies, particularly below 800 Hz, it was not possible to extract viable noise

localization results. Even at frequencies between 800 Hz to 1,000 Hz, the beamforming resolution

was low and thereby appeared distributed over a large region, making it difficult to pinpoint the

exact location of a noise source. However, these results are also shared to draw a contrast between

the relative locations of noise sources for the baseline landing gear at low and high frequencies.

For clarity, the frequencies for which beamforming maps are presented are marked with arrows

in Figure 3.3 (b).

17

(a) (b)

(c) (d)

(e) (f)

Figure 3.4 Noise source maps for the baseline nose landing gear at low to high frequencies.

Flow is from left to right.

18

The noise localization maps in Figure 3.4 (a) and (b) indicate that the dominant noise at 800

Hz and 904 Hz originate from a region covering the lower portion of the strut and the whole

wheel. At a frequency of 1,504 Hz, the dominant noise source migrates upward, towards the

downstream of the intersection of the strut and oleo, as seen in Figure 3.4 (c). At 3,000 Hz, the

beamforming map in Figure 3.4 (d) identifies the region covering the downstream of the strut,

oleo and upper half of the wheels as significant noise sources. For both the frequencies of 4,504

and 6,000 Hz, two distinct noise sources appear dominant in Figure 3.4 (e) and (f). These sources

are concentrated near the upper oleo region as well as the downstream of the wheels, with 4,504

Hz being prominent downstream of the upper half of the wheel and 6,000 Hz being concentrated

around the downstream of the near-central region of the wheel. It is important to note that the

noise shown to be propagating from around the wheels can possibly be generated in the inter-

wheel region where axle-strut and axle-wheel junctions are also present. However, since these

regions are shielded by the wheels, it is not possible to determine the exact location of the noise

source based on the phased array measurements made from a sideline position.

3.3 Tow Hook

Figure 3.5 compares the overhead and sideline SPL spectra of the landing gear model that has

the tow hook attachment against those of the baseline landing gear configuration (without tow

hook). Here, the landing gear model with tow hook is identical to the baseline configuration, with

the only difference being the tow hook attachment to the upstream of the axle. Initially focusing

on the overhead position, shown in Figure 3.5 (a), it can be seen that the overhead tonal peak that

was observed at around 324 Hz in the baseline configuration (in Figure 3.3 (a)) is not affected by

the addition of the tow hook upstream of the axle, suggesting that this noise does not originate

from the axle or the axle-strut junction. Furthermore, the sideline noise spectra, given in Figure

3.5 (b), are almost identical for the configurations with and without tow hook (the latter being the

baseline case). This is expected, since the majority of the tow hook is shielded by the wheels in

the sideline direction. Regarding the second spectral peak centred at around 648 Hz in the

overhead direction, Figure 3.5 (a) shows a decrease of up to 4 dB with the addition of the tow

hook upstream of the axle. Considering this, and the highly directional attributes of the peak at

648 Hz in the flyover direction for both configurations with and without the tow hook, it becomes

highly possible that this peak is generated in the axle and/or the axle-strut interaction region, and

is shielded by the wheels and directed only between the wheels in the overhead direction.

19

(a) Overhead (b) Sideline

Figure 3.5 Comparison of the narrowband noise spectra between landing gear

configurations with and without tow hook for the (a) overhead (θ = 0°) and (b) sideline (φ =

-90°) microphone positions.

3.4 Torque Link

A total of eight different torque link configurations were tested. Four of these configurations

were mounted downstream of the strut, with the distance between the bottom of the axle and the

bottom of the oleo (which is denoted as L in Figure 3.6) varying from 0.8D to 1.1D in increments

of 0.1D. Varying the L distance altered the inclination of the fixed-length torque link arms with

respect to the strut. The remaining four configurations involved the same L variations with torque

links mounted upstream of the strut.

Figure 3.6 The definition of the distance L, which is measured from the bottom of the axle

to the bottom of the oleo.

20

The acoustic effects of the different torque link configurations studied in this investigation can

be observed in Figure 3.7, where the SPL spectra measured directly overhead and to the side of

the gear model with varying torque link configurations are presented. Each plot in Figure 3.7

depicts the noise spectra of the landing gear model with torque links placed both upstream and

downstream of the strut with a given L distance, along with the SPL spectrum of the baseline gear

model (without a torque link attachment).

For frequencies below 1 kHz, the overhead spectra given in Figure 3.7 (a) for both the upstream

and downstream placements of the torque link with a separation distance of L = 0.8D show only

a small (up to 1.5 dB) decrease compared to the baseline case in the first low-frequency spectral

peak at around 324 Hz. Further increase in the torque link arm separation distance to L = 0.9D

and beyond induces an additional 2 to 3 dB of suppression of this overhead peak (in Figure 3.7

(c), (e) and (g)). The second low-frequency peak encountered in the overhead direction at 648 Hz

reduces by 4 to 5 dB with the addition of an upstream torque link for all L separations (as seen

from Figure 3.7 (a), (c), (e) and (g)). A similar reduction in this peak was also observed with the

addition of the tow hook in Figure 3.5 (a), further indicating that the components placed upstream

of the axle seem to suppress the related noise source. This aspect is further reinforced by the

observation that the effect of a torque link placed downstream of the strut is minimal on this

frequency in the flyover direction for the range of the L values considered (as seen from all the

overhead microphone results of Figure 3.7). Furthermore, the addition of a torque link

downstream of the strut causes a SPL increase of up to 4 dB in the overhead direction between

the two low-frequency peaks, broadening the acoustic response in low frequencies.

For the mid- and high-frequency ranges in Figure 3.7 (a), the torque link placed downstream

of the strut with a torque link arm separation of L = 0.8D induces up to 3 dB broadband increase

in the overhead direction while the same torque link placed upstream of the strut causes as much

as 8 dB of increase when compared to the SPL of the baseline landing gear. Increasing the arm

separation distance to L = 0.9D does not seem to have a major effect in these mid- and high-

frequency noise levels for the overhead direction, as seen in Figure 3.7 (c). However, with

increasing arm separation distance to L = 1.0D and 1.1D, Figure 3.7 (e) and (g) shows that the

mid- and high-frequency response of the configurations fitted with the upstream and downstream

torque links approach one another. Although the reason for this observation is not yet fully

understood, one possible scenario could be that as the torque link arm separation increases, the

21

(a) L = 0.8D - Overhead (b) L = 0.8D - Sideline

(c) L = 0.9D - Overhead (d) L = 0.9D - Sideline

(e) L = 1.0D - Overhead (f) L = 1.0D - Sideline

(g) L = 1.1D - Overhead (h) L = 1.1D - Sideline

Figure 3.7 Overhead and sideline spectra for landing gear models with the torque link

installed downstream of the strut, the torque link installed upstream of the strut and

baseline gear. Arrows indicate frequencies at which beamforming maps are generated.

22

torque link arms get closer to the strut, and as a result the torque link-strut pair may start acting

like a single bluff body over a larger length of the strut. Flow visualization experiments are

required to shed further light into these acoustic observations.

Focusing on the sideline direction, Figure 3.7 (b) shows that the noise spectrum corresponding

to the landing gear equipped with the upstream torque link, having a separation distance of L =

0.8D, features two prominent low-frequency peaks at 630 Hz and 820 Hz. The same torque link

was observed to suppress the overhead noise by 4 to 5 dB at around 630 Hz in Figure 3.7 (a).

Hence, while reducing the noise radiated in the overhead direction, the upstream torque link

increases the sideline noise levels at 630 Hz by up to 10 dB when compared to the baseline gear,

indicating that the direction of the radiated noise at this frequency is shifted from the overhead

direction to the sideline direction with the presence of the upstream torque link. Inspecting the

sideline spectra for all configurations fitted with an upstream torque link at varying torque link

separations (given in Figure 3.7 (b), (d), (f) and (h)) reveals that the SPL at the low-frequency

peak of 630 Hz stays relatively constant (at around 70 dB) for all torque link arm separation

distances. However, the peak seen at 820 Hz rises to 76 dB at a separation of L = 0.9D (in Figure

3.7 (d)), and decreases to 73 dB and 66 dB with torque link separations of L = 1.0D and L = 1.1D,

respectively (as shown in Figure 3.7 (f) and (h)). Hence, the dependence of this peak to the

upstream torque link separation suggests that the peak at 820 Hz may be a result of the interaction

between the upstream torque link and the strut.

