Introduction
Atmospheric icing affects the operational performance of many man-made
devices and ground structures. Examples include icing of electrical power
networks (), wind turbines (),
communication towers () and aircraft
().
Based upon the understanding of the physics and chemistry of clouds
(), several icing mechanisms are well-known to atmospheric
scientists and engineers: supercooled droplet freezing ()
– including freezing drizzle and rain (), snow accretion
() and mixed-phase icing ().
Particularly, in the field of aircraft icing, supercooled large droplet (SLD)
icing with median droplet diameters above 50 µm is considered as
an independent category ().
Supercooled means that the droplets are in a metastable liquid state although
their temperature is below the freezing point. Solidification can take place
through homogeneous or heterogeneous nucleation. In absence of any impurities
like aerosols, homogeneous nucleation starts with exceeding a
barrier of free enthalpy triggering the creation of critical embryos. The
change of free enthalpy is a function of latent heat of fusion (promoting
phase change) and energy needed to create a new water–solid interface
(hindering phase change). With increasing supercooling the latent heat of
fusion increases thus making homogeneous nucleation more probable. Droplet
impact on a solid substrate promotes heterogeneous nucleation
initiating a fast development of ice dendrites in the liquid
().
The following important parameters are of particular significance in defining the boundary conditions of the icing process:
Classical, fluid-mechanical testing parameters like Reynolds number Re=ϱairμairU∞c,
Mach number Ma=U∞a∞ and
angle of attack α, where ϱair is the density and μair the dynamic viscosity of air,
U∞ the free-stream velocity, c a reference length of the test model and a∞ the speed of sound. All kind of test models can
be considered for this dimensional analysis, e.g. aircraft, airfoils, probes, et cetera.
Water concentration in the air. On the one hand, this is the liquid water content (LWC), a measure of the mass of water per unit volume of air.
Typical values for atmospheric icing conditions range from 0.1 to 3 g m-3 and have been measured in the updraft of
continental convective systems e.g. above the Amazon basin; see and . On the other hand,
in the presence of ice crystals, the ice water content (IWC) has to be specified. Inside continental and oceanic mesoscale
convective systems, IWC values are about 1 g m-3 (). While ice water content in mid-latitude cirrus
is generally low (), peek values of 6 g m-3 have been observed in deep convective systems.
Commonly, total water content (TWC) is defined as the sum of LWC and IWC.
Median volume diameter (MVD) of the statistical distribution of water droplets in the liquid cloud. MVD attempts
to reduce the size distribution to a single, representative scalar diameter. This idea is attributed to Langmuir
(). In the presence of ice crystals, a median mass diameter (MMD) is defined because of their variable density.
Static air temperature for water droplet icing close to 0 ∘C promote glaze ice formations,
whereas rime ice formation becomes predominant at lower static air temperatures ().
Humidity, i.e. the amount of water vapour in the air, is an important parameter for mixed-phase ice accretion
(). showed its relevance for pure droplet icing. A distinction is made between absolute
humidity AH and relative humidity RH. Absolute humidity is defined by the mass of water vapour per unit volume of dry air.
With the partial pressure of water vapour pvapour, the static air temperature Tair and the specific
gas constant of the vapour Rvapour, this yields AH=pvapourRvapourTair.
Relative humidity is defined as the ratio of the partial pressure of water vapour pvapour to the equilibrium vapour
pressure of water pvapour*, i.e. RH=pvapourpvapour*.
Accumulation time tacc, for which the test model is exposed to the cloud of supercooled droplets and/or ice crystals.
Among the above enumeration of boundary conditions, temperature is of
particular importance, because it governs the freezing dynamics of impacting
supercooled water droplets. At very low temperatures, the droplets will
solidify shortly after their impingement and entrap the surrounding air. The
resulting ice accretion is called rime ice. With increasing temperature, the
solidification process of impacting droplets is retarded, yielding
wind-driven water film dynamics. At locations with increased convective heat
transfer, the water film freezes, yielding a nearly transparent – glaze ice
shape with typical horn formations.
Rime ice and glaze ice shapes over time at the leading edge of a
NACA0012 airfoil, chord length 0.53 m. One blue line represents one minute
of ice accretion. Further boundary conditions (rime):
U∞=58 m s-1, tacc=8 min; (glaze):
U∞=102.8 m s-1, tacc=7 min. Data from
computations with TAUICE by Jan Steiner.
The choice of boundary conditions for conducting a specific icing wind tunnel
experiment depends on the application under investigation. When testing to
support industrial product developments, certification requirements often
define the boundary conditions. In Europe, icing of large transport
aircraft is addressed in the document CS25 () of the
European Aviation Safety Agency (EASA). In its Appendix C, information on the
standard LWC–MVD envelope can be found, where the liquid water content is in
the range between 0.1 and 2.9 g m-3 and MVD between 15 and
50 µm. Icing certification for wind energy converters is
governed by the technical standards of the International Electrotechnical
Commission (IEC). Among others, the Det Norske Veritas Germanischer
Lloyd (DNV GL) is an accredited registrar and classification society
responsible for certification of wind energy converters, having created
calculation rules and detailed specification based on the IEC; see . It has to be noted that industrial products are in many
cases significantly larger than the dimension of a typical icing wind tunnel.
Hence, appropriate scaling laws are necessary to overcome the limited ranges
of air speed, test section dimension and icing cloud characteristics in icing
wind tunnels (see ).
Besides supercooled droplet icing, the aircraft industry has become aware of
another icing hazard in the recent past. The phenomenon called “ice
crystal icing” has been identified to effect ice accretion on heated aircraft
assembly such as engine compressor blades or stagnation pressure probes.
Thrust losses and engine damages as well as biased flight parameter display
and loss of the autopilot can be caused. Ice crystal icing mostly appears in
tropical regions in the vicinity of convective cloud systems. A comprehensive
treatment on the topic was published by . The
Federal Aviation Administration (FAA) of the United States and EASA extended
their icing regulations to Appendices D and P for ice crystal icing
conditions in 2010 and 2011. Enhanced research activities on the topic have
been promoted lately by NASA () and the National
Research Council of Canada ().
Icing is a very challenging field of study that incorporates aspects of
meteorology, fluid mechanics, thermodynamics, physics and engineering.
Despite the fact that many researchers have been involved in the study of ice
accretion over numerous decades, the physics of this phenomenon is still not
completely understood. Icing wind tunnels are therefore an essential pillar
to advance our knowledge in face of this multidisciplinary challenge.
Enduring changes in technical standards or certification processes to
continuously improve the safety against icing related incidents further
emphasize the industrial need for icing wind tunnel testing. The cost
estimations of the Ice Protection Harmonization Working Group (IPHWG) for
aircraft certification with respect to SLD-icing underlines this requirement; see Fig. .
Cost estimation for aircraft icing certification according to FAR
Part 25 Appendix O, based on the working group report on supercooled large
droplet rule-making .
Against this background, the design and construction of the Braunschweig
Icing Wind Tunnel began in 2010. It was a goal to contribute a tunnel with
sufficiently large dimensions in the test section to support both fundamental
and applied icing research with reasonably low operating costs. During the
design, construction and commissioning process, many lessons have been
learned, which can not be found in the literature and thus form the outline
of this publication. The major components of the Braunschweig Icing Wind
Tunnel and their design constraints are presented in Sect. 2. The special
topic of mixed phase icing, where both supercooled droplets and ice crystals
are involved in the icing process, is treated in Sect. 3. Together with the
international co-authoring partners, the commissioning of the tunnel was
realized, which is presented in Sect. 4. Finally, some applications and test
results that show the tunnel's capabilities are highlighted, with some
concluding remarks on health and safety considerations.
Design and construction of the Braunschweig Icing Wind Tunnel
Overall design
The aim to build a low-budget icing wind tunnel for research purposes
influenced major choices on the operational envelope. The main targets in icing
wind tunnel design are to obtain a homogeneous distribution of flow velocity,
temperature, and a uniform icing cloud in the test section. Therefore, it is
the test section that governs the design choices.
The first choice was on the dimension of the test section. Addressing
customers in aviation, automotive and energy industries requires reasonable
aerodynamic testing where the wind-tunnel wall boundary layers are
significantly smaller than the size of the test section. In this regard,
investigations with Reynolds numbers up to 2×106 are a frequent
demand. Moreover, the test section shall allow for mounting significant
aircraft subsystems and particle sizing instrumentation. Hence, a cross
section of 0.5 m × 0.5 m was considered as a lower bound for
sizing the icing wind tunnel. Additionally, it has to be taken into account
that the test section size is in relation with the overall dimension of the
icing wind tunnel. The space available at the installation site suddenly
becomes a further design constraint.
The maximum velocity is primarily determined by the power of the wind tunnel
drive; see Sect. . However, when considering the heat
balance in icing wind tunnels (see Sect. ) the
importance of the chilling device becomes evident.
