Embed Size (px)
Citation preview
1
An Experimental Study on the Transient Ice Accretion Process
over the Blade Surfaces of a Rotating UAS Propeller
Yang Liu1, Linkai Li2, Zhe Ning3, Wei Tian4 and Hui Hu5()
Department of Aerospace Engineering, Iowa State University, Ames, Iowa, 50010
In the present study, a comprehensive experimental study was conducted to investigate the
transient ice accretion process over the blade surfaces of a rotating UAS propeller. The
experiments were performed in the Iowa State University Icing Research Tunnel (i.e., ISU-
IRT) with a scaled UAS propeller model operated under a variety of icing conditions (i.e.,
ranged from rime to glaze ice accretion). A “phase-locked” high-speed imaging technique was
developed to resolve the transient details of the unsteady icing processes. Simultaneously, the
dynamic aerodynamic forces acting on the propeller was also measured quantitatively along
with the input power measurements in driving a constant propeller rotation during the ice
accretion processes. Based on such temporally-synchronized-and-resolved measurements, the
detailed ice accretion phases were correlated with the dynamic aerodynamic force data. The
dynamics of the leading-edge ice accretion was also quantified as a function of time and space.
The ice accretion features as well as their effects on the propeller performance were evaluated
in moderate to severe icing conditions. Such quantitative measurements of the transient ice
accretion provided further insight into the underlying physics for the icing processes on the
rotating UAS propeller.
Nomenclature
R = Radius of the propeller, m
H = Height of the propeller mount, m
dmount = Diameter of the mount tube, m
dsupport = Diameter of the supporting tube, m
a1 = Length of the mount, m
a2 = Length of the coniptical cone, m
a = Total length of the mount and cone, m
r = Local rotational radius, m
CT = Thrust coefficient
CP = Power coefficient
η = Propeller efficiency
J = Advance ratio of the propeller
U = Freestream velocity of airflow, m/s
n = rotational speed of the propeller, rpm
D = Diameter of the propeller, m
T = Thrust force acting on the propeller, N
ρ = Air density, kg/m3
P = Power to drive the propeller, W
β = Water collection efficiency
1 Post-Doctoral Research Associate, Department of Aerospace Engineering. 2 Graduate Student, Department of Aerospace Engineering. 3 Graduate Student, Department of Aerospace Engineering. 4 Assistant Professor, School of Aero. & Astro., Shanghai Jiaotong University. 5 Martin C. Jischke Professor, Department of Aerospace Engineering, AIAA Associate Fellow,
Email: [email protected]
LWC = Liquid water content, g/m3
K = Droplet inertia parameter
ρw = Water density, kg/m3
d = Droplet diameter, m
µ = Air viscosity, Pa∙s
L = Characteristic length of a body, m
Fc = Centrifugal force, N
m = Water mass at a blade span location, kg
ω = Angular speed of the propeller, rad/s
We = Weber number
σ = Surface tension, Pa
T∞ = Ambient temperature, °C
θ = Twist angle of blade, °
α = Effective angle of attack, °
V1 = Local resultant velocity, m/s
V2 = Local tangential velocity, m/s
Hice,img = Imaged leading-edge ice thickness, mm
Hice = Actual leading-edge ice thickness, mm
uice = Leading-edge ice growth rate, mm/s
Dow
nloa
ded
by I
OW
A S
TA
TE
UN
IVE
RSI
TY
on
Febr
uary
13,
201
7 | h
ttp://
arc.
aiaa
.org
| D
OI:
10.
2514
/6.2
017-
0727
55th AIAA Aerospace Sciences Meeting
9 - 13 January 2017, Grapevine, Texas
AIAA 2017-0727
Copyright © 2017 by Yang Liu, Linkai Li, Zhe Ning and Hui Hu. Published by the American Institute of Aeronautics and Astronautics, Inc., with permission.
AIAA SciTech Forum
2
I. Introduction
cing has been widely recognized as a big threat to the flight safety and performance of aircraft in cold weather. Even
small amounts of ice accretion can reshape the aerodynamic outlines of aircraft and negatively affect the flight
performance. The icing process and its effect on aircraft wings have been extensively studied [1–4], and much effort
has been made to predict icing conditions using theoretical and computational models [5–8]. The ice formation on
aircraft wings can be in a variety of structures, depending on the flight conditions and environmental parameters [9].
For example, when the ambient temperature is cold (typically below -10 ºC) and the air has a small liquid water content
(LWC), the supercooled water droplets freeze immediately upon impact with the wings, forming rime ice. At warmer
operating temperatures, i.e., just below freezing, if the air has a larger LWC, the impinged water does not immediately
freeze, but transports along the wing surface prior freezing downstream, forming the complicated ice shapes, called
glaze ice. Glaze ice is considered as the most dangerous ice accretion as it usually forms in “horns” and larger
“feathers” growing outward into the airflow [10]. These irregular ice shapes cause large scale flow separation which
produces dramatic increases in drag and decreases in lift.
In recent years, unmanned aircraft system (UAS) has become one of the most remarkable developments in aviation,
for both commercial and military uses. Free from having to accommodate the safety needs and endurance limits of
onboard pilots, UASs are capable of flying extended missions and venturing into hazardous and remote locations.
However, the vulnerability of UAS was revealed when they were exposed to icing conditions [11]. Unlike most large
manned aircraft using turbo jet engines for propulsion, almost all the UASs are powered by propellers. Since ice may
accumulate on every exposed frontal surfaces of UAS, not only on wings, but also on rotating blades of propellers,
the aerodynamic performance of UAS can be significantly degraded in icing conditions. In moderate to severe
conditions, the propellers can become so iced up that continued flight would become impossible. Consequently, UAS
missions in cold weather are usually delayed or canceled when icing conditions persist, which substantially limits the
applicability of UASs.
In looking to develop the all-weather capabilities of UAS, much effort has been made to look for solutions of UAS
operations in icing conditions. Siquig [12] evaluated the coupling of icing properties and vehicle characteristics by
comparing two different unmanned aerial vehicle (UAV) systems (i.e., High flying long endurance UAVs vs. Low
flying short endurance UAVs). The impacts of icing on the UAV operations were revealed. Bottyán [13] developed
an in-flight structural icing estimation method based on a simple 2D ice accretion model predictions. The effects of
ambient air temperature, LWC, airfoil geometry and airspeed on the icing process on UAS wings were evaluated.
Szilder and McIlwain [11] revealed the influence of Reynolds number on the ice accretion process over a NACA 0012
airfoil for UAS applications. The regimes of rime and glaze formation as well as the ice accretion extent as a function
of meteorological conditions were identified. Armanini et al. [14] proposed an icing-related decision-making system
(IRDMS) to quantify in-flight icing based on changes in aircraft performance and measurements of environmental
properties. Sørensen [15] integrated a power control system and an electrically conductive carbon-nano-material-
based coating for temperature control of UAS airfoil surfaces to address the issue of structural change due to ice
accretion.