Focusing the attention to the mid- and high- frequency range of the sideline spectra for a torque

link arm separation of L = 0.8D (shown in Figure 3.7 (b)), addition of the downstream torque link

induces an increase of up to 1 dB in the 2 to 3 kHz range and 3 dB in the 3 to 6 kHz range.

Increasing the separation to L = 0.9D does not seem to cause a significant change in this acoustic

character (as seen in Figure 3.7 (d)). Further increasing the separation to L = 1.0D and L = 1.1D

induces a small increase of 2 dB in SPL in the 1 to 2 kHz range, and up to 3 dB at frequencies

beyond 2 kHz when compared against the response of the baseline gear in Figure 3.4 (f) and (h).

Therefore, in general, the downstream torque link appears to increase the mid- to high-frequency

range only slightly. On the other hand, the addition of a torque link upstream of the strut with an

arm separation distance of L = 0.8D (in Figure 3.7 (b)) induces a much more significant increase

in noise. The broadband noise rises by 6 dB at frequencies of 3 kHz and higher. An additional

increase can be seen around 2.65 kHz that induces a 9 dB increase of noise when compared to the

23

response of the baseline landing gear at that same frequency. Furthermore, a broadband noise

level rise of 8 dB can be seen between 1.1 to 2 kHz. Increasing the torque link arm separation to

L = 0.9D (Figure 3.7 (d)) does not seem to cause any significant changes in the mid- and high-

frequency spectra for the gear equipped with the upstream torque link when compared to the L =

0.8D counterpart. However, with a further increase in arm separation, to L = 1.0D, the increases

in noise levels in the mid- and high- frequencies become lesser, and eventually for L = 1.1D, the

noise levels of the gear with the upstream and downstream torque links in the mid- and high-

frequencies collapse under a similar spectrum. A comparable trend was also observed earlier in

the overhead SPL in Figure 3.7 (g). As indicated earlier, this similarity in noise spectra for the

upstream and downstream placements of a torque link at large torque link arm separation distances

may be linked to the coupling of the torque link and strut as a single bluff body due to their close

proximity.

Beamforming results for landing gears with torque link arm separation distances of L = 0.8D,

0.9D, and 1.0D are found to depict similar locations for the noise sources, despite slight variations

in sound pressure levels. Therefore, the beamforming results that will be displayed in this section

for the landing gear with a torque link separation of L = 0.9D can be taken as representative of all

three L distances. Beamforming maps that will be discussed here are constructed at selected

frequencies that correspond to significant noise level contributions in the presence of the torque

links. Arrows in Figure 3.7 (d) mark these frequencies.

For the upstream and downstream placement of the torque link with L = 0.9D, Figure 3.8 shows

the beamforming maps at 824 Hz, which corresponds to the second prominent low-frequency

peak recorded by the sideline microphone in presence of the upstream torque link in Figure 3.7

(d). Due to the poor resolution of the beamforming array at low frequencies, it is not possible to

distinctively identify the individual noise sources within the landing gear, and therefore, Figure

3.8 (a) shows the source of this low frequency noise as one large region covering the entire wheel,

axle and the torque link for the landing gear with downstream torque link. In the case of the gear

with upstream torque link, Figure 3.8 (b) shows the upper torque link arm and the oleo intersection

region as the major contributor to noise at 824 Hz. In fact, the presence of such a sharp main lobe

at such a low frequency in Figure 3.8 (b) is in accord with the high level of the tonal response

seen at around 820 Hz in Figure 3.7 (d).

24

(a) Downstream torque link (b) Upstream torque link

Figure 3.8 Low-frequency beamforming maps of the landing gear equipped with (a) a torque

link downstream of the strut and (b) a torque link upstream of the strut. The separation in

torque link arms is L = 0.9D. Flow is from left to right.

Mid- and high-frequency beamforming maps for the same landing gear configurations with

downstream and upstream torque links (having L = 0.9 D) are shown in Figure 3.9. Based on

Figure 3.9 (a), the source of major noise generation at 2.6 kHz is the junction between the strut

and the upper arm of the torque link for the landing gear with the downstream torque link.

However, Figure 3.9 (b) shows that, for the gear with upstream torque link, the dominant noise at

2.6 kHz originates from an area between the lower arm of the torque link and the strut, and is 8

dB more intense when compared to the acoustic response of the gear with downstream torque

link, given in Figure 3.9 (a). The difference in the levels of acoustic response at this frequency is

in agreement with the spectra results shared in Figure 3.7 (d).

From Figure 3.9 (c) and (d), it can be seen that the dominant noise generation at 4.5 kHz occurs

from the upper arm of the torque links, including the connection between the upper arm and the

oleo regions, for both landing gear configurations (with a torque link in the downstream as well

as upstream of the strut). However, the acoustic response originating from the upper arm of the

upstream torque link is approximately 2 dB higher than that of the downstream torque link. The

major contributor to noise at 6 kHz for the gear with downstream torque link is observed to be the

connection region between the upper arm of the torque link and the oleo in Figure 3.9 (e). Notice

that for the case of the gear with downstream torque link, as the frequency increases, the dominant

25

Downstream Torque Link Upstream Torque Link

(a) (b)

(c) (d)

(e) (f)

Figure 3.9 Mid- and high-frequency beamforming maps of the landing gear with

downstream and upstream torque links with an arm separation length of L = 0.9D. Flow is

from left to right.

26

noise source associated with that frequency migrates upwards from the wheel region to the oleo.

This trend can be observed by inspecting Figure 3.8 (a), Figure 3.9 (a), (c) and (f). For the case

with the upstream torque link, Figure 3.9 (f) shows two distinct major noise sources at 6 kHz. The

first major source comes from the region in which the upper arm connects to the oleo, and the

second is the lower section of the upper torque link arm, just above the connection points between

the two torque link arms.

Figure 3.10 depicts the beamforming map of the landing gear equipped with an upstream torque

link having a separation of L = 1.1D. The case for this configuration is presented separately from

the rest due to its clear distinction in high frequency noise localization results. The dominant noise

source at 6 kHz that was observed for the L = 0.8D, 0.9D, and 1.0D cases (as shown in Figure 3.9

(f)), at the lower portion of the upper torque link arm, no longer appears with the increase of arm

separation to L = 1.1D. Without further study of the flow field, it is difficult to deduce the reason

behind this phenomenon or the other trends observed for torque link noise sources, however, one

possible reason behind the latter observation could be associated with the two torque link arms no

longer meeting at sharp angles relative to each other, due to their increased separation.

Figure 3.10 Beamforming map at 6 kHz for the landing gear equipped with the upstream

torque link that has a torque link arm separation length of L = 1.1D. Flow is from left to

right.

Noise spectra observed by each microphone located in the flyover and sideline linear arrays

were integrated over the 200 – 20,000 Hz range to determine the OASPL directivity patterns for

each landing gear configuration equipped with a torque link. Figure 3.11 presents the OASPL

27

directivities for all torque-link-fitted landing gear configurations considered in this study and for

the baseline landing gear.

Figure 3.11 (a) shows that, in the flyover direction, the noise directivities of landing gears

equipped with a downstream torque link (with any of the L values) are roughly omnidirectional

in the measured flyover angular range of θ = +30° to -30°, as is the case for the baseline gear,

despite the gear with torque links having 3 to 5 dB of higher OASPLs. On the other hand, OASPL

directivity patterns for gears equipped with an upstream torque link (for all L) have their minimum

levels at the overhead location of θ = 0° and their response intensifies by up to 6 dB with the

increase in the flyover angle, θ, in either direction. This indicates a highly directional radiation of

noise in the flyover direction for the landing gear equipped with any of the upstream torque links.

It was also seen in Figure 3.7 that the low-frequency prominent noise emitted by the gear with

upstream torque links was highly directional.