Figure shows cost estimations for both drive
and cooling power, which is based on price lists of leading manufacturers in
their respective fields. Realizing that the cost for tunnel cooling is about
10 to 20 times more expensive than for the tunnel drive, the maximum velocity
in the test section is in direct relation to the available investment costs
for building the tunnel. Given an upper bound for the tunnel drive power, a
compromise between test section size and maximum velocity has to be made.
Consistent with the above requirements of test section size, the maximum
velocity was estimated to be 40 m s-1. A larger test section size
would decrease the maximum velocity given a constant drive power. Since many
aeronautical applications even demand for higher speeds, the decisions of a
maximum velocity of 40 m s-1 and a test section of
0.5 m × 0.5 m were taken.
Pricing as of 2017 for asynchronous motors from different
manufactures and for CO2 refrigerating plants before tax and
tolls.
The lower temperature limit inside the test section was set to
-20 ∘C. Below this bound, one usually observes rime ice
accretions on the test models. In contrast, glaze ice formation with
temperatures closer to 0 ∘C is significantly more complex from the
perspective of fundamental research that shall be conducted in the present
icing wind tunnel. An overview of the entire tunnel and its adjacent
components is given in Fig. . A
comparison with the specifications of selected other icing wind tunnels can
be found in Table .
Overview of the Braunschweig Icing Wind Tunnel and its adjacent
components.
Performance of different icing wind tunnels compared to the
Braunschweig Icing Wind Tunnel. U∞,max represents the
maximum air speed in the test section, Pchill is the installed
cooling power. Data extracted from , ,
and .
Tunnel name
Test section size
U∞,max
Pfan
Pchill
λ
PchillPfan
Altitude
m×m
m s-1
kW
kW
–
–
capability
NASA IRT
1.83 × 2.75
174.3
3730
–
0.20
–
no
CIRA
2.35 × 1.15
225.0
4000
6400
0.19
1.60
yes
Cox
0.71 × 1.17
98.3
149
–
0.27
–
no
NRC AIWT
0.57 × 0.57
100.0
450
420
1.99
0.93
yes
0.52 × 0.33
165.0
Braunschweig
0.50 × 0.50
40.0
37
80
3.35
2.16
no
Tunnel drive
Given an air temperature of -20 ∘C, a maximum speed U∞
and a cross-sectional area Atsec of the tunnel test section, the
required jet power Pjet can be calculated to the following:
Pjet=12ϱairU∞2⋅V˙air=12ϱairU∞2⋅U∞⋅Atsec=12ϱairU∞3⋅Atsec=12⋅1.38kgm-3⋅40ms-13⋅0.25m2≈11kW.
In wind-tunnel design, a power factor λ is introduced, which
indicates the ratio of fan power Pfan and jet power
Pjet. For a conventional, closed-loop, low-speed wind tunnel, a
value of λ≈1.5 can be assumed. The fan power has to
compensate pressure and skin friction losses of the individual tunnel
components. High losses are caused by the wall boundary layers of the test
section and the tunnel diffuser. Furthermore, the first and the second corner
after the diffuser create significant losses. Since the vanes of the first
corner are subject to ice accretion, their losses are tremendously high
compared to conventional closed-loop tunnels. Consequently, the power factor
was dimensioned to a large value of λ=2.3, yielding a fan power of
Pfan≈ 25 kW. An axial fan with an efficiency
ηfan≈67 % was selected for the tunnel drive, the
electrical power input is therefore 37 kW. Using a frequency converter,
the speed of the three-phase asynchronous motor of the fan drive can be
controlled, allowing variable tunnel speeds from 5 to 40 m s-1.
Chilling device
To adjust and maintain a constant temperature of the air inside the icing
wind tunnel, which is well below the freezing point of water, a chilling
device is necessary, whose cooling capacity has to exceed the power of the
wind tunnel drive.
Power estimation
There are three major heat sources that have to be compensated by the
chilling device: the power of the tunnel drive, the heat input through the
wind-tunnel walls and the heat transfer of the water spray. The latter is
composed of the sensible heat, Q˙sensible to
supercool the water spray from the temperature at which it leaves the
pneumatic atomizer (about 20 ∘C) down to the air temperature
inside the tunnel test-section and the latent heat
Q˙latent. As soon as the supercooled water droplet impacts
on a solid substrate, the drop solidifies due to the heterogeneous nucleation
and releases the latent heat. Both sensible and latent heat are a function of
the liquid water content:
m˙water=LWC⋅V˙air=LWC⋅U∞⋅Atsec,Q˙sensible=m˙water⋅cp,water⋅ΔT,Q˙latent=m˙water⋅clatent,water,
where m˙water is the water mass flow rate of the spray
atomizers and ΔT the temperature difference, with the values for
demineralized water as follows:
cp,water=4.183kJkg-1K-1,clatent,water=334kJkg-1.
For typical operational conditions of the icing wind tunnel, i.e.
U∞= 40 m s-1, ΔT= 30 K, and approximations
for the fan power and the heat input through the wind-tunnel walls, the
necessary cooling capacity can be estimated; see
Fig. . When LWC is in the range of
1 g m-3, a continuous cooling power of 30 kW is necessary. To
allow for a five minute peak load of LWC = 3 g m-3 and static
air temperatures down to -20 ∘C in the tunnel, the system was
dimensioned for a maximum cooling power of 80 kW.
Required chilling power decomposition as a function of the liquid
water content for typical operational conditions of the Braunschweig Icing
Wind Tunnel.
Refrigeration unit
To provide the icing wind tunnel with 30 kW continuous cooling capacity
and 80 kW maximum cooling capacity, a customized cooling unit was built
(see the upper right part of Fig. ). The core
of the system is a refrigeration unit, ➃, that cools the heat
transfer fluid (Therminol® D12) down to a
temperature of -32 ∘C. To compensate for the volume change of
the fluid, a surge tank, ➆, filled with nitrogen is installed. The
cold fluid is then stored in a 4 m3-sized buffer tank, ➄,
which acts as a hydraulic switch; on its right side, it is connected to the
refrigeration unit. On its left side, it is connected to the heat exchanger
in the wind tunnel. As a result, the operation of the refrigeration unit is
decoupled from the heat exchanger in the wind tunnel. Using a pump and
electrically operated valves (➅), the volume flow from the buffer
tank to the heat exchanger in the wind tunnel is controlled. Depending on
this flow rate, the continuous cooling power of 30 kW of the
refrigeration unit can be well exceeded. For a 80 kW peak load
operation, the complete buffer tank is first cooled to -32 ∘C,
and afterwards the chilled heat transfer fluid is pumped through the heat
exchanger at the wind tunnel within a short period of 7 min. The recovered
waste heat from the cooling process is fed into the regenerative heating
system of the institute building. Because of its cold-resistant properties,
the piping of the cooling system and the storage tanks are made of
stainless steel 1.4541, and then insulated with a 4 cm thick insulation
layer made of Kaiflex®. The entire system is
controlled by programmable logic control.
Heat exchanger
The heat transfer between the Therminol® D12
fluid and the circulating wind tunnel air takes place through two
consecutive heat exchangers. While flowing from the bottom to the top of
the first heat exchanger, the Therminol® D12
fluid increases in temperature. With only one heat exchanger, this would
yield an undesired vertical temperature gradient in the passing air flow. The
second compartment, in which the
Therminol® D12 fluid flows the opposite
way, thus compensates the vertical gradient in air temperature. The entire
assembly is installed in front of the third corner in the wind tunnel, see
also the illustration in Sect. . The large
distance to the test section promotes a homogenization of the temperature
distribution over the tunnel cross-section through turbulent effects.
Mechanically, the heat exchangers have a cross-sectional area of
1.6 m × 1.6 m and are composed of elliptical tubes made of
stainless steel and rectangular, smooth steel fins, which are firmly
connected to each other by hot galvanizing. The fin pitch is 3.5 mm. To
protect the galvanized steel surface against corrosion, it is coated with
ZACOSIN® 2000Q, an epoxy resin-based
protective coating that provides high thermal conductivity by embedded
aluminium particles; see also Sect. . Alternatively,
the heat exchanger could also have been made entirely of stainless steel.
However, compared to standard steel, stainless steel has a significantly
lower thermal conductivity (50 Wm-1K-1 compared to
15 Wm-1K-1), which is why the surface of the heat
exchanger would have to be much larger.
The heat exchangers also include a condensation drainage. Condensation
occurs predominantly in the initial cooling phase from room temperature to
cold operational temperatures. When continuously running at low temperatures,
the effect of condensation is only minor.