Surprisingly, while a number of studies were performed to address ice accretion and anti/de-icing over UAS wings,
almost no research work can be found in literatures to examine the ice accretion process and resultant aerodynamic
performance of rotating UAS propellers. In comparison to that over fixed-wings, the dynamic ice accretion process
over the blade surfaces of rotating UAS propellers would become even more complicated, due to the combined effects
of aerodynamics shear forces and centrifugal forces. Leveraging the previous studies of rotor icing posed on
helicopters and wind turbines, it is suggested that the ice accretion on rotor blades can be critically dangerous as it
modifies the aerodynamic profiles, creates excessive vibration, increases weight and drag, and introduces mass and
aerodynamic unbalance concerns as ice sheds off [16,17]. Lamraoui et al. [16] revealed that the uneven distribution
of relative velocities along a rotor blade caused a dramatic variation in ice accretion mass, with the maximum mass
rate located on the outer section and the minimum near the blade root zone. Similarly, a linear distribution of the
stagnation ice thickness along the blade (i.e., a linear increase from the hub to the tip) was also suggested by Fortin
and Perron [18]. Such unbalanced three-dimensional mass distribution drastically degraded blade aerodynamic
performance and increased blade fatigue. While the mass of rime ice accretion was dependent on blade span locations,
the glaze ice growth was suggested to be even more complicated due to the surface water transport along the blade
span driven by the action of centrifugal forces [19]. In comparison to the icing processes on helicopter and wind
turbine rotors, the ice accretion on UAS propellers can be rather different due to the unique blade design and airfoil
profiles adopted in UAS applications as well as the different flow and rotational conditions. The availability of
I
Dow
nloa
ded
by I
OW
A S
TA
TE
UN
IVE
RSI
TY
on
Febr
uary
13,
201
7 | h
ttp://
arc.
aiaa
.org
| D
OI:
10.
2514
/6.2
017-
0727
3
temporally-resolved ice accretion measurements on the blades of a rotating UAS propeller is highly desirable in order
to elucidate underlying physics to improve our understanding about the transient ice accretion process during dynamic
rotation and the aerodynamic responses of the iced UAS propellers.
For most of the previous ice accretion studies, the accreted ice shapes were measured by tracing an outline of the
ice accretion, which is intrusive and provides measurement of the ice shape at only a single time [20,21]. The
nonintrusive measurements of ice accretion have been achieved by using photography techniques [20,22–24]. Vargas
and Tsao [20] used macrophotography images of the ice shapes with a tape measure to characterize the ice accretion
shapes. Kraj and Bibeau [22] digitized the snapshots of ice accretion to measure the growth rate of leading-edge ice
thickness during active de-icing operations. Most of these previous experiments have examined the accreted ice shapes
and their aerodynamic characteristics at a specified time, which ignores many important transient details in ice
accretion. More recently, Rye and Hu [23] developed a high-speed imaging technique capable of providing time-
resolved measurements of spatial distribution of water and ice throughout the icing processes to quantify the transient
ice accretion details over a NACA0012 airfoil.
In the present study, a comprehensive experimental study was conducted to investigate the transient ice accretion
processes over the blade surfaces of a rotating UAS propeller. The experiments were performed in the Iowa State
University Icing Research Tunnel (i.e., ISU-IRT) with a scaled UAS propeller operated under a variety of icing
conditions (i.e., ranged from rime to glaze ice accretion). A “phase-locked” high-speed imaging technique was
developed to resolve the transient details of ice accretion over the rotating propeller blades. Simultaneously, the
dynamic aerodynamic forces acting on the propeller was also measured quantitatively along with the input power
measurements in driving a constant propeller rotation throughout the ice accretion processes. During the experiments,
while the freestream velocity of incoming airflow and the rotational speed of the UAS propeller were kept at pre-
scribed levels, varying the ambient temperature and LWC level led to different types of ice accretion on the rotating
propeller blades. Based on such temporally-synchronized-and-resolved measurements, the detailed ice accretion
phases were correlated with the dynamic aerodynamic force data. The dynamics of the leading-edge ice accretion was
also quantified as a function of time and space. The ice accretion features and their effects on the propeller performance
were evaluated in moderate to severe icing conditions as well.
II. Experiments
A. Propeller Design
The model propeller used in the present study was designed based on a typical three-blade propeller seen in modern
propeller-powered UASs. Figure 1 shows a schematic of the model propeller along with typical cross section profiles
of the propeller blade. The model propeller has a radius of 100 mm and the coniptical cone of the propeller has a
diameter 33 mm. With the scale ratio of 1:14, the model propeller would represent 600 kW turboprops with a propeller
diameter of about 2.8 m commonly seen in modern propeller-powered UASs. The blades and the coniptical cone of
the model propeller are made of a hard plastic material (i.e., VeroWhitePlus, RGD835, manufactured by Stratasys,
Inc.) by using a rapid prototyping machine (i.e., 3D printer).
Figure 1: Design of the model UAS propeller used in the present study. (a) schematic of the model propeller
layout and blade design details; (b) Variation of blade chord length and twist angle along the span direction.
Dow
nloa
ded
by I
OW
A S
TA
TE
UN
IVE
RSI
TY
on
Febr
uary
13,
201
7 | h
ttp://
arc.
aiaa
.org
| D
OI:
10.
2514
/6.2
017-
0727
4
The propeller blades have the typical airfoil cross sections and platform profiles used in modern propeller design.
As shown in Fig. 1(a), two airfoil profiles (i.e., ARA-D 13% and ARA-D 20%) are used at different spanwise locations
along the propeller blade. The ARA-D 20% airfoil profile is used between 0.10R and 0.30R, while the ARA-D 13%
airfoil is positioned from 0.30R until the blade tip. With the prescribed blade platform profiles and twist angles (i.e.,
optimized based on the designed testing freestream velocity of airflow and rotational speed of the propeller) as shown
in Fig. 1(b), a spline function is used to interpolate the prescribed cross section profiles to generate the three-
dimensional model of the propeller blade using SolidWorks software.
TABLE 1. The design parameters of the UAS propeller model.
Parameter R (mm) H (mm) dmount (mm) dsupport (mm) a1 (mm) a2 (mm) a (mm)
Dimension 100.00 196.10 32.20 30.00 100.96 41.00 146.86
The primary design parameters of the model propeller are listed in Table 1. In the present study, the propeller
blades were mounted on a hard plastic tube with a coniptical cone headed at zero pitch angle. A brushless motor (DJI
2212, 940KV) was installed inside the supporting tube, which would drive the UAS propeller to rotate. An aluminum
tube with streamlined cross section was used to support the model propeller.
B. Experimental Setup and Methods
This experimental study was performed in the ISU Icing Research Tunnel (ISU-IRT) as schematically shown in
Fig. 2. The icing research tunnel provides a unique facility for conducting fundamental experimental studies in icing-
related scenarios. The icing tunnel can run over a range of test conditions to duplicate various atmospheric icing
phenomena (e.g., from rime, mixed to glaze icing). The facility provides the capabilities to perform experiments at
temperatures as low as -25 °C and at wind speed up to 60 m/s. A pneumatic spray system is integrated in the wind
tunnel, which is capable of generating water droplets of 10–100 μm in diameter with the LWC adjustable to more than
5.0 g/m3.
Figure 2: A schematic of the ISU-Icing Research Tunnel (ISU-IRT).