Focusing on Figure 3.11 (b), where the sideline OASPL directivities are presented, the overall

shapes of the directivities between the landing gear with upstream and downstream torque links

look roughly similar, despite greatly varying in OASPLs. After the baseline landing gear, the

quietest configuration is observed to be the model equipped with the downstream torque link

having a separation of L = 1.0D. Its OASPLs are close to those of the baseline landing gear

without torque links. Although, as the sideline angle, φ, moves away from -90°, the OASPL

difference between this configuration and the baseline increases slightly. On the other hand, the

gears equipped with the upstream counterparts of the torque links generate approximately 15 dB

higher acoustic response. It is important to note that with the increase of separation to L =1.1D,

the OASPLs of the two configurations start gradually getting closer to one another. In other words,

the quietest upstream torque link is the one with a separation of L =1.1D while the loudest

downstream configuration is the one with the same separation, as seen in Figure 3.11 (b).

Flyover noise emissions for the landing gears equipped with an upstream torque link were seen

to be highly directional in Figure 3.11 (a). Hence, a further spectral analysis throughout the entire

θ = ±30° range is necessary to understand the origins of this directionality. Figure 3.12 (a) depicts

the OASPL directivity pattern in the flyover direction for the landing gear model equipped with

the upstream torque link having a separation of L = 0.9D, and Figure 3.12 (b) to (h) compares the

noise level spectra of the same landing gear configuration recorded by each microphone in the

flyover linear array along with those of the baseline landing gear.

28

(a) Flyover noise radiation directivities (b) Sideline noise radiation directivities

Figure 3.11 Flyover and Sideline directivities of landing gear with and without torque links.

(d) θ = -10° (e) θ = 0° (f) θ = +10°

(c) θ = -20° (a) Flyover Directivities (g) θ = +20°

Figure 3.12 The OASPL directivity map in the flyover direction and the SPL spectra of the

landing gear between θ = -30° to +30° positions with the upstream torque link having L =

0.9D (shown in green) and without the torque link (shown in black).

(b) θ = -30° (h) θ = +30°

29

In Figure 3.12, for the landing gear with upstream torque link, as the angular location of the

microphone in flyover shifts from being at θ = 0° to θ = ±30° (approaching the sideline direction),

the low frequency peak at 630 Hz gradually grows in amplitude and a new peak at 820 Hz starts

gradually appearing, with the latter eventually becoming the most pronounced at the largest θ

angles. The 630 Hz peak that was observed only in the flyover direction without the presence of

torque links (in Figure 3.3 (a)) gradually intensifies with the presence of the torque link placed

upstream of the strut as one moves from the θ = 0° to θ = ±30°, reaching its maximum directly in

the sideline direction (shown previously in Figure 3.11 (a)). A peak at 820 Hz was not observed

in the overhead spectra given in Fig 3.7 (c), because, as evident from Figure 3.12, it completely

disappears as one moves from θ = ±30° to the overhead location of θ = 0°. Combining this

information with the beamforming map in Figure 3.8 (b), it is evident that the tones observed at

820 Hz originate from the upper arm of the torque links and radiate in the sideline (±y) directions.

3.5 Steering System

This section describes the acoustic response of the nose landing gear with varying steering

system configurations. A total of 20 steering systems were tested, with the first 10 being flat-

tipped steering actuators set between α = 0° to 90°, relative to the freestream flow, as shown in

Figure 3.13. Later, the steering actuators were fitted with rounded caps and the experiments were

repeated. The acoustic results of the gear equipped with flat-tipped actuators are presented here

before the results of the gear with round-tipped actuators. Here, 𝛼 = 0° defines a steering actuator

aligned in the freestream direction with its free end pointing upstream. Additionally, Figure 3.14

shows the steering actuators and steering brackets that make up the steering system on the landing

gear model.

Figure 3.13 Schematic defining the angle α between the axis of the steering actuator and

the freestream direction. Here, the flow direction is from top to bottom.

30

Figure 3.14 Frontal and isometric views of the landing gear with steering actuators

installed at α = 60°.

3.5.1 Flat-tipped Steering Actuators

Narrowband overhead noise level spectra of landing gears equipped with varying orientations

of the flat-tipped steering actuators (from 𝛼 = 0° to 90° in 10° increments) are compared to those

of the baseline landing gear in Figure 3.15. Spectral results for the steering actuator orientations

of 𝛼 = 0° to 40° are collected in Figure 3.15 (a) along with the spectrum of the baseline landing

gear configuration, and the spectra for 𝛼 = 40° to 90° along with that of the baseline case are given

in Figure 3.15 (b). From Figure 3.15 (a) and (b), two general trends can be observed: first, steering

systems mainly contribute to overhead noise that is in the frequency range of 1 kHz and 8 kHz,

and secondly, the noise levels increase from α = 0° to α = 40°, then gradually reduce with

increasing α up to α = 90°. In fact, α = 90°, which is the landing gear configuration with the

actuators placed perpendicular to the freestream flow, produces the lowest noise levels amongst

the landing gear configurations with an actuator assembly.

Figure 3.15 (a) and (b) show that the steering systems do not induce significant acoustic

contribution to the low-frequency peaks that were observed for the baseline configuration at 324

Hz and 648 Hz in the overhead direction. This suggests that these low frequency noise sources do

not originate from anywhere near the steering system installation location or the oleo region.

Between 1 kHz and 2 kHz, a hump exists in the overhead noise spectra for all configurations. This

hump gradually increases in amplitude with the addition of the actuators from α = 0° to 30° (in

Figure 3.15 (a)), and with further increase in α, it gradually decreases till α = 60° (as seen Figure

3.15 (b)). Beyond this angle, the hump in the 1 to 2 kHz range is roughly equal to that found on

the baseline landing gear without a steering system (Figure 3.15 (b)). In Figure 3.15 (a), for α =

31

0°, a sharp peak at 2.5 kHz can be observed with an amplitude of 60 dB, which is a 10 dB increase

from the acoustic response of the baseline gear at the same frequency. This peak gradually reduces

with increasing α from α = 0° to 90°, falling nearly to the baseline level at α = 90° (as can be seen

from a combined inspection of Figure 3.15 (a) and (b)). A further peak in the overhead direction

is notable near 5 kHz, amplitude of which increases with an increase in the steering actuator

orientation from 𝛼 = 0° to 40° and the decreases from 𝛼 = 40° to 90°.

Figure 3.15 Overhead noise spectra of the landing gear with and without the steering

systems. Black arrows indicate frequencies at which the beamforming maps are generated.

Narrowband sideline noise spectra for the landing gear equipped with steering systems

compared to that of the baseline gear are shown in Figure 3.16. Again, results from 𝛼 = 0° to 40°

are collected in Figure 3.16 (a) and from 𝛼 = 40° to 90° are presented in Figure 3.16 (b). From

Figure 3.16 (a) and (b), two general trends can be observed: first, the steering systems are a

significant noise in the sideline direction only for frequencies above 2 kHz, second, despite the

change in noise levels with the steering actuator orientation α, the overall shape of the spectra

stays the same in the presence of the steering systems. The α = 80° orientation produces the lowest

sideline noise levels amongst all the steering actuator orientations considered.

A close inspection of the spectra in Figure 3.16 (a) and (b) above 2 kHz shows, for all

configurations with a steering system, two notable excursions centered near 3 kHz and 5.8 kHz.

The spectral peak at 3 kHz becomes more apparent with its level surpassing the peak at 5.8 kHz

between α = 20° to 60°. For α = 40°, for example, the response at 3 kHz is approximately 9 dB

(a) w/o steering and with steering

at α = 0° to 40°

(b) w/o steering and with steering at

α = 40° to 90°

32

higher than that observed for the baseline landing gear. Beyond α = 60°, the acoustic response

sharply reduces to a level that is only about 1.5 dB above that found on the landing gear without

a steering system. The peak that is centered around 6.9 kHz stays unchanged of the α variations

between α = 0° to 60°. After α = 60°, although this peak is reduced by abound 2 dB, it still remains

3.5 dB higher than the peak of the baseline scenario at the same frequency.

The change of noise level spectra with respect to the steering actuator angle, α, may not be a

result of the steering actuator only. Each variation in α required a slightly modified steering

bracket to be constructed. The variable dimensions of the steering bracket are depicted in Figure

3.3 (a). It is therefore possible that the variance in overhead and sideline acoustic response with

respect to α may be a result of the geometric modifications made to the steering brackets. Thus,

the steering system noise may be predominantly generated within the steering bracket region, or

by the interaction between the wake of the steering actuators and the steering bracket.

Figure 3.16 Sideline noise spectra of the landing gear with and without the steering systems.

Black arrows indicate frequencies at which the beamforming maps are generated.