Settling chamber
To obtain a low turbulence level inside the test section of the tunnel, a
settling chamber with honeycomb and screens is installed. An influence on the
longitudinal (parallel to the main flow direction) and lateral flow
structures (perpendicular to the main flow direction) has to be distinguished
(). The honeycomb reduces both the lateral components
of mean velocity and of the larger turbulent eddies, whereas the screens
reduce the longitudinal components of turbulence or mean-velocity variations
across the sectional area. For the Braunschweig Icing Wind Tunnel, a
combination of screen–honeycomb–screen–screen was selected. The
screens have a mesh size of 6 mm, the honeycomb diameter is 10 mm,
its length is 100 mm.
Spray system
A spray system is required to produce a uniform drop distribution in the
tunnel test section. In the past five years, two different spray systems have
been developed; one system suited for large droplets
(MVD≈80 µm) and a high liquid water content
(1.1 g m-3 <LWC< 3 g m-3) and a second
spray system for low MVD (8 µm <MVD< 48 µm) and low LWC (0.1 g m-3 <LWC< 2 g m-3).
The spray system for large droplets is composed by a grid of 5×5
air-assisted atomizers. Inside one atomizer, a thin jet of water is
destabilized by the shear forces of a transversely directed air stream and
finally breaks down into small droplets. The diameter of the water jet, which
is in the range of 100 µm, determines the water volume flux as
well as the drop size. The higher the applied air pressure, the higher the
aerodynamic shear forces, and the smaller the drop size. In order to enable a
high variability of the drop size distribution for basic tests, the atomizers
are designed in a modular manner, that is, the water jet and the air atomizer
cap can be exchanged separately.
Stainless steel has been chosen as a material for the atomizers, since
otherwise conventional atomizers made of nickel-plated brass are prone to
wear due to frequent temperature changes inside the icing tunnel (from
+20 to -20 ∘C). However, the poor thermal conductivity of
stainless steel (15 Wm-1K-1) compared to brass
(120 Wm-1K-1) has to be considered. Since the spray
atomizers are directly exposed to the cold tunnel air, they would freeze
without corresponding countermeasures. Therefore, a heating coil is installed
close-by.
The high liquid water content (1.1 g m-3 <LWC< 3 g m-3) provided by the spray system is a technical constraint
of air-assisted atomizers. Below a certain water flow rate that is needed for
low LWC, their operation becomes unstable. This is because the low water
pressure inside the atomizer cap prevents the formation of a stable water
jet, which is needed prior to atomization. The second spray system
circumvents that problem and was specially designed for low liquid water
contents; see Fig. . To achieve a LWC of
0.1 g m-3, a water volume flux can be estimated to the following:
V˙water=LWC⋅V˙airϱwater=LWC⋅Atsec⋅U∞ϱwater=0.1gm-3⋅0.25m2⋅40ms-11000kgm-3=3.6Lh-1.
Using a spray matrix of 5×5 atomizers, this yields a volume flux of
0.14 Lh-1 per atomizer, which is ten times lower than the
typical minimum flow rate of air-assisted atomizers in industrial
environments. In consequence, electrically actuated air-assisted atomizers
have been chosen for the spray system. The atomizers are pulsating up to
10 000 times a minute, making the spray appear to be constant. By changing
the duty cycle, the LWC can be adjusted without modifying the pressure of
water and air supply, which is a major advantage compared to classical
air-assisted atomizers.
Spray system for low MVD and low LWC, and its supply system.
Water and air are conditioned for the operation of the air-assisted
atomizers; see again Fig. . A pressurized vessel of
stainless steel serves as a reservoir of water, whose temperature can be
regulated between 2 and 80 ∘C. Water pressure and air pressure
can be adjusted separately between 0.5 and 9 bar. Water flow rate and air
flow rate are measured and adjusted by flow regulators. Possible pollutant
particles, which could block the nozzles, are separated with water filters.
In larger icing test facilities, where tunnel air temperatures below
-40 ∘C can be adjusted, the air duct of the spray atomizers is
heated in order to prevent droplet freeze out. In our setup, droplet freeze
out is less likely due to the minimum temperatures of around
-20 ∘C, which are well above the homogeneous nucleation limit of
water. Therefore, the atomizer's air duct is not heated and air at room
temperature is used to operate the atomizers.
Droplet trajectories in the wind tunnel nozzle and their deviation
from the air streamlines under the influence of gravity. Large droplets may
collide with the tunnel walls thereby altering the droplet size
distribution.
Tunnel nozzle design
Having passed the settling chamber, the flow is accelerated in the tunnel
nozzle that contracts towards the dimensions of the test section. Beyond
altering the turbulence structure (), the shape of the
tunnel nozzle also influences the trajectories of the water droplets that are
injected by the spray system. Especially large droplets and ice particles
are affected.
Two different nozzle geometries have been considered. These are the
following:
the AVA nozzle, with its contour given by the following:zxztsec=K-11-xL32-1-xL3+1
the Witoszynski nozzle, with its contour given by the following:zxztsec=11+1K-11-xL221+13xL23,
where x,z are spatial coordinates, L is the nozzle length,
zsc,ztsec are half of the inlet and outlet diameter of
the nozzle and K is given by
K=zscztsec2. To provide enough
time for the supercooling process of the droplets on their way to the tunnel
test section, a large nozzle length of 3.5 m is foreseen.
The droplet trajectories are mainly governed by inertial forces, drag and
gravity. A water droplet with diameter d, velocity vector
udwd is transported by the air velocity
uairwair resulting in a drag force of the
following:
D=ϱair2Δu2d2cD⋅ΔuΔu,
where Δu=uairwair-udwd represents the
slip velocity. Assuming laminar flow around the spherical droplets with a
mass of md at low Reynolds numbers
Red=ϱair⋅Δu⋅dμair, the drag coefficient cD can
be approximated to cD=6πRed. Applying
the principle of d'Alembert, this yields a system of differential equations,
given by the following:
md∂udwd∂t=D+mdg,
where t represents the time and g is the acceleration due to
gravity.
Velocity and temperature slope for air and two different sized
droplets inside two tunnel nozzle configurations.
The effect of gravity on the droplet trajectories is depicted in
Fig. , assuming an air velocity of
40 m s-1 at the nozzle exit. Small droplets with a size of
60 µm follow the streamlines of air flow inside the nozzle.
With increasing droplet size, the gravitational deflection becomes
predominant, particularly at the low speed regions of the nozzle. For
300 µm droplets, their trajectory from the lowest spray bar even
collides with the nozzle wall. Such undesired collisions alter the size
distribution of the droplet cloud and have to be avoided.
Figure demonstrates the effect
of the nozzle geometry on the droplet slip velocity. Note that after the
nozzle length of L= 3.5 m, a straight segment of 0.5 m in length is
attached. In the low speed region x< 1.5 m, the AVA nozzle
accelerates the flow less than the Witoszinski nozzle. Consequently, the slip
in velocity between droplet and air is larger in that region. Downstream, for
x>3 m, the slip is caught up. While the small droplet of 60 µm approaches the air speed of 40 m s-1 when entering the test
section at x= 4 m, the large droplet of 300 µm still has
a slip of about 5 m s-1. Different initial droplet velocities
created by the air-assisted atomizers were considered likewise. However, we
observed that the surrounding low speed air flow in the settling chamber
rapidly decelerates the droplet. This effect on the droplet trajectory is
thus negligible in our tunnel.
Furthermore, the cooling process of the droplet on its trajectory has to be
considered. Cooling is promoted by a convective heat transfer at the
interface between droplet and air. When the droplet is smaller than
100 µm, the surface temperature penetrates the entire droplet
volume in about ten milliseconds, assuming a Fourier number of 1 for the
heat conduction problem; see also . Hence, isothermal
cooling can be assumed as an approximation for the thermal energy balance,
yielding a differential equation for the droplet temperature Td:
π6d3ϱdcp,d∂Td∂t=Nu⋅πd2Tair-Td⋅kaird,
where ϱd and cp,d represent the density and
the specific heat capacity of the droplet, kair is the heat
conductivity of air and Nu is the Nusselt number, estimated by
Nu=2+0.6Pr13Red12; see
. Herein, Pr=0.7 is the Prandtl number of air.
Using the above equation, the temperature evolution of the droplet inside the
tunnel nozzle is plotted in
Fig. for a given static air
temperature of Tair=263.15 K. After one metre travelling
distance, the droplet of 60 µm, which had an initial temperature
of Tair=293.15 K, has already reached the air temperature. In
contrast, the droplet of 300 µm has still a temperature above
the freezing point of water, when entering the test section at x= 4 m.
In summary, droplet conditions up to a maximum size of 150 µm
can be well provided in the present tunnel environment, taking into account
a reasonably low deflection due to gravity, a low slip velocity and
sufficient droplet cooling. Furthermore, the AVA geometry was selected for
the wind tunnel nozzle, because it enables a slightly higher heat transfer to
supercool the droplets.