Figure 3 shows the schematic of the experimental setup for measuring the ice accretion on the blade surfaces of
the rotating model UAS propeller. In the present study, the model propeller was driven by a brushless motor (DJI
2212, 940KV), which was powered by a direct current (DC) power supply (VOLTEQ HY3050EX). This high quality
regulated power supply can be continuously adjustable at 0-30V DC and 0-50A. During the experiments, the rotational
speed of the model propeller was adjusted with a speed controller by changing the signal duty cycle of the brushless
motor. In the meantime, the rotational speed of the model propeller was measured by a tachometer (MONARCH
PLT200) that can generate a pulse signal from each rotation of the propeller. The tachometer-generated pulse signal
was then simultaneously scanned by a 16-bit data acquisition system (NI USB-6218) and sent to a digital delay/pulse
generator (BNC Model-577). Thus, a control system for the model propeller with feedback compensation was
integrated. A proportional-integral-derivative (PID) algorithm (i.e., a control loop feedback mechanism) was then
formulated to achieve automatic correction of the propeller rotational speed as a disturbance occurs (e.g., ice accretion
or ice shedding). In the present study, the currents and voltages of the DC power supply were also scanned by the data
acquisition system, which provides the temporally-resolved input power measurements of the model propeller.
Dow
nloa
ded
by I
OW
A S
TA
TE
UN
IVE
RSI
TY
on
Febr
uary
13,
201
7 | h
ttp://
arc.
aiaa
.org
| D
OI:
10.
2514
/6.2
017-
0727
5
Figure 3: A schematic of the experimental setup for the ice accretion imaging and force & power measurements
on the rotating model UAS propeller.
As shown in Fig. 3, the aluminum supporting tube for the propeller model was connected to a high-sensitivity
force-moment sensor (JR3 load cell, model 30E12A-I40) to measure the dynamic aerodynamic forces acting on the
model propeller. The JR3 load cell is composed of foil strain gage bridges, which are capable of measuring the forces
on three orthogonal axes, and the moment (torque) about each axis. The precision of the force-moment sensor cell for
force measurements is ±0.25% of the full range (40 N). During the experiments, the aerodynamic force data was
sampled at a rate of 5000 Hz for each test case.
In addition to the aerodynamic force measurements, a high-speed imaging system was also used in the present
study to conduct temporally-resolved imaging of ice accretion on the blade surfaces of the rotating propeller under
controlled environmental conditions. The imaging of the transient ice accretion process was achieved by using a high-
speed camera (PCO Tech, pco.dimax S4) with a 50 mm macrolens (Nikon, 50 mm Nikkor 1.8D). The camera was
positioned normal to the freestream direction of airflow, providing a view with a 2014 × 526 pixels2 field of view and
a pixel resolution of 11.56 pixels/mm. To provide high quality imaging data, low-flicker illumination was provided
by a pair of 100 W Studio-LED lights (RPS Studio Light, Model RS-5610 and RS-5620).
In the present study, both the force measurement system and the high-speed imaging system were connected to a
digital delay/pulse generator. As a trigger signal (i.e., tachometer-generated pulse signal) was received by the
delay/pulse generator, two separate pulse signals would be simultaneously sent out from two channels of the
delay/pulse generator to the force measurement system and the imaging system, to initiate the temporally-
synchronized measurements. For the high-speed imaging system, each pulse signal would further burst three pulses
for image acquisition at a specified frequency (i.e., which is the triple of the propeller rotational frequency). By
adjusting the time delay between the trigger signal from the tachometer and the signal output to the imaging system,
images of ice accretion on the propeller blades would be captured when the blades passed through the focal plane of
the high-speed camera. Thus, “phase-locked” high-speed imaging of the transient ice accretion processes over the
blade surfaces of the rotating model propeller was accomplished.
C. Evaluation of Propeller Performance
The propeller performance is characterized using the thrust coefficient (CT), power coefficient (CP), and propeller
efficiency (η) [25], which are usually plotted against the advance ratio (J), i.e., the ratio of the freestream velocity of
airflow to the propeller tip speed. The definitions for the advance ratio, thrust and power coefficients, and propeller
efficiency are given by
U
JnD
(1)
where U is the freestream velocity of the airflow, n is rotational speed of the propeller, D is the diameter of the
propeller.
Dow
nloa
ded
by I
OW
A S
TA
TE
UN
IVE
RSI
TY
on
Febr
uary
13,
201
7 | h
ttp://
arc.
aiaa
.org
| D
OI:
10.
2514
/6.2
017-
0727
6
2 4T
TC
n D (2)
where T is thrust force acting on the propeller, and ρ is the air density.
3 5P
PC
n D (3)
where P is the input power to drive the propeller.
T
P
CJ
C (4)
Figure 4: Performance of the model propeller with no ice accretion. (a) Thrust and power coefficients vs.
advance ratio; (b) Propeller efficiency as a function of advance ratio.
In order to determine the optimum advance ratio of the model propeller, we examined the propeller performance
at a series of advance ratio values (i.e., from J = 0.6 to 1.8) by changing the freestream velocity of airflow while the
rotational speed of the propeller was kept constant (i.e., n = 3000 rpm in the present study). Figure 4 illustrates the
performance of the model propeller used in the present study. As clearly shown in Fig. 4(a), when the advance ratio
increases from J = 0.6 to 1.8, the thrust coefficient decreases from CT = 0.15 to 0.03 (i.e., with a factor of 5), while the
power coefficient decreases from CP = 0.4 to 0.1 (i.e., with a factor of 4). By using Eq. (4), the curve of propeller
efficiency vs. advance ratio was acquired as shown in Fig. 4(b). It is noticed that as the advance ratio increases from
J = 0.6 to 1.6, the propeller efficiency continuously increases to the maximum (i.e., from η = 0.23 to 0.78). When the
advance ratio is further increased beyond J = 1.6, the propeller efficiency begins to decrease, and has a dramatic drop
at J = 1.8 (i.e., η = 0.50). Therefore, the optimum advance ratio for the model propeller used in the present study (at
the rotational speed of n = 3000 rpm) was determined to be J = 1.6 (i.e., with the corresponding freestream velocity
of the airflow being U = 16 m/s).
III. Icing Process over a Rotating Propeller Blade
As mentioned above, icing occurs when the supercooled water droplets suspended in the airflow impact on a body.
Depending on the environmental parameters and the body geometry, the ice formation can be in a variety of structures.
One of the key icing parameters is the water collection efficiency, β, which specifies the distribution of water collection
upon impact. It is mainly determined by the relative size of the droplets compared with the body, and the droplet
inertia upon impact, which is characterized by the droplet inertia parameter [26]:
2
18
wd U
KL
(5)
where ρw is the water density, d is the droplet diameter, µ is the air viscosity, and L is the characteristic length of the
body (i.e., which is usually taken as the radius of leading-edge curvature for airfoils).
When the impinged water is deposited on the body, depending on the ambient temperature, it is either frozen
immediately (i.e., rime icing) or transported downstream before freezing (i.e., glaze icing). In the latter case, if the
Dow
nloa
ded
by I
OW
A S
TA
TE
UN
IVE
RSI
TY
on
Febr
uary
13,
201
7 | h
ttp://
arc.
aiaa
.org
| D
OI:
10.
2514
/6.2
017-
0727
7
water is deposited on a two-dimensional airfoil, the surface water transport is only driven by aerodynamic shear
stresses in a Couette flow [27, 28]. However, for the water/ice accretion over a rotating propeller blade, the surface
water transport would become even more complicated, due to the combined effects of aerodynamics shear force and
centrifugal force. Figure 5 shows a schematic of the surface water transport over the blade surface of a rotating
propeller. The impinged water forms a surface water film and transports downstream driven by the boundary layer
airflow. Due to the action of the centrifugal force, the advancing behavior of the water film becomes three-
dimensional, transporting along both blade chord and span. The definition of the centrifugal force is given by
2
cF m r (6)
where m is the water mass at a blade span location, ω is the angular speed of the rotating propeller, r is the local
rotation radius.