Figure 3.17 presents the band limited OASPL directivities of landing gear in the sideline

direction (from 𝜑 = -60° to -120°) for varying steering actuator orientations (from 𝛼 = 0° to 90°

in 10° increments), in comparison to that of the landing gear without steering actuators. The range

of frequencies in which the band limited OASPLs have been calculated correspond to the two

excursions observed in sideline spectra given in Figure 3.16 (a) and (b) (one being centred around

3 kHz and the other around 5.8 kHz).

(a) w/o steering and with steering

at α = 0° to 40°

(b) w/o steering and with steering at

α = 40° to 90°

33

(a) In the 2.5 kHz to 4 kHz range (b) In the 5 kHz to 7 kHz range

Figure 3.17 Sideline noise directivities of landing gear with and without steering systems at

two selected frequency ranges.

Figure 3.17 (a) shows that at low α values, the sideline response is asymmetrical compared to

the φ = -90° location between 2.5 kHz to 4 kHz, such that the acoustic response is up to 2 dB

more intense in the upstream sideline direction (φ = -120°) than it is in the downstream sideline

direction (φ = -60°) for α ≤ 50°. For α > 50°, the response in the same frequency range flattens

and becomes omnidirectional similar to the baseline landing gear. Conversely, between 5 kHz to

7 kHz, the baseline configuration as well as the configurations with steering actuators at high α

values, α > 50°, exhibit asymmetrical directivity characteristics, while the steering actuators with

lower α values show an omnidirectional response.

Based on the flyover and sideline microphone measurements in Figure 3.15 and 3.16, and the

directivity results in Figure 3.17, it is evident that, for landing gears equipped with steering

actuators, placing the actuators at α = 40° generates the most intense response, whereas α = 80°

gives the quietist. Also, as α increases, the landing gear starts exhibiting an acoustic response

similar to the baseline gear in terms of the directivity.

It was observed in Figure 3.16 that the steering systems do not contribute to the sideline noise

levels below 2 kHz. These results are further validated by the low frequency beamforming maps,

given in Figure 3.18. In a similar manner as shown earlier in the baseline beamforming results (in

Figure 3.4 (a) and (b)), at lower frequencies such as 824 Hz and 1,504 Hz (in Figure 3.18 (a) and

(b) respectively), the dominant noise generation occurs in the general area of the wheel and strut

region.

34

It is important to note that only a small subset of the beamforming results is depicted here for

brevity. During the course of the investigation, beamforming maps for landing gear configurations

equipped with steering systems spanning all α values have been studied, and it was found that

despite the change in the overall sound pressure levels for each case, the regions in which noise

generation occurred at any given frequency stayed the same. Hence, the gear equipped with the

steering actuators set to α = 40° was selected to draw the most contrast, since this angle was

observed to produce the highest acoustic response. It should also be noted that the frequencies at

which the beamforming maps have been formed for the α = 40° case correspond to the frequencies

where important spectral features were detected in the sideline and overhead noise spectra. The

locations of these frequencies are marked for clarity with arrows in Figures 3.15 (b) and 3.16 (b).

(a) (b)

Figure 3.18 Low frequency beamforming maps of the landing gear installed with steering

actuator having α = 40°. Flow is from left to right.

Figure 3.19 (a) depicts the beamforming maps generated at the sharp peak (near 2.5 kHz) seen

in the overhead noise spectra (in Figure 3.15 (b)) corresponding to the α =40° case, and Figure

3.19 (b) shows near 3-kHz-peak detected in the sideline noise spectra (in Figure 3.16 (b)) for the

α = 40° case. Unlike the low frequency beamforming results presented in Figure 3.18, the

dominant noise source at mid-frequencies is identified to be the steering system. Figure 3.19 (a)

shows that the dominant noise generation at 2,648 Hz occurs around the steering brackets and the

root of the steering actuators. One possible reason for such intense acoustic response in this region

may be the interaction of the flow separating from the steering actuators with the steering actuator

brackets, which is further discussed in Chapter 4. In Figure 3.19 (b), the noise at 3,152 Hz is

35

propagating from the uppermost section of the steering bracket and the upper connection point of

the steering system to the oleo.

(a) (b)

Figure 3.19 Mid-frequency beamforming maps of the landing gear installed with steering

actuator having α = 40°. Flow is from left to right.

Figure 3.20 depicts far-field noise propagation in the sideline direction at 5.8 kHz originating

from the region around the steering brackets. This frequency corresponds to the second peak

observed in the sideline noise level spectrum shown in Figure 3.16 (b).

Figure 3.20 High frequency beamforming map of the landing gear installed with steering

actuator having α = 40°. Flow is from left to right.

36

3.5.2 Round-tipped Steering Actuators

Steering actuators fitted with rounded caps were investigated in order to evaluate the

contribution of the tip shape of the actuator to the noise spectra of the landing gear and to observe

whether the rounding the actuator tips can help reduce noise propagating from this region. A top

view drawing of the gear equipped with round-tipped actuators is shown in Figure 3.21 (a).

Additionally, a photo of the gear equipped with round-tipped actuators with the α = 90°

configuration viewed from downstream can be found in Figure 3.21 (b).

(a) Top view drawing of the gear with

round-tipped actuators

Figure 3.21 Steering actuators equipped with rounded caps viewed from downstream.

Overhead noise spectra of the landing gear equipped with flat and rounded-tip steering

actuators are shown in Figure 3.22 for varying steering actuator orientations (from 𝛼 = 0° to 90°

in 10° increments) along with the spectrum of the baseline landing gear. Based on Figure 3.22 (a)

to (k), four general trends can be observed: first, broadband noise increase of varying levels can

be seen between 1.5 – 2.2 kHz with the addition of rounded caps when compared to the response

of the baseline gear and the gear with flat-tipped actuators. For the α ≤ 70° cases, the response in

this frequency range is up to 6 dB higher for the gear with round-tipped actuators than that of the

gear with the flat-tipped actuators, however, with the increase of α to 80° and 90°, this uptick is

somewhat suppressed, and the difference decreases to 3 dB. Second, the addition of rounded caps

(b) View from downstream of the gear

with round tipped actuators at α = 90°

37

induce varying levels of noise suppression in the 2.2 – 2.8 kHz range for all of the cases, when

compared to the response of the gear with flat-tipped actuators. This suppression is also more

significant for the α ≤ 70° cases, since the suppression also includes the spectral peak at 2.5 kHz,

which is observed only in the landing gears with flat-tipped steering actuators with low α

configurations. Third, the addition of rounded caps induces tonal responses at 2.9 kHz and 3.8

kHz particularly visible for the configurations with α ≥ 30° as can be seen in (c) – (k). These peaks

dominate the round-tipped steering system acoustic response for the α = 60° and 70° cases. Fourth,

some suppression of varying levels of broadband noise in the 0.9 – 1.5 kHz range is seen with the

addition of rounded caps for all α configurations except α = 50°, 60° and 70°.

(a) α = 0° (b) α = 10°

(c) α = 20° (d) α = 30°

Figure continued on the next page.

38

(e) α = 40° (f) α = 50°

(g) α = 60° (h) α = 70°

(j) α = 80° (k) α = 90°

Figure 3.22 Overhead noise level spectra of landing gear equipped with rounded and flat

steering actuators compared to the overhead spectrum of the baseline gear.

Sideline noise level spectra of the landing gear equipped with flat-tipped and round-tipped

steering actuators as well as the baseline gear are shown in Figure 3.23. Observing Figure 3.23

39

(a) to (k), four general trends can be seen regarding the spectra of gear with round-tipped steering

actuators. First, presence of steering actuators (flat- or round-tipped) does not contribute to noise

under 800 Hz. This is expected, since, in Section 3.5.1 steering systems were found not to

contribute to low-frequency noise in the sideline direction. Second, addition of rounded caps

induces a slight rise in noise levels in the 0.9 – 1.5 kHz range in the sideline direction. This uptick

is likely omnidirectional, as it was also seen in the overhead direction in Figure 3.22. Third, the

spectral peak seen at around 3.8 kHz in the overhead direction for the α ≥ 30° configurations in

Figure 3.22 is also seen in the sideline direction, indicating an omnidirectional response. However,

for the α = 0°, 10°, 80° and 90° cases, the response at 3.8 kHz is broadband in the sideline

direction. Fourth, the OASPLs of the gear with either flat-tipped or round-tipped actuators

compared to that of the baseline increases in intensity from α = 0° to α = 40° and sharply decreases

until α = 80°, approaching the response of the baseline landing gear.