Tunnel walls
Mechanical design
Mechanical design of the icing wind tunnel walls.
To minimize the heat transport through the wind tunnel walls, they consist of
three different layers with a total thickness of 80 mm; see
Fig. . Between two layers of aluminium, an
insulating foam with low heat conductivity is introduced. The inner aluminium
shell which is potentially exposed to the icing cloud is anodized for a
protection against corrosion. The flange of each wall segment is made of
purenit®, a highly compressed material based
on PUR/PIR rigid foam providing a high thermal insulation. A rubber seal
prevents the water from penetrating into the wall. The flanges were fastened
with many screws, thus avoiding water leakage. In order to make the joints
fully watertight, the inner groove was sealed with an acrylic compound. Due
to an excellent manufacturing accuracy, no steps between adjacent wall
segments are observed.
Aerodynamic wall effects
Icing wind tunnel tests usually involve models with large dimensions to
minimize uncertainties when applying scaling laws for ice accretion. The flow
around these models thus interferes with the walls of the test section. Two
different wall effects have to be considered. On the one hand this is the
viscous effect of junction flow (), which occurs when
the boundary layer of the wind tunnel wall encounters the test model attached
to the same wall. The adverse pressure gradient forces the wall boundary
layer to separate with horseshoe vortices that wrap around the obstacle.
These unsteady vortices result in high turbulence intensities, high surface
pressure fluctuations and heat transfer rates, which finally alter the ice
accretion in that area. This effect was computationally studied for a
NACA0012 airfoil with the icing code FENSAP (); see
Fig. . FENSAP solves the Reynolds-averaged
Navier–Stokes equations using a finite element approach, calculates droplet
trajectories with an Eulerian scheme and finally determines the ice accretion
using a model of . The upper part of
Fig. clearly identifies the horseshoe vortex while
plotting the spanwise velocity in multiple slices and an isosurface along the
NACA0012 airfoil. Especially in the vicinity of the leading edge at the
junction between airfoil and wind-tunnel wall, the two-dimensional flow
behaviour is significantly deteriorated. Because a glaze ice case was
simulated (T∞= -10 ∘C, LWC = 1 g m-3,
ice accumulation time tacc=120 s), the velocity field
strongly influences the heat transfer. At the junction, the decreased heat
transfer attenuates the ice accretion, which is shown in the lower part of
Fig. . However, the two-dimensional ice shape
distant to the tunnel wall, which covers 80 % of the wing span, remains
unaffected.
The second case of wind-tunnel wall interaction is an inviscid effect. The
distant walls that are not connected to the test model will influence the
streamlines around the model. Thereby, both pressure and shear stress
distribution along the model surface are altered. Interestingly, at low
speeds and sufficiently large droplet sizes (MVD = 15 µm), the
ice accretion is nearly unaffected, when tunnel walls are introduced; see the
computational results of the icing code TAUICE in
Fig. . TAUICE is developed by the German
Aerospace Center. It solves the Reynolds-averaged Navier–Stokes equations
using a finite volume approach, calculates droplet trajectories with a
Lagrangian scheme and finally determines the ice accretion using a model
of .
Corrosion
Since the operation of the icing wind tunnel entails the usage of water
droplets or ice particles, corrosion has to be considered during the design
process. The most vulnerable part of the tunnel in this regard is the heat
exchanger, as it is made from a mix of steels, which is galvanized
together; see Sect. . Without any proper
protection, many factors decide on the rate of corrosion of a galvanized
surface, the most important for this application being the chemical
composition of the water and the type of surface contact.
Water with a high content of carbonates supports the creation of a protective
layer of alkaline zinc carbonates ZnCO3 on the galvanized surface.
However, with no or only few carbonates and a relatively high content of
oxygen, the water reacts with the zinc to form zinc hydroxide 2Zn+2H2O+O2=2ZnOH2. This
further forms a compound 2ZnCO3⋅3ZnOH2, also known as white rust,
which does not have the protective function of zinc carbonates, but only
connects loosely to the surface. Similar to normal rust this porous layer
keeps water near the surface and delays drying, which further increases the
rate of corrosion. Over time, particles flake off, until the zinc is used up.
The type of surface contact between water and the galvanized surface is
important for the rate of corrosion. If the water is sprayed onto the surface
and remains as small droplets, each droplet forms a small corrosion element
with a large surface that promotes the supply of oxygen. This increases the
corrosion rate in comparison to the case where a galvanized surface is
completely submerged in water with only a small contact area to the
surrounding air.
To protect the galvanized steel surface of the heat exchanger against
corrosion, it is coated with ZACOSIN® 2000Q,
an epoxy resin-based protective coating that provides high thermal
conductivity by embedded aluminium particles.
Another metal in contact with the deionized water is brass, which is used in
several valves, pipe connections, etc. Brass forms tarnish, a thin dark layer
of stable metal oxides, that protects the base metal. Unlike rust, the
oxidation reaction is self limiting. As soon as the layer is formed, no more
metal is oxidized. The tarnish did not pose any problem in all parts.
The models for the tunnel are made from aluminium or glass-fibre reinforced
plastics (GRP). None of these materials showed any signs of reaction with the
water so far. The same holds for the inner wall of the wind tunnel which is
made of anodized aluminium. Even with scratches from installing models no
signs of corrosion showed up.
Basic instrumentation
To monitor the icing wind tunnel operation and thus the boundary conditions
of an experiment, several probes are necessary; see
Fig. . The velocity U∞ in the test
section is determined by two static pressure probes psc and
ptsec. Their typical accuracy for long term stability is
±0.3 % of the full scale output. Using the Bernoulli equation:
psc+ϱair2usc2=ptsec+ϱair2U∞2+ζnozzleϱair2U∞2,
where usc is the velocity in the settling chamber at
psc, ϱair is the density of air,
ζnozzle is the specific loss coefficient of the wind tunnel
nozzle, and a continuity equation as follows:
usc⋅Asc=U∞⋅Atsec,
where Asc and Atsec are the cross sectional areas of
settling chamber and test section, one can estimate the following:
U∞=ptsec-pscϱair2AtsecAsc2-1-ζnozzle.
Note that the effect of pressure losses in the wind tunnel nozzle is usually
neglected, since AtsecAsc2≫1
and ζnozzle≪1.
A Vaisala HMT-337 monitors both the relative humidity RH and the
total temperature Ttot in the settling chamber. The uncertainty
for the humidity measurement is ±1.8 % RH, the accuracy for the
temperature measurement is around ±0.3 ∘C. Ttot is
used to calculate the static air temperature T∞ in the test section,
T∞=Ttot-U∞22cp,air,
where cp,air is the specific heat capacity of air. Moreover,
the temperature information is necessary to determine the density
ϱair of air in Eq. (). The power of the
tunnel drive Pfan and the pressure loss over the heat exchanger
ΔpHE indicate the performance degradation of the icing wind
tunnel due to ice accretion. Furthermore, the state of the spray system
Sspray is controlled; see Sect. .
The junction flow between a NACA0012 airfoil (c=0.75 m,
α=0∘) and wind-tunnel side wall creates a horseshoe vortex
(upper image), which is altering the ice accretion (lower image).
U∞=40 m s-1, T∞=-10 ∘C,
LWC = 1 g m-3 and IWC = 0.
Inviscid wall effect on ice accretion for a AH-94-W-145 airfoil, c=0.75 m, U∞=40 m s-1, T∞=-10 ∘C,
LWC = 0.3 g m-3, MVD = 15 µm and IWC = 0.
Ice crystals and mixed phase capability
Basic instrumentation of the Braunschweig Icing Wind Tunnel.
Ice crystals in the atmosphere have to be considered for aircraft safety,
since they partially melt in warm environments and develop a “sticky”
character. In particular, heated stagnation pressure probes and engine
compressor stator blades can be maleffected by ice crystal icing
(). Icing of aircraft probes can cause false
flight parameters displayed inside the cockpit. Ice accretion inside the
compressor causes flow blockage, forcing the compressor to operate towards
stall conditions. The compressor encounters a decay in rotational speed
resulting in significant thrust losses (rollback event)
(). Moreover, total engine flame
out may appear if a huge mass of accumulated ice is shed into the combustor.
To provide experimental capability on ice crystal icing, the Braunschweig
Icing Wind Tunnel was upgraded with an ice crystal generation and conveyance
system, which is presented in this section; see also
Fig. .
Morphology of ice crystals
Ice crystal icing conditions are typically encountered in wide ranging
convective cloud systems of high ice water content at flight altitude. Such
conditions can especially be found in the vicinity of mesoscale convective
cloud systems in tropical regions; see
and . In order to better document ice crystal
icing conditions, two flight campaigns have been conducted in the course of
the HAIC and HIWC project
(). The first campaign took
place in Darwin, Australia, in 2014 during the monsoon period, the second
campaign in Cayenne, French Guiana, in 2015 during the rainy season. The
flight measurements have been conducted in high ice water content cloud areas
in large tropical mesoscale convective systems, mostly over the oceans.