Figure 5: A schematic of the surface water transport over the blade of a rotating propeller.
Driven by the aerodynamic shear stresses and the centrifugal force, the deposited water is redistributed to other
regions of the propeller blade, and may break up into rivulets due to the interplay of surface tension forces and inertial
forces [23], which can be characterized using the Weber number [26]:
2U L
We
(7)
where ρ is the air density, σ is the surface tension at the water/air interface.
In the present study, typical ice accretion trials (i.e., rime and glaze) on the rotating propeller blades were conducted
at the advance ratio of J = 1.6 (i.e., with the freestream velocity being U = 16 m∕s while the rotational speed of the
propeller was kept at n = 3000 rpm). The effects of ambient temperature (i.e., varied form T∞ = -15 to -5 °C) and LWC
(i.e., varied form LWC = 0.5 to 2.0 g/m3) on the ice accretions process were examined.
IV. Measurement Results
A. “Phase-locked” High-speed Imaging of Ice Accretion over the Rotating UAS Propeller Blade
As described above, “phase-locked” high-speed imaging was conducted to provide “frozen” images of the ice
features accreted on the rotating propeller blades as a function of time. Figure 6 shows the icing process over the
propeller blade at the rotational speed of n = 3000 rpm, with the freestream velocity of U = 16 m/s, LWC = 1.0 g/m3,
and the ambient temperature of T∞ = −15°C. The ice accretion exhibited typical rime characteristics (i.e., with white
and opaque appearances) as described in previous works [29]. As the supercooled water droplets impinged on the
propeller blade, since the ambient temperature was freezing cold, the cooling actions of heat convection and
conduction were sufficiently strong to remove the latent heat of fusion released during the phase change from liquid
water to solid ice. As clearly shown in Fig. 6(a), the impinged water froze upon impact and formed an ice layer
Dow
nloa
ded
by I
OW
A S
TA
TE
UN
IVE
RSI
TY
on
Febr
uary
13,
201
7 | h
ttp://
arc.
aiaa
.org
| D
OI:
10.
2514
/6.2
017-
0727
8
conforming well with the blade leading-edge. Marks of surface water transport (i.e., rivulets-shaped features) were
not observed. It is suggested that the formation of rime ice on the rotating blade is not affected by the centrifugal force
due to the immediate freezing of the impinged droplets. As more water impinged and froze on the blade, the ice layer
at the leading-edge was thickened as can be seen in Fig. 6(b) to 6(f). Along with the thickening of the leading-edge
ice layer, growth of ice “feathers” around the mid-span (i.e., between 0.5R and 0.8R) was also observed as shown in
Fig. 6(c) to 6(f). The ice “feathers” were found to grow back toward the leading-edge, accreting together with the
leading-edge ice layer. It is suggested that the ice “feathers” formed the wake regions, whereby flow separation was
formed behind the thickening ice layer and prevented water from impinging behind the ice layer. Only the ice layer
that reached out into the outer flow continued to grow outward and into the oncoming flow as can be seen in Fig. 6(g)
and 6(h).
Figure 6: “Phase-locked” images of the ice accretion process on the rotating propeller blade at n = 3000 rpm
with the freestream velocity, U = 16 m/s; LWC = 1.0 g/m3; and T∞ = −15°C.
At warmer ambient temperatures, e.g., for the test case of T∞ = −5°C while the other environmental parameters
were kept the same, the ice accretion exhibited typical glaze characteristics (i.e., with transparent and clear
appearances) as shown in Fig. 7. For the ice accretion at such warmer temperatures, since the heat transfer (i.e., heat
convection and heat conduction) is not adequate to remove all of the latent heat of fusion in the impinged water
droplets, only a fraction of the water froze upon impact while the remaining transported on the blade surfaces of the
rotating propeller driven by the aerodynamic shear force and the centrifugal force. As can be seen in Fig. 7(a) and
7(b), when the water droplets impinged on the blade, an ice layer formed at the leading-edge. In comparison to the
rime icing mentioned above, the ice layer formed in this case has an extended coverage on the inner blade section.
Such larger ice coverage to downstream is suggested to be caused by the surface water film run-back driven by the
aerodynamic shear force. It is also noticed that there is a “needle-shaped” ice tip formed at the blade tip as can be seen
in Fig. 7(b), which is due to the water transport along the blade span as driven by the centrifugal force. As more water
impinged on the blade, the ice layer at the leading-edge was thickened while the “needle-shaped” ice at the blade tip
became more evident as can be seen in Fig. 7(c) to 7(f). It is also found that the ice roughness began to grow behind
the ice layer as shown in Fig. 7(f). It has been revealed by Waldman and Hu [23] that in a typical glaze ice accretion,
water would bead up at the advancing film front near the airfoil leading-edge, where the roughness would exhibit. As
the ice layer at the leading-edge was further thickened, the ice thickness along the blade span was rather uniform in
the icing images as shown in Fig. 7(g) and 7(h). Since the imaging direction is normal to the blade span in the present
study, yet the blade has twist angles along the span, the actual ice thickness at the leading-edge is suggested to vary
along the blade span, which is to be discussed in detail later.
Dow
nloa
ded
by I
OW
A S
TA
TE
UN
IVE
RSI
TY
on
Febr
uary
13,
201
7 | h
ttp://
arc.
aiaa
.org
| D
OI:
10.
2514
/6.2
017-
0727
9
Figure. 7: “Phase-locked” images of the ice accretion process on the rotating propeller blade at n = 3000 rpm
with the freestream velocity, U = 16 m/s; LWC = 1.0 g/m3; and T∞ = −5°C.
Figure. 8: “Phase-locked” images of the ice accretion process on the rotating propeller blade at n = 3000 rpm
with the freestream velocity, U = 16 m/s; LWC = 2.0 g/m3; and T∞ = −5°C.
As the ambient temperature was still kept at T∞ = −5°C, while the LWC was increased from LWC = 1.0 g/m3 to
2.0 g/m3, the ice accretion on the rotating blade was still glaze, but with more ragged three-dimensional ice features
as clearly shown in Fig. 8. In this test case, since the LWC was doubled in quantity while the freestream velocity of
the airflow was still kept constant at U = 16 m/s, the mass flow rate of the water impinged on the blade was significantly
Dow
nloa
ded
by I
OW
A S
TA
TE
UN
IVE
RSI
TY
on
Febr
uary
13,
201
7 | h
ttp://
arc.
aiaa
.org
| D
OI:
10.
2514
/6.2
017-
0727
10
increased, while the aerodynamic shear force remained the same. As the water droplets impinged on the blade, a water
film quickly formed at the leading-edge, and ran back under the action of the aerodynamic shear force as can be seen
in Fig. 8(a) and 8(b). In the meantime, the water film advancing to the downstream locations quickly broke into
rivulets, which transported along the blade span under the action of the centrifugal force as can be found in Fig. 8(b).
It is also noticed that there was a “needle-shaped” ice formed at the blade tip due to the centrifugal effect. As more
water impinged on the blade, the transverse rivulets became more evident as can be seen in Fig. 8(c) and 8(d). Near
the leading-edge, the ice roughness began to exhibit, while the “needle-shaped” ice at the blade tip grew into multiple
ice branches as shown in Fig. 8(d). As the ice layer thickened and the roughness elements grew, the ice branches
reached out into the outer flow continued to grow as shown in Fig. 8(e) and 8(f), due to the large amount of water
transport towards the blade tip driven by the centrifugal force. Combined with the aerodynamic shear force, the water
film at the leading-edge broke and froze into the “lobster-tail-liked” ice features as clearly shown in Fig. 8(g) and Fig.