(a) α = 0° (b) α = 10°

(c) α = 20° (d) α = 30°

Figure continued on the next page.

40

(e) α = 40° (f) α = 50°

(g) α = 60° (h) α = 70°

(j) α = 80° (k) α = 90°

Figure 3.23 Sideline noise level spectra of landing gear equipped with rounded and flat

steering actuators compared to the spectrum of the baseline gear.

For all α, the sideline spectra of the gear equipped with flat-tipped and round-tipped steering

actuators predominantly differ in the 1.0 – 2.25 kHz and 3.3 – 5.5 kHz ranges as shown in Figure

41

3.23 (a) – (k). Initially thought to be a result of the vortex shedding at the steering actuator tips,

PIV results – to be discussed in Chapter 4 – indicated that the addition of rounded caps suppress

vortex formation, delay separation and induce a higher velocity wake that likely interacts with the

steering brackets.

3.6 Fully-Dressed Landing Gear with Downstream Torque

Link

The tow hook, torque link and the steering system were added to the baseline landing gear to

form the fully-dressed landing gear. As was the case for the previous experiments, different

versions of the fully dressed landing gear were made by varying the torque link distance, L, and

varying the angle between the steering actuators and the freestream flow, α. This section will

discuss the fully-dressed landing gears with the torque link component placed downstream of the

strut as it was not possible to simultaneously install the tow hook, upstream torque link and

steering systems onto the test model due to the restrictions in the steering bracket design. Thus,

only the results from fully-dressed gear with downstream torque links are discussed here. These

discussions will help establish a general trend regarding the interaction noise between the tow

hook, torque link and steering system. It is assumed that there is no interaction noise between the

tow hook and steering systems due to the large distance between these two components.

Furthermore, it was shown in Figure 3.5 and Figures 3.16 and 3.17 that the tow hook and the

steering actuators affect the noise spectrum in different frequency ranges.

The noise level spectra of the fully-dressed landing gear can be separated into two categories:

low α (0° ≤ α ≤ 60°) and high α (70° ≤ α ≤ 90°)) cases, where α shows the angle between the

steering actuator centreline and the freestream flow, since each spectrum in their respective

category followed a similar trend and the trends seemed to be unaffected by the change in torque

link arm separation distance, L. Thus, only one case from each of the categories is shared in this

section.

Overhead noise spectra of the fully-dressed landing gear with its torque link placed

downstream of the strut with a separation distance of L = 1.0D and its steering actuators placed at

α = 30° and α = 90° are shown in Figure 3.24. In the figure, the “measured” spectra, shown in red,

represent the experimental results for the corresponding models. The “calculated” spectra, shown

in black, are formed by subtracting the sound pressure contributions of the landing gear models

42

formed by the individual installment of each of the components to the baseline landing gear from

the baseline model and then adding their contribution linearly to the spectra of the baseline landing

gear model. With such a linear summation of the pressure contribution of each component to the

baseline gear, the interaction noise between components can be identified as the difference

between the “calculated” and “measured” spectra.

Focusing on Figure 3.24 (a), where steering actuator angle is α = 30°, a trough at 640 Hz can

be seen in the calculated overhead spectrum. This is a direct result of the tow hook and torque

link suppressing some of the response at around 640 Hz when they are equipped on the landing

gear model on their own, as was shown in Figures 3.5 and 3.7 (e). The tow hook and the

downstream placement of the torque link were shown to reduce the noise around 640 Hz by

approximately 4 and 1 dB, respectively, when compared to the baseline model. Thus, when their

suppression effects are added linearly, it was assumed that without any interaction between these

components, the levels at 640 Hz would decrease down to 53 dB, as shown in Figure 3.24 (a).

However, when combined, as seen in the measured spectra in Figure 3.24 (a), no noise suppression

can be observed at 640 Hz. Furthermore, the suppression that was induced by the presence of

either the tow hook or the torque link is disappeared, and that the noise level at 640 Hz has reached

almost up to the levels of the baseline gear to roughly 61 dB. It was previously concluded that the

noise at this frequency was originating from the axle and strut area in the inter-wheel region. This

indicates that these two components – being directly upstream and downstream of the strut – may

be acoustically interacting with each other and inducing additional noise that negates the noise

reduction seen by individual installment of those components. However, it is also possible that

the addition of multiple components at once can vary the noise reflection patterns from the

surfaces of the components and cause the sound waves to propagate into the directions that are

not the overhead direction. It was shown earlier in Section 3.5 and in Figures 3.15 that the peaks

at 2.5 kHz and 5.0 kHz in the overhead direction originated from the steering systems. Focusing

on the peaks at 2.5 and 5.0 kHz in Figure 3.24 (a), the noise levels in the presence of both the tow

hook, torque link and steering system fall short by about 2 dB in the measured results compared

to the calculated spectra, indicating that the overhead noise originating from the steering system

at 2.5 kHz and 5.0 kHz are marginally suppressed in the presence of torque link downstream of

the steering system. Focusing on Figure 3.24 (b), for the steering actuator orientation of α = 90°,

the differences in peaks seen earlier in Figure 3.24 (a) for the α = 30° case at 2.5 and 5.0 kHz

between the calculated and measured spectra can no longer be seen. This is likely a result of the

43

larger portionratio of the steering brackets and actuators being positioned far out in ±y directions

at high α configurations (shown in Figure 2.4 (c) as the distance “AB”) such that the torque links

are not situated directly downstream of them, thus, minimizing the interaction between these two

components. Besides the vast difference in spectra at 640 Hz in Figure 3.24 (b), reasons of which

were discussed previously, the calculated and measured spectra share a similar overall shape and

noise levels, indicating minimum interaction between the torque links and steering system for α

= 90°.

(a) α = 30° (b) α = 90°

Figure 3.24: Overhead calculated (black) and measured (red) noise level spectra of the fully

dressed landing gear.

Narrowband sideline noise level spectra of the fully-dressed landing gear with its torque link

placed downstream of the strut with a separation distance of L = 1.0D and its steering actuators

placed at α = 30° and 90° are shown in Figure 3.25. Examining Figure 3.25 (a), where steering

actuator angle is α = 30°, the difference in sideline noise levels between the calculated and

measured spectra is minimal in the entire frequency range, indicating that the flow interactions

between the components do not have a significant impact on the acoustic response of these

components in the sideline direction. The discrepancy at 640 Hz that was seen in the overhead

direction in Figure 3.24 (a) is not observed in the sideline direction in Figure 3.25 (a), further

indicating that this noise is generated in the axle and strut junction region, and propagated between

the wheels in the overhead direction while being shielded by the wheels against sideline

propagation. A slight rise of 2 dB is seen in Figure 3.25 (a) and (b) in the calculated spectra

compared to the measured spectra between 1.3– 1.8 kHz. Noise at this frequency range was found

out to have been caused in the presence of either the torque link or the steering system in Section

44

3.4 and 3.5 and Figures 3.7 and 3.16. This slight difference seen between the measured and

calculated noise levels at this frequency range regardless of the angle, α, may result from the

interaction of upper arm of the torque link with the steering brackets. However, it should be noted

that this slight discrepancy between the measured and calculated spectra may also be a result of

the changed reflection patterns of noise in the presence of both torque link and steering systems

when compared to that in the presence of either torque link or steering systems. Concentrating on

Figure 3.25 (b), a slight downtick (of up to 3 dB) in the 200 – 600 Hz range is seen in the measured

spectrum compared to the calculated spectrum for the steering actuator angle of α = 90°. However,

this is not present for the α = 30° configuration. In Figure 3.25 (a), the reason for this actuator-

orientation-dependent slight uptick is not clear from the present results.

(a) α = 30° (b) α = 90°

Figure 3.25: Sideline calculated (black) and measured (red) noise level spectra of the fully

dressed landing gear.