Details about the campaigns and about data treatment can be found
in , and .
Ice particle images captured by 2D-S probe during the Darwin
Campaign. (a) Stratiform cloud region, MMD
≈ 150 µm, IWC ≈ 1.2 g m-3,
T∞≈ -10 ∘C. (b) Convective cloud
region, MMD ≈ 80 µm, IWC ≈ 3 g m-3
and
T∞≈ -10 ∘C.
Atmospheric ice particles feature a broad diversity of sizes and shapes
which depend on the individual ice particle growth history affected by
ambient temperature and super-saturation. Figure (a)
shows examples of ice particle images captured close to -10 ∘C
during the Darwin campaign with the 2-D-Stereo probe (see
Sect. for a description of the instrument). Images from
Fig. correspond to a stratiform part of the cloud
where rather constant ice water content close to 1.0 g m-3 was
sampled. Images from Fig. (b) were recorded in
convective cores with IWC peak values exceeding 3 g m-3. In the
convective part, ice particles are more numerous and smaller. Close to
-10 ∘C, column and capped column type crystals have been found.
Larger ice particles (> 600 µm) are rare and resemble
graupel (dense and roundish particles). On the contrary, in the stratiform
part of the cloud, ice particles larger than 600 µm are more
frequent and appear to be less fragile, consisting of aggregates of pristine
shapes. Since particle growth is affected by vapour deposition and
aggregation and encounters different temperature regimes, a lot of different
and irregular shapes are possible at all altitudes in mesoscale convective
cloud systems.
Ice particle size distribution (MMD) based on HAIC Falcon 20 flight
measurements from Darwin and Cayenne in dependence of ambient air
temperature.
Figure shows ice particle size and mass
distributions (PSD and MSDs) depending on ambient temperature using the
equivalent diameter deq for size definition. The equivalent
diameter intends to provide a size information on a non-spherical particle.
It is defined as the diameter of a circle of the same area as the shaded
pixel number for each particle. With the distribution of deq it is
therefore possible to compare the composition of icing clouds in both
atmosphere and icing wind tunnel environment. The size and mass distributions
have been averaged for the selected temperature regimes, only cloud areas
with total water content above 1.0 g m-3 have been taken into
account. The concentrations of small (< 200 µm) and large
(> 1 mm) ice particles vary in opposite ways with temperature: for
colder temperatures, concentrations of small ice particles increase, whereas
the number of large particles decreases (cf.
Fig. ). This temperature dependency might be a
consequence of several cloud processes. Nucleation of new ice particles is
favoured at low temperatures, creating new small ice crystals. On the
opposite, growing of ice crystals by collection processes requires larger
particle sizes and might be more efficient at higher temperature. Regarding
the dynamics of the cloud, small ice particles can also easily be carried
aloft by updraft winds. Therefore, there is still no clear and unique
scenario describing the formation of high IWC areas in clouds. Cumulative ice
particle mass distributions are plotted in the bottom part of
Fig. . PSDs have been converted in to MSDs,
following the work of and
. Bin masses are linked to the particle size using a
power-law relationship m(d)=βdeqγ. However, β
and γ are not constant; for each time step, γ is deduced from
the analysis of particle images and β is constrained by additional TWC
measurements, thus ensuring that the total mass from the MSDs equals the
measured TWC. With decreasing atmospheric temperature, the cumulative particle
mass is carried more and more by small ice particles. The median mass
diameter (MMD), reduces from roughly 750 µm at
-10 ∘C to 320 µm at -50 ∘C. More
details about MMDs in high IWC cloud regions can be found in
.
As mentioned above, ice crystal icing can cause malfunctions of aircraft
engines due to inner ice accretions. Inside-engine conditions are
characterized by ice particle sizes of about 20 µm as ice
particles fragment in the fan stage. Due to centrifugal forces, high ice
particle concentrations can be found in the casing region of the core engine.
The local TWC can exceed the atmospheric TWC by a factor of 4 or even more; see .
Ice crystal production in a cloud chamber
Ice crystal generators. A jet of air (a) circulates the
flow, allowing for ice crystal growth. When the crystals are large enough,
they fall into a chest freezer (b). Microscopic images of ice
particles inside the cloud chambers (c). Both plate crystals and
needles can be generated.
It has been aimed to reproduce closest possible replicates of natural ice
crystals. Cloud chamber technology has been identified to be appropriate,
because natural ice crystal growth is simulated in an artificial cloud.
Usually, cloud chamber technology is applied in meteorological science to
investigate ice crystal formation and growth mechanisms and to study the
interaction of individual ice crystals; see . Thus, the
productivity of cloud chambers is accepted to be rather low. For icing wind
tunnel studies, huge amounts of ice particles are required. In collaboration
with the Austrian Neuschnee GmbH, two highly productive cloud chambers have
been developed and installed inside a cooling room.
Figures and
show both cloud chambers and auxiliary equipment. Basically, atomized
droplets are forced to freeze out inside the chamber by the expansion of cool
pressurized air. Strong circulation of air keeps the particles in suspense
until they grow to certain size and settle down to the bottom of the chamber.
The production rate of the cloud chamber is not sufficient to enable a direct
supply of ice crystals into the wind tunnel. Thus, the particles are
collected inside a chest freezer, which is directly connected to the cloud
chambers. The chest freezers operate at -70 ∘C to minimize
degeneration and sintering of the particles. Currently, the production rate
for both cloud chambers is limited to about 1 kgh-1. This
production rate is sufficient to perform between 10 to 15 wind tunnel tests a
day.
Microscopic images of ice crystals grown inside the cloud chamber are shown
in the upper right part of Fig. . The primary ice
particle habit can be adjusted by variation of ambient temperature inside the
cooling chamber. After storing in the chest freezer, the ice particles
typically feature aggregates of individual crystals as further illustrated in
the following section.
Ice crystal conveyance
To establish defined cloud conditions inside the test section of the icing
wind tunnel, ice particles are fluidized into an airflow and guided into the
wind tunnel; see Fig. . The conveying
airflow is extracted from the icing wind tunnel by a bypass construction
including a radial fan. An external heat exchanger (aftercooler) cools the
air to compensate heat input of the fan and environment. The aftercooler is
connected to the refrigeration system of the cooling chamber. Downstream of
the heat exchanger, the piping system enters the cooling room where ice
particles are supplied to the airflow. Further downstream, the particle-laden
flow is discharged into the icing wind tunnel.
Assembly for ice particle dosing and fluidization. Ice particle
structure after storage, dosing and sieving procedure.
To adjust the ice water content inside the test section a defined mass-flow of
ice particles m˙ice,inj has to be supplied at a constant
feed rate. Particle dosing is realized by a volumetric dosing machine, shown
in Fig. . The dosing machine has a very linear
operating behaviour: the higher the frequency of the dosing machine motor,
the higher the supplied ice particle massflow; see
Fig. a. After storage and dosing, the
particles are partially interlocked and exist as accumulated ice clumps,
which are by far larger than the desired particle sizes of several tens of
microns. Thus, the particles have to be sieved to desired sizes, which is
realized by a custom made sieving machine; see again
Fig. . The sieving machine is driven by an
electric motor that forces a 800 mm × 800 mm sieve to
oscillate at a frequency of 8 Hz. The total mass of ice provided by the
dosing machine does not pass the sieving machine. Greater clumps of ice are
not fully broken by the sieving procedure and depose on the sieve. The
average deposit is about 15 to 20 % of the ice mass provided by the sieving
machine.
Ice particle massflow, (a) dependence on dosing machine
frequency and (b) ice particle supply over time at a dosing frequency
of fdosing=21Hz.
The sieving machine is connected to the conveyance piping system by a conical
flexible bag. Ice particles passing the sieve directly fall into the piping
system and get dragged by the pipe airflow. To adapt the local pressure
inside the piping system to the ambient pressure inside the cooling room, the
piping system implies an injector nozzle that induces a local jet inside the
pipe; see the lower part of Fig. . The nozzle
diameter is carefully adjusted based on preliminary calculations of the
pressure distribution inside the whole piping system. The temporal stability
of the mass flow is depicted in Fig. b.
The dashed line shows the ice particle mass supplied by the dosing machine at
a frequency of 21 Hz. It is a linear fit, a constant feed-rate has been
proven by weighing, as indicated by the red data points. The blue dots and
the solid blue line represents masses of ice weighed at the outlet of the
sieving machine. This curve shows a very linear behaviour as well. One can
conclude that after about 30 s the ice particle massflow injected into the
wind tunnel is temporally constant. Weighing of ice mass passed by the
sieving machine has been repeated several times, also for lower feed-rates and
it turned out that after about 30 s the provided ice massflow is stable.
