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## B03: Dynamic impingement cooling by pulsed impinging jet arrays

**Principal investigator: **Dr.-Ing.
Frank Haucke [1] (mai [2]l)

### Summary

The experimental investigation of dynamic impingement cooling in three-dimensional array arrangements is to be further advanced. The high flow-mechanical complexity of the impinging jet interaction is to be extended by the influence of a dynamically imposed cross-flow, which increases the flow-mechanical similarity to turbine blades through which the flow passes. In order to further improve the similarity to the inner geometry of real turbine blades, the influence of a curved flow path on convective heat transfer will be analysed in detail. In cooperation with the numerical subproject B04 and the control engineering subproject B06, a significant reduction of the measurement time for the parametric variations is aimed at.

## 2nd funding period 2016-2020

- Figure 1: Test rig impingement cooling.
- [3]
- © SFB1029

Within the first project phase of the subproject a
test facility for the investigation of dynamically forced impinging
jet arrays was designed and built. The main focus of the design
process was the inclusion of a 7x7 impinging jet array as well as the
variation of the nozzle distance (S/D<=5), the variation of the
impingement distance (H/D<=5) and the variation of dynamic
parameters, such as frequency (f_{D}<=1000 Hz), phase shift
of adjacent impinging jets (Φ<=100 %), Reynolds number
(Re_{D} <= 7200) and duty cycle (DC<=100 %), planned in
the second funding phase. The developed test rig has the possibility
to freely traverse the target and the nozzle plate, which allows both
the adjustment of the impingement distance andthe accurate positioning
of the nozzle array to the target plate.

The compressed air supply to the nozzle array was implemented via four mass flow, each supplying a cooling air distributor, which in turn distributed the cooling air to 12 or 13 nozzles. The implementation of the required dynamics within the cooling air mass flow was achieved by 49 modified solenoid switching valves, whereby each nozzle could be controlled individually with regard to the required dynamic parameters. The impinging jet nozzles used [18] [23] [19] were based on the nozzle design presented by Janetzke [5] [6].

- Figure 2 Schematic LCT-setup.
- [4]
- © SFB1029

For the determination of the wall temperatures and thus for the evaluation of the cooling effect, liquid crystal thermography (LCT) is also used as a measuring method in the second funding phase. In this method, the impingment plate consists of an electrically heated 0.5 mm thick metal foil, which is equipped with a self-adhesive liquid crystal film (LCT film) on the back. The foil construct is fixed on a glass plate. This provides a stable base and favourable thermal and optical properties for heat transfer investigations (see Figure 2). To exclude an isolation effect between the individual layers, all air between the glass plate and the foils is evacuated. The LCT film is a film with encapsulated liquid crystals, which reflect different wavelengths of the illuminating light source depending on the local wall temperature, resulting in a defined, visible temperature-dependent color distribution. This can be recorded through the glass plate with a colour camera and converted into a two-dimensional temperature distribution under consideration of a colour temperature calibration. With the help of the flat temperature distribution, the local Nusselt number distribution can be calculated with the help of the known electrical heating power, the ambient parameters and the cooling fluid temperature, provided that the measurement has taken place in a state of thermal equilibrium. This similarity parameter is used to evaluate the increase of the convective heat transfer of the system by the respective set parameter combinations.

- Figure 3: Schematic PIV-setup.
- [5]
- © SFB1029

In addition to the LCT measurements, a phase-triggered stereo PIV system is used to gain insight into the flow field and the physical mechanisms responsible for the change in heat transfer (see Figure 3). For each case studied, 40 equidistantly distributed phase points within an actuation phase are recorded. To make the tracer particles (di-ethyl-hexyl-sebacate (DEHS)) visible, a Δz = 1.5mm thick light section is positioned centrally under the central row of nozzles using a Nd:YLF laser. The tracer particles are introduced into the cooling air at the cooling air distributor (see figure 4 left), whereby each nozzle is evenly supplied with seeding. The two cameras (resolution: 2560 x 2150 pixels) are stereoscopically arranged at an angle of 37° to the measuring plane and each takes 100 images for averaging. This results in an optical resolution for the measurements of Res = 0.94 vectors per mm.

The required optical components are positioned downstream of the nozzle array, whereby an influence of the flow by the components can be excluded. Two measurement planes are recorded to obtain the data. The first covers nozzles three, four and five and the second covers nozzles five, six and seven. A recording of all seven nozzles is optically not possible due to structural conditions. The overlapping of both measuring planes is used to couple the recorded data by interpolation in the overlap area.