8(h). As more water transported towards the outer blade and froze into the irregular ice features around the blade tip,
the centrifugal force acting on the ice features continuously increased. Once the centrifugal force overcame the
adhesion force within the ice or between the ice layer and the blade surface, ice shedding would occur. The shedding
of ice pieces due to delamination of ice layers was also observed in glaze icing.
Here, the “phase-locked” high-speed imaging of the ice accretion over the rotating propeller blades provides a
qualitative view of the transient processes involved during icing, which lends some insight into the physics of ice
formation (i.e., as a result of combined effect of the aerodynamic force and the centrifugal force). However, more
quantitative measurements of the ice features as well as their effects on the propeller performance are needed to study
the icing process in detail.
B. Dynamic Thrust and Power Measurements of the Rotating UAS Propeller in Transient Icing Processes
As described above, the JR3 force-moment sensor used in the present study can provide time-resolved
measurements of all three components of the aerodynamic forces and the moment (torque) about each axis. While
similar features were also revealed by the other components of the aerodynamic forces and the moments, only the
measured thrust coefficient, CT, is given in the present study for analysis for conciseness.
Figure 9 gives examples of the aerodynamic force measurement results in term of the instantaneous thrust
coefficients with the model propeller operated at a constant rotational speed of n = 3000 rpm and advance ratio of J =
1.6 in three typical icing conditions. With the freestream velocity of the airflow being kept constant at U = 16 m/s, the
time history of the measured instantaneous thrust coefficient acting on the ice accreting propeller at the ambient
temperature of T∞ = -15 °C and LWC = 1.0 g/m3 is shown in Fig. 9(a). The Gaussian-filtered mean values and the ±
1 standard deviation (std) bounds were also given in the plot for comparison. In the present study, the aerodynamic
forces were first sampled for 15 seconds before the ice accretion began, and then sampled for more than 105 seconds
during the icing process. The aerodynamic forces acting on the propeller were found to be highly unsteady with their
magnitudes fluctuating significantly as a function of time. Compare the thrust coefficients acting on the propeller
before and during the ice accretion, the force fluctuation amplitude (i.e., as indicated by the ± 1 std bounds) was found
to increase slightly as ice accreted while the mean value of the thrust coefficients slightly increased as well. As has
been revealed above, the ice accretion in this test case is a typical rime icing process. The accreted ice profile at the
leading-edge had a streamlined shape. As the supercooled droplets continuously impinged on the blade, the ice layer
at the leading-edge was thickened, which effectively extended the areas of the blade lifting-surfaces so that more lift
force was generated. In the meantime, the aerodynamic drag acting on the propeller was also increased due to the ice
accretion on the rotating blades. The slight increase of the thrust coefficient was suggested to be a result of the addition
of the extra lift and drag forces generated during the ice accretion. As reported in previous works, ice accretion is a
unsteady (and sometimes random) process 6. Even for the blades with totally the same shape and dimensions, the
accreted ice mass and shape may vary in the different trials with the same icing conditions. The slight increase of the
force fluctuation amplitude observed here is suggested to be caused by the unsymmetrical ice accretion on the three
propeller blades due to the uncertainties in icing process.
Figure 9(b) shows the time history of the measured instantaneous thrust coefficient acting on the ice accreting
propeller at the warmer ambient temperature of T∞ = -5 °C while the other environmental parameters were kept the
same (i.e., U = 16 m/s; LWC = 1.0 g/m3). A similar behavior of the thrust coefficient loaded on the propeller was
observed, with the mean thrust value increasing slightly as the ice began to build up, while the force fluctuation
amplitude varied only in small scales as indicated by the almost unchanged width of the ± 1 std bounds during the
icing process. As mentioned above, typical glaze ice was accreted on the blade in this test case. Since the LWC was
relatively small, only the “needle-shaped” ice feature at the blade tip was observed in addition to the streamlined
leading-ice layer. The effect of such “needle-shaped” ice feature on the blade aerodynamic performance is small as
Dow
nloa
ded
by I
OW
A S
TA
TE
UN
IVE
RSI
TY
on
Febr
uary
13,
201
7 | h
ttp://
arc.
aiaa
.org
| D
OI:
10.
2514
/6.2
017-
0727
11
compared with the thickening leading-edge ice layer. Therefore, the lift and drag forces acting on the blades were
suggested to increase due to the expanding lifting-surfaces. The relatively stable force fluctuation indicated that the
amount of ice accretion on each propeller blade was fairly equivalent.
Figure. 9: Dynamic thrust coefficients measured at the rotational speed of n = 3000 rpm and U = 16 m/s, when
(a) T∞ = −15°C & LWC = 1.0 g/m3; (b) T∞ = −5°C & LWC = 1.0 g/m3; (c) T∞ = −5°C & LWC = 2.0 g/m3.
While the ambient temperature was still kept at T∞ = −5°C, as the LWC was increased to LWC = 2.0 g/m3, the ice
accretion on the rotating blade was typical glaze with evident surface water transport and irregular ice features accreted
on the propeller blades as described above. Figure 9(c) shows the time history of the measured instantaneous thrust
coefficient acting on the ice accreting propeller in such icing conditions. It is clearly seen that as the ice began to build
up on the propeller blades, there was a slight increase of the mean thrust value before t = 30 s, while the fluctuation
amplitude remained unchanged. As more water impinged on the blade (i.e., t > 30 s), the thrust acting on the propeller
started to decrease while the force fluctuation became smaller as well. Since there were more ice features of complex
structures growing during the icing process, the mass distribution of the accreted ice over the propeller blade was of
great difference compared to that of the aerodynamically-shaped ice accretion (i.e., in the ice accretion processes at
T∞ = -5 or 15 °C and LWC = 1.0 g/m3). Due to the combined effect of the aerodynamic shear force and the centrifugal
force, more water mass was transported to the outer section of the blades and froze into the “lobster-tail-liked” ice
features. Consequently, the structural vibration as the propeller rotates was damped due to the increasing rotating
Dow
nloa
ded
by I
OW
A S
TA
TE
UN
IVE
RSI
TY
on
Febr
uary
13,
201
7 | h
ttp://
arc.
aiaa
.org
| D
OI:
10.
2514
/6.2
017-
0727
12
mass, which explains the decrease of the force fluctuation amplitude. The decrease of the mean thrust force is
suggested to be caused by the complex ice structures that contaminated the aerodynamic blade profiles. As the ice
structures continued to grow outward, the centrifugal force acting on the ice structures became larger. Once the
centrifugal force overcame the adhesion force within the ice structures, a piece of ice was shedded as can be seen in
Fig. 9(c). The ice shedding caused a significant unbalance of the ice mass distribution on the three-blade propeller,
which resulted in a dramatic increase of the force fluctuation amplitude as can be seen at t = 51.7 s in Fig. 9(c). As ice
continuously accreted on the rotating blades, the unbalance effect diverged as indicated by the quick increase of the
force fluctuation amplitude. Meanwhile, the mean thrust value kept decreasing due to the growth of the irregular ice
features. As time goes on, a second ice shedding occurred at t = 102.5 s, which further unbalanced the rotating iced
propeller as indicated in Fig. 9(c). The larger fluctuations of the dynamic aerodynamic forces would indicate much
significant fatigue loads acting on the propeller when operated in the severe glaze icing conditions. The increased
fatigue loads are believed to be closely related to the growth of the irregular ice features towards the outer blade
section and the ice sheddings due to the centrifugal force. Such quantitative aerodynamic force measurement results
highlight the importance of taking the icing conditions into account for the safe operations of UAS in cold weather.