Based on the sideline spectra of the fully dressed landing gear in Figure 3.25, a discrepancy

between the calculated and measured in the 1.3 – 1.8 kHz range was identified. The reason for

this discrepancy was either the flow interactions between the torque link and steering system

imposing interaction noise, or the changed noise reflection patterns in the presence of both the

torque link and the steering system together on the landing gear. The source of this noise is further

enlightened in the beamforming map conducted at a frequency of 1.5 kHz and the results can be

seen in Figure 3.26. Figure 3.26 (a) shows the beamforming map in the presence of just the

steering system at α = 30° on the landing gear model. The major contributor to noise for this

configuration is the upper wheel region and the downstream area of the strut, and the maximum

contour level is 84 dB. Moving onto Figure 3.26 (b) for the landing gear configuration with just

the torque links, the maximum contour level stays the same at 84 dB, however, moves to the

45

connection point of the two torque link arms and its surroundings. However, with addition of both

the steering actuator and the torque link in the fully dressed configuration seen in Figure 3.26 (c),

the maximum noise levels increase up to 86 dB, and the primary noise source moves to the steering

bracket, strut and torque link connection region. Moreover, a more localized map observed in

Figure 3.26 (c) with a smaller main lobe when compared to the case with just the steering system

in Figure 3.26 (a) indicates a sharper response in that region, thus, further strengthening the case

for acoustic interaction between the upper torque link arm and the steering bracket.

(a) w/ steering, α = 30° (b) w/ torque link, L = 1.0D

Figure 3.26: Beamforming maps of landing gear equipped with steering system, torque link,

and fully dressed landing gear.

4.7 Mach Number Scaling

It was previously mentioned by Ref [15, 17] that the intensity of landing gear noise is

considered to scale with the 6th power of the Mach number for low frequencies and 7th or more

power for high frequencies. Validation experiments were conducted between 54 to 68 m/s of

freestream velocity to find out whether this property holds for the specific landing gear model.

This property can be used to approximate the acoustic response of the landing gear at various

velocities, or can be combined with Strouhal scaling to approximate the acoustic response at

various velocities and model scales, for instance, for interpolating the response of a test model to

that of an full-scale landing gear at operational approach speeds.

Overhead spectra of the fully dressed landing gear with a torque link separation of L = 0.9D

and a steering actuator angle of α = 10° is shown at varying freestream velocities in Figure 3.27

(a). The results for different configurations gave similar trends, however, only one configuration

is shown here as a representation of the phenomenon. As expected, the overall shape of the

spectrum does not depend on the freestream velocity. The spectra are scaled up to a reference

(c) α = 30°, L = 1.0D

46

freestream velocity of U∞ = 68 m/s using M6, M7 and M7.5 as shown in Figure 3.27 (b), (c), and

(d), respectively. Note that these experiments were run consecutively, with a constant outside

temperature. For the M6 scaling, given in Figure 3.27 (a), the data generally collapses to a similar

spectrum for frequencies below 1.5 kHz, with some discrepancies in the 280 – 450 Hz and 850 –

950 Hz ranges. However, for most frequencies above 1.5 kHz (with the exception of the 2.2 – 3.5

kHz range), the spectra of the lower velocities fail to reach the levels of the reference velocity of

U∞ = 68 m/s. Moving onto the spectra scaled with M7, spectra between 1.5 – 20 kHz collapse

under the same spectrum, while frequencies below 1.5 kHz are overcompensated with the M7

scaling. Furthermore, spectra that are scaled up with M7.5 overcompensate the entire frequency

range. Similar results for the sideline direction are shown in Figure 3.28. They follow a

comparable trend to that of the overhead direction and are shown as a reference.

(a) not scaled (b) M6

(c) M7 (d) M7.5

Figure 3.27 Overhead spectra of the fully dressed landing gear model with steering actuators

at α = 10° and a torque link separation of L = 0.9D scaled up to the various powers of the

Mach number with a reference freestream velocity of U∞ = 68 m/s.

47

Ref. [15] and ref. [17] have concluded that a sixth power of Mach scaling was not sufficient

for the entire frequency range. It was found out that despite that M6 successfully collapsing the

spectra in low frequencies, higher frequencies required a higher power of scaling with respect to

the Mach number. Ref [17] demonstrated that M7 scaling succeeded at doing so. Findings from

this study agree with these previous studies in the literature, that for higher frequencies –

approximately above 1.5 kHz for this case – M7 scaling successfully scales the spectra and brings

them under the same reference spectrum.

(a) not scaled (b) M6

(c) M7 (d) M7.5

Figure 3.28 Sideline spectra of the fully dressed lading gear model with steering actuators

at α = 10° and a torque link separation of L = 0.9D scaled up to the various powers of the

Mach number with a reference freestream velocity of U∞ = 68 m/s.

48

Chapter 4

Flow Visualization Results and Discussion

This chapter discusses the Particle Image Velocimetry tests conducted on the horizontal (x-y

plane) plane, focusing on the tip of the steering actuators of the landing gear equipped only with

steering systems (i.e., without tow hook or torque link). Figure 4.1 shows the field of view

employed during these experiments. For a more detailed explanation of the setup, see Figure 2.10

Although experiments were conducted for all α values considered in this study from α = 0° to α =

90° in 10° increments, only a subset of these results are shared in this thesis for brevity.

Figure 4.1 Field of view of the PIV setup.

4.1 Flat-Tipped Steering Actuators

Instantaneous velocity vector fields, skipped every 4 grids in the x direction for visual clarity,

are shown superimposed on instantaneous contours of vorticity for the flow region around the

49

flat-tipped steering actuators in Figure 4.2. Focusing on Figure 4.2 (a), where a representative

snapshot for α = 0° is presented, formations of small-scale vortical structures and separated shear

layer are seen beginning slightly upstream of the steering actuator tip. Looking at the α = 20°

configuration in Figure 4.2 (b), flow separates at the sharp tip of the actuator, and further

downstream, it re-attaches on the actuator surface. Inspection of a series of instantaneous images

depicts also occasional occurrence of a second flow separation from the downstream surface of

the actuator, close to the end of the field of view. At α = 40°, shown in Figure 4.2 (c), the flow

reattaches to the steering actuator surface after a separation at the tip, but this reattachment occurs

at a location much further away from the tip than it was for the α = 20° case. Further downstream,

flow undergoes an occasional second separation with a larger wake formation compared to the α

= 20° case. Starting from the α = 60°, as α is increased to α = 90°, as can be seen from Figure

4.2 (d) to (f), unlike the α = 20° and α = 40° cases, the flow does not re-attach to the actuator

surface within the field of view, forming a spatially larger low-velocity wake region behind the

actuator.

Corresponding results of the time-averaged velocity fields, skipped every 4 grids in the x

direction, are presented together with the time-averaged vorticity contours in Figure 4.3. At α =

0°, given in Figure 4.3 (a), the formation of the separated shear layer slightly upstream of the flat

tip and re-attachment of the flow on the actuator surface, seen earlier in the instantaneous

representation in Fig. 4.2 (a), are clearly evident in the time-averaged pattern. Moving onto Figure

4.3 (b), at both α = 20°, the low-velocity wake region appears downstream of the upper portion of

the actuator, whereas high-velocity downstream flow is directed towards the root of the actuator.

The flow regime of the α = 40°, shown in Figure 4.3 (c), is similar to that of the α = 20° case,

however, with one difference: the time-averaged flow shows a second separation before reaching

the steering actuator root. The existence of a second separation towards the actuator root in the

time-averaged results depicts the strong prevalence of this phenomenon, seen also in the

instantaneous image given in Figure 4.2 (c). This separation is most likely a result of the higher

value of α imposing a greater challenge to the flow to stay attached when compared to the α = 20°

case. Past α = 40°, the character of the flow completely changes, such that the low-velocity wake

region persists till the root of the actuator (within the field of view). Staring from the α = 60°, as

α is increased to α = 90°, in Figures 4.3 (d) to (f), low-velocity wake region spatially enlarges

downstream of the actuator.

50

(a) α = 0° (b) α = 20°

(c) α = 40° (d) α = 60°

(e) α = 70° (f) α = 90°

Figure 4.2 Instantaneous skipped velocity field superimposed on the instantaneous vorticity

contour for the flat-tipped steering actuator. Flow from left to right.

51

(a) α = 0° (b) α = 20°

(c) α = 40° (d) α = 60°

(e) α = 70° (f) α = 90°

Figure 4.3 Time averaged skipped velocity field superimposed on the time averaged vorticity

contour for the flat-tipped steering actuator. Flow from left to right.