Correlations between the injected ice particle massflow
m˙ice,inj and IWC inside the test section are discussed in
Sect. . The lower left part of
Fig. shows a microscopic image of two ice
particles which have been captured at the exit of the sieving machine. As one
can see, the particles feature aggregates of tiny ice particles and can
thus be considered as close replicates of the natural ice particles shown
in Fig. .
Isokinetic probe to measure the total water content, TWC, developed at Cranfield University.
The ice particles are injected into the settling chamber of the icing wind
tunnel, upstream of the spraybar system. The velocity of the particle-laden
jet is about 20 m s-1, which is five times higher than the local wind
tunnel airspeed. Consequently, the jet mixes and expands by turbulent mixing
with the ambient airflow so that the particles are spread among the flow
field and mix with the droplets atomized by the spraybar system. The particle
trajectories get contracted inside the wind tunnel nozzle. Due to the
circular pipe exit the ice particles also cover a circular cross sectional
area inside the test section. It has been tried to extend this area by using
adapters at the pipe exit and by the use of several injection pipes
distributed along the settling chamber. These efforts did not work out due to
ice deposits inside the additional assembly; see .
Consequently, the simplest approach of the single pipe outlet has been
chosen. For studies at glaciated and mixed phase conditions, screens and
honeycombs in the settling chamber of the tunnel have to be unmounted to
avoid ice accretion on these elements.
Commissioning of the Braunschweig Icing Tunnel
General approach
The icing wind tunnel calibration has been performed with respect to the
requirements specified in SAE ARP 5905. Accordingly, both an aero-thermal
calibration of the airflow and a calibration of the icing cloud had to be
performed. After presenting the deployed measurement techniques, selected
calibration results are presented. Finally, an instrumentation
inter-comparison exercise was carried out.
PDI probe
The measurement principle of Phase Doppler interferometry (PDI) is based on
light scattering interferometry () that uses as
measurement scale the wavelength of light and as such its calibration is not
as easily degraded. The scattering by spheres much larger than the instrument
wavelength is approximated by geometrical optics. Size and velocity are
determined by measuring sinusoidal scattering signals on adjacent
photo-detectors as particles move through an interference fringe pattern
formed in the intersection of two laser beams of same wavelength.
The method does not require frequent calibration as the light wavelength
does not change and the detector separations that affect the size measurement
are fixed. The sinusoidal nature of the signals detected may be used with the
Fourier analysis approach to detect signals reliably even in low
signal-to-noise (SNR) environments. Off-axis light scatter detection is used.
The advantage of this approach is that a very small, well-defined sample
volume can be formed which minimizes uncertainty when computing absolute
concentration and reduces the possibility of coincident events (more than one
particle residing in the sample volume at one time). The sample volume can be
adjusted to balance count rate with coincidence rate to suit the user's
preference. The PDI only needs an initial factory calibration since
parameters affecting the measurements as the laser wavelength, beam
intersection angle, transmitter and receiver focal lengths, and the detector
separation do not change with age or use of the instrument.
The method has been demonstrated to be capable of measuring drop size
distributions, droplet velocity distributions, size–velocity correlation,
droplet time of arrival, droplet spacing, droplet number density, liquid
volume flux and liquid water content. At the Braunschweig Icing Wind
Tunnel, the PDI has been used for size and concentration measurements of cold
clouds and for liquid or ice discrimination in mixed phase conditions. Although
the PDI method cannot provide a size measurement for ice particles, it is
able to extract a velocity information due to the particle's reflectivity.
The resulting velocity histogram was used in an instrumentation
inter-comparison campaign (see Sect. ) in order
to check differences between the mean velocity of ice particles and airspeed
upstream of the 2D-S probe volume.
Cranfield isokinetic probe (IKP)
The probe consists of two tubes with one pipe faced towards the
upstream direction, thereby collecting a representative sample of air,
droplets and ice particles suspended within the flow; see
Fig. . The goal is to measure the total water content
(TWC) of the flow. The volume flux Q˙front has to be adjusted
to obtain isokineticity at the probe head. The second tube is faced
downstream with its pipe end. Thus, a second sample of air, without any
particles or droplets in it, can be collected to determine the water vapour
content so that the quantity of water in condensed form (solid and liquid)
can be determined.
A high electric current is passing through the metallic structure of the
probe head thus enabling its resistive heating. Consequently, the probe
stays free of an ice build-up so that the flow conditions at the entrance to
the probe can be carefully matched to oncoming flow in order to achieve the
correct sampling, not too rich or depleted with respect to the condensed
water. The heating of the inside of the probe drives all sampled water
droplets and ice particles into the vapour phase. A subsequent cooling
system composed of copper tubes plunged into a water tank has been added in
order to bring the air at a reasonable temperature (about
30–35 ∘C) before entering the measurement system.
The measurement system itself is mainly comprised of two parts. The first
part is the mass flow meter measurement and valve to automatically control
the mass flow at the probe inlet to maintain iso-kinetic conditions. The
second part is the water vapour concentration measurement which is done using
a LICOR 7000 system.
The probe features a double wall construction making it possible to use it as
a pitot static probe and also providing a means to heat it. A large current
is passed through the outer wall, via the tip and back along the inner tube.
Tube wall thicknesses are chosen to get the appropriate split of heat between
the different parts of the probe.
HSI probe
The high speed imaging (HSI) probe () illuminates the
volume using six laser beams and takes shadow images of traversing cloud
particles. Thereby, depth of field errors and sampling bias due to particle
obscuration are minimized. The shadow images are sampled at a frequency of
300 fps by a CMOS-chip of 640 by 480 pixels. Images of particles in the
range of 5 to 1200 µm can be recorded, the resolution is
3.795 µmpixel-1. The entire device including lasers and
camera is remotely controlled. A trigger beam on a different wavelength is
coaxially aligned to the laser beams. A receiver, including an aperture of
appropriate size and shape, is elevated to 40∘ to the transmitted beam
and detects particles passing the object plane. It will trigger the lasers
(laser pulse duration 12 ns) when particles are within the desired
measurement volume. The HSI hardware is integrated in a cylindrical canister.
Three arms, including optical components, are mounted on the front end of the
canister that allow for the taking of intrusive particle images. The image processing
software is able to detect and to evaluate irregular shaped particles, and
includes several particle validation criteria, e.g. out-of-focus detection.
Spatial distribution of axial flow velocity U∞,x, (a) ice-crystal-icing setup and (b) standard setup of the icing wind tunnel.
2D-S probe
The two-dimensional stereo (2D-S) probe
(, ) detects the size and
concentration of cloud droplets and ice particles in the size range of 10 to
1280 µm using shadow images of the cloud particles. Two
orthogonal diode laser beams illuminate two linear diode arrays consisting of
128 photodiodes with 10 µmpixel-1 resolution. When a
particle crosses the laser beam in the sampling volume, its shadow image on
the photodiode array is recorded by high-speed electronics. The 2D-S probe
used during the airborne and wind tunnel measurements was equipped with
anti-shattering tips to reduce shattering of large particles on the leading
edge of the probe arms. The diode lasers operate at 45 W and are single-mode
and temperature-stabilized. This design with two lasers better defines the
sampling volume boundaries and thus minimizes errors associated with the
depths of field and the sizing of small particles.
The LaMP 2D-S data processing has been described in detail in
, as it has previously been used for the treatment of the
data sets collected during the HAIC field campaigns. The software is able to
extract various size parameters from the particle images (area equivalent
diameter, maximum dimension, etc.) and treats all major artefacts related to
optical array probe's measurements, e.g. shattering, splashing, multiple
particles in a single image, out of focus images. An area equivalent diameter
was used in the present manuscript. For truncated images, the computation of
the particle's size is based on the work of : all valid
images are considered for the retrieval of the particle size distribution.
The size of out-of-focus particles is corrected according to the work of
. The only difference between the analysis of the field
campaign and the wind tunnel datasets lies in the shattering treatment as
this artefact removal feature has been turned off for the wind tunnel
measurements. While most shattering should be minimized by the
anti-shattering tips, the shattering treatment in regions of large effective
diameters and high IWC further corrects for shattering artefacts
(). Shattering of large particles is a function of IWC; see . Since hardly any particles were larger than the full
array width (see the results in Sect. ) we
argue that large particles that would have caused shattering artefacts are
not present in the icing wind tunnel. Therefore, the shattering treatment was
turned off.
Calibration results
Aerothermal calibration
An aerothermal calibration has been conducted to ensure and prove adequate
airflow quality inside the test section. Investigations on airflow uniformity
have been made by means of a five hole probe. The probe has been calibrated
a priori based on the procedure described by . The
measurements have been performed for the standard setup of the icing wind
tunnel with screens and honeycombs installed in the settling chamber and also
for the ice-crystal-icing setup, where these elements have been unmounted.