Figure. 10: Input power response measured at the rotational speed of n = 3000 rpm and U = 16 m/s under the
different icing conditions.
In addition to the aerodynamic force measurements, the input power required to drive the propeller at constant
rotation rate (i.e., n = 3000 rpm) was also measured in the present study. Figure 10 shows the input power response
as a function of time under the different icing conditions, which is represented by the ratio of input power with and
without ice accretion (i.e., Cpice/Cpno-ice). It is clearly seen that at the ambient temperature of T∞ = -15 °C and LWC =
1.0 g/m3, the required input power increased slightly as ice began to build up on the rotating propeller blades. As
mentioned above, the ice accretion around the blade leading-edge in such conditions is typical rime ice, which can
increase the aerodynamic drag acting on the propeller as well as the total rotating mass. Therefore, more input power
was needed to retain the rotational speed of the ice accreting propeller. A similar input power performance was
observed for the glaze ice accretion at the ambient temperature of T∞ = -5 °C and LWC = 1.0 g/m3 as shown in Fig.
10. Compare to the former rime icing process, slightly more input power (i.e., 10% more power) was required during
the glaze ice accretion. Since the accreted ice mass was equivalent, a larger aerodynamic drag was suggested in the
glaze ice accretion due to the surface water transport and roughness formation. For the glaze icing process with a
greater LWC (i.e., LWC = 2.0 g/m3), the input power increased exponentially as shown in Fig. 10. The ice shapes
formed in this test case were highly irregular with complex structures growing outwards along the blades, which
drastically contaminated the aerodynamic profiles of the propeller blades. The aerodynamic drag acting on the
propeller and the total rotating mass were significantly increased, and much more input power was required to keep
the propeller rotating at the same speed. Once the ice shedding occurred, a decrease or even a dramatic drop of the
input power would exhibit due to the sudden decrease of the rotating mass as also shown in Fig. 10.
Dow
nloa
ded
by I
OW
A S
TA
TE
UN
IVE
RSI
TY
on
Febr
uary
13,
201
7 | h
ttp://
arc.
aiaa
.org
| D
OI:
10.
2514
/6.2
017-
0727
13
V. Discussion
As described above, the ice profiles formed in rime icing, or glaze icing with a small LWC (i.e., ≤ 1.0 g/m3),
usually conform well with the blade leading-edge. Figure 11 shows a schematic of the ice accretion around the leading-
edge of a propeller blade in such conditions. As clearly shown in Fig. 11(a) and 11(b), the ice layer accreted around
the blade leading-edge has a rather streamlined shape that grows outward and into the oncoming flow. Such
streamlined ice layers could effectively increase the lift and drag forces acting on the rotating propeller blades, as
highlighted in the dynamic thrust and power measurement results. In the present study, more efforts were made to
quantitatively investigate the leading-edge ice accretion as a function of time and blade span locations. As mentioned
above, images of the transient ice accretion processes were acquired by using the “phase-locked” high-speed imaging
technique. Previous works [23] has validated that the leading-edge ice thickness can be extracted by digitizing the ice
features in the recorded images of ice accretion. A similar image-processing method was used in the present study to
examine the transient process of the leading-edge ice accretion.
Figure. 11: A schematic of the leading-edge ice accretion on a propeller blade. (a) Snapshot of a typical rime ice
accretion (with a cut) conforming with the blade profiles; (b) Zoom-in (view from side) of the ice accreted propeller
blade; (c) A schematic of the ice accretion at the leading-edge of a blade element.
Since the recorded images are maps of the light intensity scattered or reflected from the blade, water, and ice, by
comparing the intensity texture maps derived from the sequence of images before and during ice accretion, the
evolution of ice features can be extracted [23]. The initial reference image of the blade without water or ice was
defined as I0, and the ith image as ice accreted on the blade was defined as Ii. The intensity difference maps for the
images of the iced blade thus can be derived as:
0i i
diffI I I (8)
The pixel counts in the intensity difference maps represent the amount that the image has changed from the initial
state (i.e., blade without water or ice). Such difference in intensity is caused by the presence of water or ice (neglecting
image noise). Therefore, the advancing front of the leading-edge ice layer can be identified by finding the first location
in front of the blade with meaningful change in the pixel count compared to the initial reference; namely, for every
span position y,
2firsti i
ice diffx I x (9)
where ε was chosen as six standard deviations of the typical image noise. The image noise was characterized by
calculating the rms pixel fluctuations in the blank areas in the images (i.e., areas excluded blade, ice, or water) before
the ice accretion began.
Knowing the initial pixel locations of the blade leading-edge, i.e., 0
ix , the ice thickness at the leading-edge can be
calculated:
Dow
nloa
ded
by I
OW
A S
TA
TE
UN
IVE
RSI
TY
on
Febr
uary
13,
201
7 | h
ttp://
arc.
aiaa
.org
| D
OI:
10.
2514
/6.2
017-
0727
14
, 0
i i i
ice img iceH K x x (10)
where K is the calibration constant in mm/pixel.
In the present study, the propeller blade has twist angles along the blade span. The local tangential velocity also
varies due to the radius change in span. Therefore, the local resultant velocity as well as the local stagnation point may
vary at the different blade span locations. Since the maximum ice accretion occurs at the stagnation line [30], while
the stagnation point can be determined based on the local flow direction (effective angle of attack), a guideline of ice
accretion can be estimated. Figure 11(c) shows a schematic of the ice accretion at the leading-edge of a blade element,
in which the imaging path is from the top along the vertical direction. It can be seen that the ice thickness captured in
the image is a projection in x-y plane of the actual ice thickness. The relationship between the magnitudes of the
imaged ice thickness and the actual ice thickness is:
,
sin
ice img
ice
HH
(11)
where θ is the local twist angle, and α is the local effective angle of attack, which is defined as
2
arctanU
V
(12)
where V2 is the local tangential velocity, which is defined as V2 = r∙ω.
By applying the above methodology, the actual ice thickness accreted along the blade leading-edge can be
quantitatively extracted. Figure 12 shows the leading-edge ice growth as a function of time under the icing conditions
of U = 16 m/s, T∞ = −5°C and LWC = 1.0 g/m3. The snapshots of the leading-edge ice growth as a function time are
shown in Fig. 12(a). It is clearly seen that the leading-edge ice layer (i.e., advancing line indicated in red) is
continuously thickened as time goes on. The imaged ice layer thickness seems uniformly distributed along the blade
span. With consideration of the blade twist (θ) and the effective angle of attack (α), the distribution of the leading-
edge ice thickness along the blade span corresponding to the icing snapshots are given in Fig. 12(b). It is clearly seen
that there is an increase of the leading-edge ice thickness from the blade root to tip, which was also observed in
previous works of rotor icing studies [16,18]. Since the relative velocity at a local blade span is a function of the
rotation radius (i.e., with the minimum at the root and the maximum at the tip), there would be more water impinging
on the outer blade section within a specified duration of ice accumulation. In the case of glaze ice accretion, there is
also a surface water transport towards the blade tip due to the centrifugal force. For the test case here, in less than 2
minutes, the leading-edge ice growth at the blade tip exceeds 4.5 mm, while the ice thickness near the blade root (i.e.,
r/R = 0.2) is only 2.0 mm. Compared to the chord length (c) of the blade (i.e., croot = 11.6 mm and ctip = 3.4 mm), Such
leading-edge ice thickness represents a significant change in the airfoil shape, especially at the outer blade positions.