52

4.2 Round-Tipped Steering Actuators

Instantaneous velocity vectors, skipped every 4 grids in the x direction, and instantaneous vorticity

contours for the landing gear configuration fitted with round-tipped steering actuators are

superimposed in Figure 4.4. For the α = 0° case, shown in Figure 4.4 (a), a separated shear layer

region develops starting slightly upstream of the tip. This was also observed to be the case for the

flat-tipped actuator in Figure 4.2 (a). Moving onto the α = 20° case, elongated vortical formations

are seen to follow the tip and propagate downstream with the separated flow. The instantaneous

wake region is in the downstream of the upper portion of the actuator as was in the flat-tipped

actuators, however this low-velocity region is smaller for the actuator with the rounded caps then

without. Flow reattaches back onto the actuator surface and these areas contain high-vorticity

regions directed towards the root of the actuator. At α = 40°, shown in Figure 4.4 (c), the low-

velocity wake region is smaller and much closer to the actuator surface compared to the flat-tipped

case for the same α. Perhaps the most surprising finding is for the α = 60° configuration, given in

Figure 4.4 (d), since despite its high angle of incidence, the flow does not detach from the actuator,

and instead, it is directed from the rounded tip towards the root. Increasing α to α = 70° results in

a completely different flow regime in the wake, as seen in Figure 4.4 (e). After passing the rounded

tip, the flow separates from the surface of the actuator. This separation is significantly delayed

compared to the α = 40° given in Figure 4.4 (c). Low-velocity flow dominates only the lower

portion of the actuator as opposed to the flat-tipped actuator at α = 70° (given in Figure 4.2 (e),

where wake region was observed to encompass downstream of the entire actuator. At α = 90°, in

Figure 4.4 (f), the flow downstream of the actuator is completely detached.

Time-averaged skipped velocity fields of the steering actuators superimposed on time-

averaged vorticity contours are shown in Figure 4.5. At α = 0°, in Figure 4.5 (a), the separated

shear layer is again seen to start slightly upstream of the actuator, pushed away from the tip and

then reattach on the actuator. Further downstream, flow re-separates from the actuator surface. At

α = 20° case, flow separates right near the end of the rounded caps and then is re-attached to the

surface of the actuator. This time-averaged pattern confirms what was observed in the

instantaneous image (given in Figure 4.4 (b)), that is, the low-velocity wake region in the

downstream of the upper portion of the actuator is smaller than that for the flat-tipped actuator

with the corresponding α. Increasing the angle to α = 40° in Figure 4.5 (c) shows that the flow is

only attached to the wall for a small fraction of the actuator length. However, this low-velocity

53

(a) α = 0° (b) α = 20°

(c) α = 40° (d) α = 60°

(e) α = 70° (f) α = 90°

Figure 4.4 Instantaneous skipped velocity field superimposed on the instantaneous vorticity

contour for the round-tipped steering actuator. Flow from left to right.

54

(a) α = 0° (b) α = 20°

(c) α = 40° (d) α = 60°

(e) α = 70° (f) α = 90°

Figure 4.5 Time averaged skipped velocity field superimposed on the time averaged vorticity

contour for the round-tipped steering actuator. Flow from left to right.

55

region is spatially much smaller than that of the flat-tipped actuator for the same α (shown in

Figure 4.3 (c)). Moving onto the α = 60° configuration, the time-averaged result in Figure 4.5 (d)

confirms that no significant flow detachment exists for this case. Since the velocity at the surface

is zero and that the flow follows the steering actuator surface, the resulting large velocity gradient

induces high vorticity near the walls. At α = 70°, the flow is now detached, and the low-velocity

wake formation can be clearly seen based on the velocity field. Despite this, the vortical structures

are mainly located around the tip. The low velocity wake region, which dominates downstream

of the entire actuator, does not spread from the surface as far downstream as the wake of the flat-

tipped actuator at the same α. The α = 90° case in Figure 4.5 (f) results in a similar flow regime

to the previous configuration, with the wake spreading somewhat farther away from the actuator

surface compared to α = 70°.

4.3 Effect of the Flow Field on Acoustics

Figure 4.6 OASPLs of the gear with steering actuators in the 200 – 20,000 Hz range

Sideline OASPLs of landing gear equipped with flat and round tipped steering actuators in the

200 – 20,000 Hz range are shown in Figure 4.6. The trends seen for varying α in the OASPL

variations somewhat follows the trends seen in flow behavior around the steering actuators. It was

seen in Figures 4.2 and 4.3 that, for the landing gear equipped with flat-tipped actuators, for α ≤

40°, low-velocity wake region dominates downstream of the upper portion of the actuator (vicinity

56

of the downstream of the tip), whereas high-velocity flow is directed downstream towards the root

of the actuator (within the limits of the field of the view). In Figure 4.6, OASPL slightly and

gradually increases with increasing α until α = 40° for the gear equipped with flat-tipped actuators.

Past α = 40°, OASPL trend changes in Figure 4.6 and starts gradually decreasing with increasing

α. This change in the trend of OASPL coincides with the change in the flow character. Past α =

40°, the low-velocity wake region is directed toward the root of the actuator, spatially growing

with an increase in α. Combining the OASPL trend with the trend seen in the PIV results, it seems

possible that the slower-velocity flow in the root of the steering actuators may be the cause of

suppression in noise due to interaction of the slower velocity flow with the components at the root

of the actuator, such as the steering brackets and oleo. Because of the limitations in PIV field of

view, the flow around the root of the actuators and the bracket region was not investigated in the

present study, however, further PIV investigations in these interaction regions may be helpful to

confirm this possible inference.

OASPLs in Figure 4.6 show that rounding the tip of the actuator leads to a noise reduction

from 0.6 dB to a maximum of 1.6 dB for the actuator orientations considered, which may be a

result of the gradual flow separation rather than a sharp tip. OASPLs from α = 0° to 20° appear to

be unaffected from α. Past α = 20°, OASPL shows a sudden 1 dB increase and stays approximately

at the same constant level from α = 30° to 60°. Past α = 60°, with an increase in α, OASPL

undergoes a gradual decrease. This behavior coincides with the changes in the flow behavior, such

that, at α < 60°, separated low-velocity wake region dominates the downstream of the upper

actuator region (close to the tip), at α = 60°, flow stays attached to the entire actuator surface in

the field of view, while α > 60°, low-velocity flow dominates the lower portion of the actuator

and this low-velocity region enhances with increasing α. Again, lower velocity interaction that

may be taking place around the root of the actuators with the bracket and oleo components may

be the cause for the gradual noise reduction seen for α > 60°.

57

Chapter 5

Conclusion and Future Work

5.1 Conclusion

The research discussed in this work is a component-by component experimental aeroacoustic

study of a simplified nose landing gear, with the main focus on the acoustic response of the tow

hook, and various configurations of torque link and steering systems. Acoustic experiments

consisted of far-field flyover and sideline linear array, and sideline phased array microphone

measurements. Further emphasis on the flow regimes around flat-tipped and round-tipped steering

actuators were made with PIV experiments. The current study utilized the UTIAS Anechoic Wind

Tunnel, for both acoustic and PIV experiments. Experiments, with the exception of Mach scaling

tests were conducted at a freestream velocity of U∞ = 60 m/s, whereas the Mach scaling tests were

conducted at freestream velocities between 54 – 68 m/s. The test model used in the experiments

was a 30%-scale model of a simplified nose landing gear. The modular design of the test model

enabled the individual addition and variation of tow hook, torque link and steering systems onto

the baseline gear, which consisted of the two wheels, axle, strut and the oleo. Torque link arm

separation was varied between L = 0.8D – 1.1D in 0.1D increments, with D being the wheel

diameter. Steering actuator angle was varied between α = 0° – 90° in 10° increments, where α is

the angle between the steering actuator centreline and the freestream velocity. This modular

design enabled the study of acoustic contribution of each of these components to the overall

response of the landing gear.

The acoustic response of the baseline landing gear was observed to be dominated by the low

frequencies. Two spectral peaks centred at 324 Hz and 648 Hz were seen in the overhead direction.

Sideline beamforming maps for this configuration indicated that the major regions for noise

generation were the downstream region of the oleo and the wheels. With the addition of the tow

hook, the spectral peak centred at 648 Hz is suppressed by up to 4 dB. Since this spectral peak

58

was not observed in the sideline direction and was highly dependent on the addition of the tow

hook, it was speculated that this peak is generated in the inter wheel region and radiated between

the wheels in the overhead direction while being shielded by the wheels against sideline radiation.