Figure a shows the contour of axial velocity inside the
test section for the ice-crystal-icing setup. The pneumatic system for
ice-particle conveyance was activated, the exit velocity of the conveyance
pipe inside the settling chamber was about 20 m s-1, however, no ice
particles were fed into the system. The plot covers a cross-sectional area of
240 mm × 400 mm around the centre-line position.
Measurements have been taken in vertical and horizontal steps of 40 mm.
One can clearly observe a footprint of the pneumatic-conveyance jet inside
the flow field. Velocity deviations relative to the centre-line velocity are
about 0.4 m s-1, which complies with ARP 5905 requirements of
±1 % spatial flow uniformity. Moreover, vectors of orthogonal velocity
components are plotted, indicating a vortex inside the flow field. Since no
screens and no honeycombs are installed, the vortex most likely originates
from the turbulent flow interaction between conveyance piping, jet flow and
surrounding tunnel air flow. The vortex magnitude is very low, local flow
angularities comply with ARP 5905 specifications. For the standard setup of
the icing wind tunnel, no vortex can be observed and the flow field is very
uniform as shown in Fig. b.
Turbulence intensity has been measured by means of hot-wire anemometry. The
upper part of Fig. shows the contour plot of turbulence
intensity for the ice-crystal-icing setup. The peak value of 1.4 %
complies with ARP specifications. Again, the footprint of the
pneumatic-conveyance jet is clearly visible.
Spatial distribution of turbulence intensity (a) and temperature variation (b) for the mixed phase configuration.
Liquid water content calibration at the tunnel centreline based on IKP measurements, (a) LWC history and (b) LWC depending on supplied water flow rate V˙water.
Comparison of HSI, PDI and IKP measurements for supercooled droplet conditions: (a) optical array probes vs. IKP and (b) MVD vs. volumetric flow rate of liquid water at the spray system.
Besides flow velocity characterization, measurements of total temperature
have been carried out. A very uniform temperature distribution could be
verified for the standard setup of the icing wind tunnel with flow
straightener and screens installed. Local temperature deviations during the
measurements correspond to temporal variation of tunnel temperature, which
can be adjusted with an accuracy of ±0.5 ∘C around the target
value. For the ice-crystal-icing setup, a footprint of the jet can be found
in the temperature field, see the lower part of Fig. .
For a static air temperature in the wind tunnel of 0 ∘C, the local
temperature at centre-line position is about -0.8 ∘C. At lower
temperatures, the discrepancy diminishes since the jet temperature and the
airflow temperature converge. In case of rather low temperatures of
-15 ∘C, the centre-line temperature is about 0.4 K higher than
the ambient temperature.
Calibration of the droplet cloud
Currently, two spray systems can be mounted inside the icing wind tunnel; see
Sect. . The spray system with high LWC and high MVD has
been calibrated by means of the Cranfield isokinetic probe (IKP), the
canister PDI probe, and the modular HSI probe. IKP measurements have been
conducted to correlate the liquid water content (LWC) in the centreline of
the test section with the water flow rate that is supplied to the spray
atomizers. The left image of Fig. shows temporal plots of LWC
obtained for various water flow rates at U∞= 40 m s-1. As
illustrated in the right image of Fig. , the dependency
between LWC and flow rate is rather linear. Test points corresponding to the
plots in the left image are highlighted by the same colour. For lower flow
velocities U∞, higher LWCs values are obtained as anticipated in
Eq. (). Nevertheless, the LWC increase is not proportional to
1/U∞, because the cross-sectional area of the icing cloud expands at
lower wind-tunnel speeds.
Comparisons of liquid water contents derived from canister-HSI and
canister-PDI measurements are plotted in the left image of
Fig. . The dashed lines stress deviations of
±10 %. The liquid water contents assigned to C-HSI and C-PDI have been
derived based on volume to mass correlations of detected droplets. It can be
observed that there is very good agreement between the measurements of all
three probes except for two test points of the PDI at
U∞= 20 m s-1. These measurements had been affected by
water condensation inside the C-PDI. The MVD varies in the range of 80 to
95 µm, see the right image of Fig. .
Higher MVDs have been observed with the C-PDI probe at
U∞= 40 m s-1. Presumably, droplet coalescence is promoted
for higher air velocities. Yet the medium droplet size is rather constant for
various liquid water contents at the same wind tunnel speed. All these
results refer to the ice-crystal-icing setup with the jet switched on. No
measurements have been taken away from centre-line position as neither the
canister probes nor the IKP could be traversed adequately. Ice accretion
studies of selected test models have proved a very uniform droplet cloud.
Calibration of the ice particle cloud
Ice-particle cloud calibration has been performed by means of the IKP and the
canister-HSI at the centreline of the tunnel test section.
Figure shows time histories of the ice water content
(IWC) determined by the IKP for dosing frequencies fdosing of 5
and 30 Hz at wind tunnel speeds U∞ of 20 and 40 m s-1. In
the left image absolute values are plotted while the right image shows plots
of IWC normalized by the target IWC. According to the IKP measurements,
stable cloud conditions are established after 20 to 60 s, which is
surprising, as steady state particle supply is established after 20 to
30 s. There are two possible explanations: on the one hand, there might be
a slight recirculation of ice particles inside the closed-loop wind tunnel –
not every injected ice particle might settle down at the forth corner of the
tunnel. On the other hand, the background humidity is slowly increasing due
to sublimation of the ice particle cloud, which might influence the IKP
measurement. Increasing the particle supply of the dosing machine to
frequencies higher than 30 Hz does not result in higher IWC, because the
permeability of the sieving mesh is limited. According to
Fig. , a maximum IWC of about 19 g m-3 can be
achieved at the centre-line position. At fdosing = 5 Hz, a
minimum IWC of 3 g m-3 can be adjusted. Lower feed-rates can be
obtained by use of a smaller feed pipe of the dosing machine but have not
been calibrated with respect to test section IWC. Figure
shows oscillations in IWC at low feed-rates because of an unsteady discharge
of the dosing machine. This issue has been solved by the use of a rotating
rod which supports particle discharge.
IWC time history based on IKP measurements: (a) IWC history and (b) normalized IWC history.
Based on the IKP measurements, a linear correlation between IWC and dosing
frequency has been determined. Corresponding plots are shown in the left
image of Fig. for wind tunnel speeds of
U∞= 20 m s-1 and U∞= 40 m s-1. Due to
the linear operational behaviour of the sieving and dosing machine, a strong
linear dependence between IWC and injected ice particle mass
m˙ice,inj could be demonstrated as shown in the right image
of Fig. . This dependency allows to assess the
accuracy of IWC set for a test run. For wind tunnel tests which include ice
particle supply, the IWC is adjusted based on the linear fits of
Fig. , left. The frequency of the dosing machine is
set according to the demanded IWC. After a test run, the ice particle deposit
on the sieving machine is weighed. Based on the correlation of dosing
frequency and supplied massflow (see
Fig. a) the ice particle massflow
m˙ice,inj can be determined respecting ice deposit on the
sieve. Figure then allows to deduce the real IWC
which has been established inside the test section. Additional test campaigns
where this procedure has been applied, showed an accuracy of IWC adjustments
of about ±15 %.
Ice water content depending on dosing frequency fdosing and injected ice particle flow m˙ice,inj.
Investigations on spatial IWC uniformity have been performed by means of a
custom-made particle-collecting tube system. The system allows to measure a
relative distribution of the ice particle concentration inside the test
section. A contour plot for U∞= 40 m s-1 is shown in
Fig. . Acceptable uniformity of ±20 % according
to ARP 5905 is given for a circular area of about 150 mm in diameter.
This area has proved to be sufficient and appropriate for valid icing
experiments. In agreement with the turbulent footprint of the pneumatic
conveyance jet, see Fig. , the peak IWC is located
slightly above the centre-line position, which has to be respected for the
mounting of aerodynamic test models.
Percentaged deviation of IWC from centerline value, U∞=40 m s-1 and fdosing=20 Hz.
Setup of the probe inter-comparison exercise. Alternatively to the
canister 2D-S probe, which is shown here, the canister HSI probe was mounted.
Upper left: C-HSI probe mounted inside the icing wind tunnel, modular PDI
and HSI laser beams are adjusted close to the C-HSI sampling volume.
Results of the probe intercomparison between 2D-S (LAMP), 2D-S
(DLR), M-HSI and C-HSI for seven different test points at glaciated icing
conditions.
Probe inter-comparison
Based on the IKP calibration, a cross comparison of various optical array
probes has been performed in the Braunschweig Icing Wind Tunnel. The CNRS and
DLR 2D-S probes, as well as the C-HSI have been mounted consecutively at the
same position inside the test section. Additionally, modular HSI and PDI
probes have been installed externally, to assess the repeatability of the
test conditions. Figure gives an
overview of the test setup. M-PDI and M-HSI laser beams have been adjusted
close to the sampling volume to perform non intrusive measurements in
parallel.