Additionally, the dotted lines in Fig. 12(b) indicate the 1 standard deviation of the roughness variations about the mean
ice thickness. As the leading-edge ice develops and grows, so does the roughness of the ice layer at the leading-edge,
which was also observed in previous works of the ice accretion measurements on a two-dimensional airfoil [23].
Figure 12: Leading-edge ice growth as a function of time at U = 16 m/s, T∞ = −5°C and LWC = 1.0 g/m3. (a)
Snapshots of leading-edge ice growth; (b) Distribution of the leading-edge ice thickness along the blade span.
Dow
nloa
ded
by I
OW
A S
TA
TE
UN
IVE
RSI
TY
on
Febr
uary
13,
201
7 | h
ttp://
arc.
aiaa
.org
| D
OI:
10.
2514
/6.2
017-
0727
15
Figure. 13: Time histories of the leading-edge ice thickness at three typical blade span locations (i.e., r/R = 0.3,
0.6, and 0.9) with the icing conditions of U = 16 m/s, T∞ = −5°C and LWC = 1.0 g/m3.
Figure 13 shows the time histories of the leading-edge ice thickness at three typical blade span locations (i.e., r/R
= 0.3, 0.6, and 0.9). It is clearly seen that the ice layer accreted at the outer blade section is much thicker. The linear
fitting lines of the measured ice thickness data are also plotted in Fig. 13, which indicates the time derivative of the
leading-edge ice accretion. The greater slope of a linear regression implies a faster growth of the leading-edge ice
layer. As clearly shown in Fig. 13, the ice growth rate at the outer span location (i.e., Kr/R=0.9 = 0.0347 mm/s) is almost
twice faster than that at the inner span location (i.e., Kr/R=0.3 = 0.0186 mm/s).
Figure. 14: A comparison of the span-variation of the leading-edge ice growth rate under various icing
conditions as the propeller rotates at a constant speed of n = 3000 rpm with the freestream velocity of U = 16 m/s.
By extracting the linear regression slopes of the temporally-dependent leading-ice thickness data along the whole
blade span, the span-variation of the leading-edge ice growth rate can be derived. A comparison of the span-variations
of the leading-edge ice growth rate under the various icing conditions is given in Fig. 14. It can be clearly seen that at
a constant cold ambient temperature (i.e., T∞ = -15 °C), the leading-edge ice growth rate increases proportionally as
the LWC increases from LWC = 0.5 to 2.0 g/m3. Since the ice accretion in such conditions is typical rime ice, the ice
accumulation parameter, a [26], is defined as
Dow
nloa
ded
by I
OW
A S
TA
TE
UN
IVE
RSI
TY
on
Febr
uary
13,
201
7 | h
ttp://
arc.
aiaa
.org
| D
OI:
10.
2514
/6.2
017-
0727
16
LWC U
aL
(13)
where τ is the duration of water and ice accumulation. The time derivative of the ice accumulation parameter relates
to the growth rate of the leading-edge ice layer.
LWCda U
d L
(14)
It is obvious that, as the freestream velocity, U, is kept constant, the rate of ice accumulation is proportional to the
LWC, which is reflected in the measured ice growth rate along the blade leading-edge as the LWC varies from LWC
= 0.5 to 2.0 g/m3 as shown in Fig. 14. It is also noticed that for the same level of LWC, varying the ambient temperature
(i.e., from T∞ = -15 to -5 °C) changes the span-variations of the leading-edge ice growth rate. At the warmer
temperature (i.e., T∞ = -5 °C), the rate of leading-edge ice accretion at the inner blade positions is slower than that at
the rime icing temperature (i.e., T∞ = -15 °C), while a faster ice accretion exhibits at the outer blade section as shown
in Fig. 13. Since the ice accretion at the warmer temperature is glaze ice growth, the surface water transport along the
blade span due to the centrifugal force is suggested to account for the changes of the leading-edge ice growth rate,
with less ice accreting at the inner blade span but more water is transported towards the blade tip and freezes at the
outer positions of the blade.
VI. Conclusions
In the present study, a comprehensive experimental study was conducted to investigate the transient ice accretion
process over the blade surfaces of a rotating UAS propeller. The experiments were performed in the Iowa State
University Icing Research Tunnel (i.e., ISU-IRT) with a scaled UAS propeller model operated under a variety of icing
conditions (i.e., ranged from rime to glaze ice accretion). A “phase-locked” high-speed imaging technique was
developed to resolve the transient details of ice accretion over the rotating propeller blades. Simultaneously, the
dynamic aerodynamic forces acting on the propeller and the power input for driving the constant rotation were also
measured quantitatively during the ice accretion processes. Based on such temporally-synchronized-and-resolved
measurements, the detailed ice accretion phases were correlated with the dynamic aerodynamic force data in order to
gain further insight into the underlying physics for the transient ice accretion processes on the rotating UAS propeller.
In the rime icing conditions, the ice layer formed at the blade leading-edge has a streamlined profile, which
effectively increases the lift and drag forces acting on the propeller blades, generating a slight increase in the resultant
thrust force, as well as the power input to keep the propeller at a constant rotational speed. A similar propeller
performance in thrust and power input is also observed in the glaze icing conditions with a small LWC. The leading-
edge ice layer formed in such moderate glaze icing conditions also has a streamlined profile, but with a “needle-shaped”
ice growing at the blade tip. Such “needle-shaped” ice feature develops into the “lobster-tail-liked” ice branches in
the severe glaze icing conditions (i.e., when the LWC is doubled), which is due to the obvious surface water transport
on the rotating propeller blades. These ice features dramatically contaminate the blade aerodynamic profiles as
indicated by the decrease of the thrust force acting on the propeller and the exponential increase of the power input.
When significant mass of ice is accumulated around the blade tip, sheddings of the ice pieces are observed, which
essentially increases the structural vibrations as indicated by the quick diverge of the force fluctuation amplitude. The
larger fluctuations of the dynamic aerodynamic forces would indicate much significant fatigue loads acting on the
propeller when operated in the severe glaze icing conditions.
With the “phase-locked” images of the transient ice accretion, efforts were also made to quantify the leading-edge
ice thickness as a function of time and blade span locations. It is found that the leading-edge ice thickness varies along
the blade span, with the maximum thickness located on the outer section and the minimum near the blade root. As the
leading-edge ice develops and grows, so does the roughness of the accreted ice layer. A linear increase of the leading-
edge ice thickness as a function of time is also observed, with a faster growth rate on the outer section of the propeller
blade. In both rime and glaze icing conditions, the ice growth rate at the blade leading-edge is proportional to the
LWC if the freestream velocity is kept constant. It is also found that in a glaze condition, the leading-edge ice layer
grows slower on the inner blade, but faster on the outer blade than that in a rime condition, which is due to the surface
water transport along the blade span driven by the centrifugal force.
Dow
nloa
ded
by I
OW
A S
TA
TE
UN
IVE
RSI
TY
on
Febr
uary
13,
201
7 | h
ttp://
arc.
aiaa
.org
| D
OI:
10.