Landing gear equipped with torque links that are placed downstream of the strut were discovered

to generate less intense noise in both overhead and sideline direction for small arm separation

distances where the inclination of torque link arms with respect to the strut are large. As the torque

link arm separation is increased, and the arms get closer to the strut, a broadband increase in the

acoustic response is observed. Gear equipped with upstream torque links were found out to

generate 18 dB higher levels of spectral peaks centred around 824 Hz than their downstream

counterparts, and overall higher levels in the entire frequency range for the configuration with a

torque link arm separation of L = 0.9D. However, as the separation is increased to L = 1.0D and

1.1D, the peaks become broader and decrease in amplitude by up to 5 dB when compared to the

peak for the L = 0.9D case. These peaks were also seen to exhibit highly dimensional behaviour,

with the peaks disappearing completely 30° off of the sideline direction. As the separation

between the torque link arms increased to L = 1.1D and the arms got closer to the strut as a result

of this, the spectra of the gear with upstream and downstream torque links approached one another

possibly as a result of the torque link-strut pair acting like one bluff body. The steering system

was discovered to predominantly induce noise in the 1 – 8 kHz range, with two prominent spectral

peaks at 2.5 kHz and 5.0 kHz. Noise levels were found out to increase from α = 0° to α = 40° and

then sharply decrease all the way up to about α = 90°. The addition of rounded caps delayed the

sharp decrease in noise levels to the configuration beyond α = 60° and induced a slight (0.6 – 1.6

dB) reduction in the OASPLs. It was discovered that with the addition of both the torque link and

the tow hook for the fully dressed landing gear, which also had the steering system, the

suppression of the low-frequency spectral peak centred at 648 Hz, seen when only the tow hook

or torque link components were installed, diminishes indicating an noise increase at this frequency

as a result of a possible interaction in the inter wheel region between the tow hook and the lower

torque link arm. Moreover, for the fully-dressed configurations with low α, the presence of both

the torque link and steering system was found out to suppress the noise at 2.5 kHz, however,

suppression disappeared for the configurations with high α values.

PIV results collected around the tip of the steering actuator and its wake indicated that for the

flat-tipped steering actuators where α ≤ 40°, low velocity wake was found out to dominate the

vicinity of the downstream of the tip, whereas the high velocity flow was directed towards the

59

root, parallel to the centreline of the actuator within the field of view. OASPLs showed the noise

levels increased up to the α = 40° case for the gear equipped with flat-tipped actuators. However,

after α = 40°, OASPLs gradually decreased with the further increase in α. This trend in the

OASPLs correspond to the trend in the flow regime around the steering actuators. For the α ≥ 40°

cases, the low-velocity wake is directed to the root of the actuator and increasing in size with the

increase in α. These trends indicate that it is possible that the low-velocity flow downstream of

the actuator and near the root may be the reason of the noise suppression for the high α

configurations, where the noise is suppressed due to the low-velocity interaction of the flow with

the steering bracket and the oleo, which are located at the root of the steering actuator. Rounding

the tips of the actuators induce up to 1.6 dB of OASPL reduction. The OASPLs of the actuators

stay relatively same between α = 30° – 60°, and they decrease for higher α configurations. This

trend coincides with the flow regime around the steering actuators, since the low-velocity wake

formation is delayed until α > 60° with the addition of rounded caps.

5.2 Future Work

Spectral analyses and noise localization studies have been conducted on the baseline gear, gear

equipped with either of the tow hook, torque link or steering systems, and the fully-dressed

landing gear which consisted f all three of tow hook, torque link and steering systems added to

the baseline gear. The spectral contribution of each component has now been observed and well

understood. Moreover, a possible mechanism of noise generation has been theorized for the

steering system through the combine consideration of the PIV and acoustic results.

Future research should utilize these findings and equip the landing gear with additional

components such as the landing lights, piping and gear bay doors in order to bring the complexity

of the geometry one step closer to the geometry of the aircraft landing gears. Current research has

utilized the knowledge and expertise of the past research conducted in the same facility while

increasing the complexity one step from previous studies. Thus, the future work should

incorporate these findings as foundational knowledge.

PIV experiments were only limited to the near-tip region of the steering actuators due to the

opaque colour of the actuators, torque link arms and the wheels. Future PIV experiments should

be designed to visualize other regions around the landing gear components such as the torque-link

strut region or the inner steering actuator region by incorporating mirrors into the experimental

60

setup. Moreover, it was theorized that the possible mechanism of noise generation in the steering

system was the impingement of high velocity wake to the steering bracket and strut region.

However, this theory was not materialized as the steering bracket and strut regions were out of

the field of view of the PIV setup. Thus, a flow visualization study in his region can help prove

or disprove this theory and suggest more defined mechanisms of noise generation. Identification

of these noise sources and noise generation mechanisms within the landing gear will eventually

help reduce noise originating from landing gear and its components.

61

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64

(A.1)

(A.2)

(A.3)

(A.4)

Appendix A

Uncertainty Analysis

A.1 Reynolds Number

Root-Sum-Square (RSS) method proposed by Moffat (1988) is used for determining the

uncertainties associated with the Reynolds number [24]. Uncertainty associated with a function,

R, is given by the formula below:

𝜀𝑅 = {∑ (𝜕𝑅

𝜕𝑥𝑖𝜀𝑥𝑖

)

𝑛

1

2

}

0.5

where 𝜀𝑥𝑖 is the uncertainty associated with each independent variable of R. Thus, the uncertainty

associated with the Reynold number can be calculated as:

𝜀𝑅𝑒 = {(𝑣𝐿

𝜇𝜀𝜌)

2

+ (𝜌𝐿

𝜇𝜀𝑣)

2

+ (𝜌𝑣

𝜇𝜀𝐿)

2

+ (−𝜌𝑣𝐿

𝜇2𝜀𝜇)

2

}

0.5

Below is a step by step calculation of each term in 𝜀𝑅𝑒.

Density of air, 𝜌 can be formulated as follows:

𝜌 =𝑃

𝑅𝑇

where P is the air pressure, R is the specific gas constant and T is the temperature. Thus, the

uncertainty of 𝜌 is given by:

𝜀𝜌 = {(1

𝑅𝑇𝜀𝑃)

2

+ (−𝑃

𝑅𝑇2𝜀𝑇)

2

}

0.5

where 𝜀𝑃 = 0.1 𝑃𝑎 and 𝜀𝑇 = 1℃. Thus, for an experiment conducted at T = 20℃ and P = 101.3

kPa, 𝜀𝜌 = 4.28 × 10−6.

Based on a transducer accuracy of 0.25% of the pitot static system, at a freestream velocity of

𝑣 = 60 𝑚/𝑠, 𝜀𝑣 = 0.15 𝑚/𝑠.

The accuracy of the lathe used for machining the parts is 0.001 in. Thus, 𝜀𝐿 = 2.54 × 10−5 𝑚

The viscosity of air can be approximated as follows:

65

(A.5)

(A.6)

(A.7)

𝜇 = 𝜇0 (𝑇

𝑇0)

0.7

Therefore, the uncertainty of the viscosity of air is given by:

𝜀𝜇 = {(𝜇0

𝑇−0.3

𝑇00.7 )

2

}

0.5

For reference values of 𝜇0 = 1.81 × 10−5 𝑘𝑔

𝑚×𝑠 and 𝑇0 = 15℃ (288 𝐾), 𝜀𝜇 = 1.4 × 10−7 𝑘𝑔

𝑚×𝑠.

Thus, for a Reynolds number of 𝑅𝑒 = 600,000 based on a wheel diameter of 𝐿 = 0.152 𝑚 in the

aforementioned conditions, the uncertainty is 𝜀𝑅𝑒 = 4800.

A.2 Spectral Analyses

Uncertainties regarding spectral analyses are calculated based on a formula given by Bendat and

Piersol [25]. The normalized uncertainty is given by:

𝜀 =1

√𝑁

where 𝑁 is the number of overlapping windows used in analysing the data with Fast Fourier

Transform. For all of the spectral analyses shown in this work, 𝑁 = 8192. Thus the normalized

uncertainty is 𝜀 = 0.011 = 1.1%