Figure presents a comparison of optical array probe
measurements at seven glaciated cloud conditions named I01, I02, I03, I05,
I06, I09 and I10. Ice water content has been derived based on the mass–size
relation described by . Up to a diameter of
97 µm, the particles have been treated as spherical for the
determination of the IWC. Above that threshold, the assumption of spherical
particles would lead to an overestimation of the IWC. Ice water contents
derived from measurements of both 2D-S probes and the C-HSI probe have been
related to the estimated wind tunnel IWC. A rather good agreement between the
measurements was proven as most of the measured values agreed with the wind
tunnel prescribed IWCs with an accuracy of ±20 %, which is well within
the limits of expected instrument accuracy.
Wind tunnel testing of the 2D-S probes and the C-HSI have generated a great
variety of ice particle images. The upper part of
Fig. shows examples of ice particles captured by
the C-HSI. The aggregate structure which has been observed for ice particles
before fluidization and conveyance, see Fig. , could be
maintained. The lower part of Fig. shows a series
of ice particle images obtained from 2D-S measurements inside the
Braunschweig Icing Wind Tunnel. Note that the ice particle images were
oriented vertically in a post-processing step. It can be observed that most
of the particles feature an irregular, elongated shape. Based on the analyses
of ice particle images one can conclude that the particles partially break up
on their trajectory from the sieving machine to the tunnel test section and
are reduced to smaller sizes.
Ice particle images captured at the test section of the Braunschweig
Icing Wind Tunnel. (a) images taken by C-HSI probe. (b) images taken
by 2D-S probe, MMD = 79 µm, IWC = 3.2 g m-3
and
T∞= -15 ∘C.
A characteristic size distribution of artificial ice particles inside the
icing wind tunnel is shown in Fig. a.
Furthermore, particle size distributions of atmospheric ice particles (Atm)
are included according to Fig. . The artificial
ice particle cloud includes higher concentrations of small ice particles with
equivalent diameters below 200 µm. The amount of particles
smaller than 200 µm yielded from the wind tunnel measurements
has not been identified during the airborne measurements, even for the
coldest temperatures. The high amount of small ice particles is reflected in
the cumulative mass distribution shown in
Fig. b. An MMD of about
79 µm was determined for the wind tunnel icing cloud. This value
agrees with the initial target of an MMD in the range of 50 to
200 µm. The HAIC flight campaigns have shown that against
previous assumptions natural ice crystal clouds are characterized by higher
MMDs in the range of 300 to 800 µm depending on temperature, see
Fig. b. Yet the high number of small
ice particles with MMDs below 80 µm may be of relevance for
particle size distributions in young and aged contrails
().
Comparison of artificial ice particle cloud to atmospheric cloud
conditions: (a) particle size distribution and (b) cumulative
particle size distribution.
Experimentally observed SLD ice shapes on a NACA0012 leading edge in
the parameter space of f0 and Ac, Re=2×106,
MVD = 80 µm and AoA =0∘. Results reproduced from
.
Side view of ice accretion shapes on a cylinder for TWC = 12 g m-3, mr=0.12 and U∞=40 m s-1 with a variation of temperature.
Side view of ice accretion when heating the cylinder, T∞=-5 ∘C, TWC = 6.8 g m-3 and mr=0.5.
It can be summarized that the initial design targets of the ice-crystal
generation system have been achieved. Irregular ice particles with close
agreement to natural ice particles can be injected into the wind tunnel.
Even hexagonal ice crystals can be grown by cloud chamber technology and
might be used for single particle impact studies. The ice particle size
distribution is characterized by an MMD of about 80 µm which is
between atmospheric and inside-engine particle sizes. Because the
ice-crystal generation process is fully separated from icing wind tunnel
operation, the MMD does not change with operating conditions. IWC values of
3 to 20 g m-3 correspond to total water contents expected for
inside-engine conditions and air data probes. Lower ice water content can
be adjusted by modifications of the dosing machine but require further wind
tunnel calibration measurements.
Some applications and test results
Supercooled large droplets icing
The consideration of supercooled large droplets (SLD) icing is part of the
certification for large transport aircraft ().
Appendix O of the document EASA CS25 () provides detailed
information on the MVD and LWC envelope. Outstanding requirement is the
simulation of liquid clouds with median volume diameters larger than
50 µm. The trajectories of such large droplets show a large
deviation from the air streamlines. With their increased inertia, the
droplets might be impacting or running back in regions on the aircraft
wing, which are not protected with de- or anti-icing systems.
Such SLD ice shapes and their surface roughness were investigated in the
Braunschweig Icing Wind Tunnel (). In order to
determine which wind tunnel conditions to use, they had to be related to the
relevant dimensionless parameters influencing ice roughness. These are the
dimensionless accumulation parameter Ac, given in
as the following:
Ac=LWC⋅U∞⋅taccϱice⋅δ,
and the stagnation freezing fraction:
f0=f0p,T∞,U∞,MVD,LWC,c,δ,
where ϱice represents the density of ice, δ the
diameter of the inscribed circle of the airfoil nose as a reference length,
free-stream pressure p and static free-stream air temperature T∞.
Within this two-dimensional (Ac and f0) parameter space, a Latin
hypercube sampling was performed to define a set of experiments. The
resulting ice shapes have been captured with a mould-and-cast method
() and are shown in Fig. .
These can be categorized into two groups: for shapes in the first group, which
are marked in blue, the ice thickness was almost constant over some distance
from the stagnation point and then a ridgeline appeared. Shapes in the second
group were marked in red and exhibited an intense yet smooth ice accretion
at the stagnation line followed by a much rougher zone with a feather
structure.
Mixed phase icing of a cylinder
Comprehensive investigations on ice accretion of generic test models at mixed
phase conditions have been performed in the recent past
(). The test models have been equipped
with heat foils to investigate the effect of internal heat conduction on the
accretion process.
Figure shows ice shapes on a cylinder
model without heat foil operation. A mixed phase cloud with a total water
content of 12 g m-3 and a melting ratio mr (defined as
LWC / TWC in this setup) of 0.12 has been adjusted. The static air
temperature has been varied between 0 and -15 ∘C. In contrast to
supercooled droplet icing, the ice accretion for mixed-phase icing exhibits a
conical shape. The ice accretion process appears to be strongly dependent on
air temperature. At constant melting ratio and constant total water content,
a decrease in icing severity can be observed for lower temperatures. This
effect can be addressed to supercooling of the liquid droplets inside the
mixed phase clouds. At lower temperatures enhanced supercooling allows a
large amount of liquid to freeze with only a short delay upon impact.
Therefore, less liquid is locally available to promote ice particle sticking
resulting in a significantly reduced ice accretion. Water imbibition into the
supposed porous ice layer seems therefore a necessary condition for increased
icing severity.
Figure gives an example of mixed phase
icing of the cylinder with activated heat foils. In contrast to
Fig. , no cone shaped ice accretion
develops as the accretion layer constantly melts due to internal heating.
Consequently, the accretion does not adhere to the cylinder surface and gets
dragged towards the upper side of the cylinder by shear forces induced by the
ambient airflow. Capillary forces prevent the accretion layer from instant
shedding. As the accretion layer convects downstream, further ice accumulates
in the flow stagnation area of the cylinder. In consequence, a cohesive ice
accretion structure detaches from the cylinder and growths further
downstream.
Development of measurement techniques
The experimental characterization of multiphase flows in icing wind tunnels
can be very challenging and is a fruitful topic for developing measurement
techniques as already shown in Sect. . Here,
we want to shortly summarize an optical technique to measure the 3-D geometry
of ice accretion.
Two properties of ice accretion are particularly demanding for optical
measurement techniques. On the one hand, the ice surface geometry is very
complex and includes both large roughness elements and deep chasms, see again
Fig. . On the other hand, glaze ice formations
appear transparent in the visible light range, which makes them hard to
detect with conventional camera systems (). Therefore, we
developed a general approach based on mid-infrared (MIR) laser scanning
(), because MIR radiation penetrates ice and water only
within a depth of less than 10 µm.
The MIR-line scanning method uses a CO2-laser light sheet, which is
first created through proper lenses, and then projected onto the scene. The
ice surface will absorb this light and emit an infrared radiation signature,
which is visible as a deformed line. Using a properly calibrated camera, the
deformed line can be transformed using a triangulation function yielding
3-D coordinates. While consecutively scanning with the laser light sheet over
the object, the entire 3-D ice geometry can be reconstructed. The scanning
accuracy was validated with 3-D ice objects of known dimensions; a pattern of
frustums was used.