2514
/6.2
017-
0727
17
Acknowledgments
The authors want to thank Dr. Rye Waldman, Mr. James Benson and Mr. Andrew Jordan of Iowa State University
for their help in operating ISU Icing Research Tunnel (ISU-IRT) Facility. The research work is partially supported by
Iowa Space Grant Consortium (ISGC) Base Program for Aircraft Icing Studies and National Science Foundation
(NSF) under award numbers of CBET- 1064196 and CBET-1435590.
References
[1] Bragg, M., Gregorek, G., and Lee, J., “Airfoil aerodynamics in icing conditions,” Journal of Aircraft, vol. 23, 1986, pp. 76–81.
[2] Broeren, A. P., Bragg, M. B., Addy, H. E., Lee, S., Moens, F., and Guffond, D., “Effect of High-Fidelity Ice-Accretion
Simulations on Full-Scale Airfoil Performance,” Journal of aircraft, vol. 47, 2010, pp. 240–254.
[3] Waldman, R., and Hu, H., “High-Speed Imaging to Quantify the Transient Ice Accretion Process on a NACA 0012
Airfoil,” 53rd AIAA Aerospace Sciences Meeting, 2015.
[4] Liu, Y., Waldman, R., and Hu, H., “An Experimental Investigation on the Unsteady Heat Transfer Process Over an Ice
Accreting NACA 0012 Airfoil,” 53rd AIAA Aerospace Sciences Meeting, American Institute of Aeronautics and Astronautics, 2015.
[5] Rothmayer, A., “On the Creation of Ice Surface Roughness by Interfacial Instabilities,” 41st Aerospace Sciences Meeting and Exhibit, American Institute of Aeronautics and Astronautics, 2003.
[6] Ueno, K., and Farzaneh, M., “Linear stability analysis of ice growth under supercooled water film driven by a laminar airflow,” Physics of Fluids, vol. 23, 2011, p. 42103.
[7] Fortin, G., Laforte, J.-L., and Ilinca, A., “Heat and mass transfer during ice accretion on aircraft wings with an improved
roughness model,” International Journal of Thermal Sciences, vol. 45, Jun. 2006, pp. 595–606.
[8] Wright, W., User’s Manual for LEWICE Version 3.2, Cleveland, OH: 2008.
[9] Potapczuk, M. G., “Aircraft Icing Research at NASA Glenn Research Center,” Journal of Aerospace Engineering, vol.
26, 2013, pp. 260–276.
[10] Campbell, S., Broeren, A., Bragg, M., and Miller, D., “Aircraft Performance Sensitivity to Icing Cloud Conditions,” 45th
AIAA Aerospace Sciences Meeting and Exhibit, Reston, Virigina: American Institute of Aeronautics and Astronautics, 2007.
[11] Szilder, K., and McIlwain, S., “In-flight icing of UAVs – the influence of flight speed coupled with chord size,” Canadian Aeronautics and Space Journal, vol. 58, Jul. 2012, pp. 83–94.
[12] Siquig, R., “Impact of icing on unmanned aerial vehicle (UAV) operations,” 1990.
[13] Bottyán, Z., “In-flight icing characteristics of unmanned aerial vehicles during special atmospheric condition over the carpathian-basin,” Acta Geographica Debrecina. Landscape &, 2013.
[14] Armanini, S. F., Polak, M., Gautrey, J. E., Lucas, A., and Whidborne, J. F., “Decision-making for unmanned aerial vehicle operation in icing conditions,” CEAS Aeronautical Journal, vol. 7, 2016, pp. 663–675.
[15] Sørensen, K. L., Helland, A. S., and Johansen, T. A., “Carbon nanomaterial-based wing temperature control system for in-flight anti-icing and de-icing of unmanned aerial vehicles,” 2015 IEEE Aerospace Conference, 2015, pp. 1–6.
[16] Lamraoui, F., Fortin, G., Benoit, R., Perron, J., and Masson, C., “Atmospheric icing impact on wind turbine production,”
Cold Regions Science and Technology, vol. 100, 2014, pp. 36–49.
[17] Palacios, J., Smith, E., Rose, J., and Royer, R., “Ultrasonic De-Icing of Wind-Tunnel Impact Icing,” Journal of Aircraft, vol. 48, May 2011, pp. 1020–1027.
[18] Fortin, G., and Perron, J., “Spinning Rotor Blade Tests in Icing Wind Tunnel,” 1st AIAA Atmospheric and Space Environments Conference, American Institute of Aeronautics and Astronautics, 2009.
[19] Han, Y., Palacios, J., and Schmitz, S., “Scaled ice accretion experiments on a rotating wind turbine blade,” Journal of Wind Engineering and Industrial Aerodynamics, vol. 109, 2012, pp. 55–67.
[20] Vargas, M., and Tsao, J., “Observations on the growth of roughness elements into icing feathers,” 45th AIAA aerospace sciences meeting and Exhibit, 2007, p. AIAA 2007-0900.
[21] Shin, J., and Bond, T. H., “Repeatability of ice shapes in the NASA Lewis Icing Research Tunnel,” Journal of Aircraft, vol. 31, Sep. 1994, pp. 1057–1063.
Dow
nloa
ded
by I
OW
A S
TA
TE
UN
IVE
RSI
TY
on
Febr
uary
13,
201
7 | h
ttp://
arc.
aiaa
.org
| D
OI:
10.
2514
/6.2
017-
0727
18
[22] Kraj, A. G., and Bibeau, E. L., “Phases of icing on wind turbine blades characterized by ice accumulation,” Renewable Energy, vol. 35, 2010, pp. 966–972.
[23] Waldman, R. M., and Hu, H., “High-Speed Imaging to Quantify Transient Ice Accretion Process over an Airfoil,” Journal of Aircraft, vol. 53, Aug. 2015, pp. 369–377.
[24] Liu, Y., and Hu, H., “An Experimental Investigation on the Convective Heat Transfer Process over an Ice Roughened
Airfoil,” 54th AIAA Aerospace Sciences Meeting, American Institute of Aeronautics and Astronautics, 2016.
[25] Brandt, J., and Selig, M., “Propeller Performance Data at Low Reynolds Numbers,” 49th AIAA Aerospace Sciences
Meeting including the New Horizons Forum and Aerospace Exposition, American Institute of Aeronautics and Astronautics, 2011.
[26] Anderson, D., “Acceptable tolerances for matching icing similarity parameters in scaling applications,” 39th Aerospace Sciences Meeting and Exhibit, American Institute of Aeronautics and Astronautics, 2001.
[27] Liu, Y., Chen, W.-L., Bond, L. J., and Hu, H., “An experimental study on the characteristics of wind-driven surface water film flows by using a multi-transducer ultrasonic pulse-echo technique,” Physics of Fluids, vol. 29, Jan. 2017, p. 12102.
[28] Rothmayer, A., “Scaling Laws for Water and Ice Layers on Airfoils,” 41st Aerospace Sciences Meeting and Exhibit, American Institute of Aeronautics and Astronautics, 2003.
[29] Hansman, R. J., and Kirby, M. S., “Comparison of wet and dry growth in artificial and flight icing conditions,” Journal of
Thermophysics and Heat Transfer, vol. 1, 1987, pp. 215–221.
[30] Özgen, S., and Canıbek, M., “Ice accretion simulation on multi-element airfoils using extended Messinger model,” Heat
and Mass Transfer, vol. 45, 2009, pp. 305–322.
Dow
nloa
ded
by I
OW
A S
TA
TE
UN
IVE
RSI
TY
on
Febr
uary
13,
201
7 | h
ttp://
arc.
aiaa
.org
| D
OI:
10.
2514
/6.2
017-
0727