### Constructional and metrological extension of the impingement cooling test stand.

For
reasons of noise and laser protection, a housing was provided around
the test facility. In addition, a thermal barrier was installed around
the test stand to improve the measuring accuracy and repeatability of
the results. The 60 mm thick thermal insulation used has a thermal
conductivity of L=0.04 W/mK and allows improved repeatability of the
measured values, even under fluctuating environmental conditions. Due
to the reduced heat and fluid exchange with the environment, less time
is required to reach the thermal equilibrium required for the
measurement, which allows for more efficient measurement operation.
Furthermore, the cooling air mass supply was modified. While the
original supply allowed four individual cooling air streams, seven
cooling air streams with seven nozzles each can now be operated
independently. Each line can be individually controlled by a mass flow
controller with an accuracy of 0.1 % - 0.5 % full-scale-error. In
addition to the improved accuracy of the cooling air supply, the
distributor system for the impinging jet nozzles has also been further
developed. By extending the cooling air supply from four to seven mass
flow controllers, it was possible to reduce the number of connected
cooling air nozzles from 12 or 13 to seven per cooling air distributor
on the basis of the cooling air distributors used at the beginning,
which enabled a considerable improvement in reliability and accuracy
during parameter variations. The integration of a triggerable camera
system, in connection with a new LabView-based control software,
enables a fully automatic measurement with time-synchronous
acquisition of image and temperature data. Thus, the accuracy of the
measurement results, in particular the colour calibration required for
the LCT measurement method, can be additionally increased, which in
turn positively influences the repeatability of the measurement
results. Due to the presented optimizations, the current maximum
deviation between comparable measurements could be determined to
σ_{max} = 2.79 % (see Figure 4 right). The investigation was
based on eight repeat measurements, which were carried out under
different atmospheric conditions (humidity, temperature, air
pressure). It is also evident that the repeatability at higher
actuation frequencies is marginally decreased than at frequencies
below f_{D} = 700 Hz.

In addition to the described optimizations, the flat experimental setup was supplemented by a cross-flow frame for the second funding phase (see Figure 5 left). This cross-flow frame is located between the nozzle plate and the flapper plate and encloses the nozzle array on three sides. The entire outgoing cooling mass flow is channeled in one direction, thus superimposing a frequency-dependent dynamic cross-flow on the array, which is comparable in its formation and thus in its flow physical properties to that inside a turbine blade. This means that the cross-flow velocity is coupled to the cross-sectional area of the flow channel. Therefore, the widening or reduction of the impingement distance H/D results in a decrease or increase of the prevailing cross-flow velocity (see Figure 5, right). The total transverse mass flow is thereby inevitably formed successively from the individual mass flows of the nozzle rows in x-direction. This leads to an increase in the cooling air mass flow rate dependent on the nozzle rows and thus also to an increase in the locally prevailing cross-flow velocity. A further optimization of the measurement methodology used was developed and implemented in cooperation with subproject B06. The thermal conditions in the vicinity of the test stand within the enclosure can be adjusted almost independently of external weather conditions by means of a controllable heat source. By adjusting the ambient temperature, both heat loss flows and the necessary measuring time can be minimized.

- Figure 5: Left: Schematic generation of cross flow. Right: Cross flow VS. impingement distance.
- [7]
- © SFB1029

### Variation of the plane impingement jet array arrangement with dynamic cross flow

In the first step of the second work
package, the classification of the test stand is the focus of the
investigation. For this purpose, the internationally accepted
empirical function presented by Florschuetz et al. [2] is used to
predict the Nusselt numbers in a static cross-flow nozzle array. A
comparison of the function values Nu_{FL} with own
experimental data Nu_{exp} is, due to the definition of the
empirical function, possible in the setup presented for the mean value
of the Nusselt number below the first row of nozzles, 90 degrees to
the cross-flow direction. Figure 6 on the left shows the calculation
basis for different Nusselt numbers used within the investigated
impingement jet array. The estimation function presented by
Florschuetz et al. has been established for impingement distances H/D
>> 3 and large Reynolds numbers and shows a possible error of up
to 11 % in these areas.

- Figure 6: Left: Calculation basis for Nusselt numbers. Right: Comparison between theoretical and experimental data according to Florschuetz.
- [8]
- © SFB1029

Figure 6 on the right shows that the test stand
behaves equivalently to the function of Florschuetz. For large
impingement distances the deviation between the theoretical and the
experimentally determined values is below 5%, for a reduced
impingement distance the deviation increases up to 9.6% if the
Reynolds number is reduced. A further reduction of the rebound
distance, which lies far outside the validity range of the estimation
function, shows that only the highest experimentally mapped Reynolds
number is still within the error window of the function. All in all,
the presented comparison shows that the experimental setup represents
the estimation function well within its validity range according to
Florschuetz [20]. The second step of the test stand classification was
focused on the determination of the cross-flow influence on the
convective heat transfer within the array. For the static case the
influence is shown in figure 7. The left side shows the spatial
development of the convective heat transfer in cross-flow direction
within the statically operating array without a cross-flow frame and
the right side shows the development with an integrated cross-flow
frame. The effect is shown as a function of the impingement distance
H/D. It can be seen that the convective heat transfer is significantly
reduced with increasing cross-flow velocity, i.e. with the number of
nozzle rows passed. Furthermore, it can also be seen that the
additionally channeled cross-flow causes a shift of the local maxima
in the convective heat transfer downstream. This shift grows with the
reduction of the impingement distance, i.e. with the increase of the
cross-flow velocity. The implementation of a dynamic impingement
cooling, here exemplarily shown for f_{D} = 500 Hz, minimizes
the heat transfer reducing effects.

- Figure 7: Cross-flow influence on heat transfer in the case of steady blowing at ReD = 7200 and S/D = 5 depending on the impingement distance H/D.
- [9]
- © SFB1029

Figure 8 shows the spatial Nusselt number
distribution, normalized to a Reynolds number of Re_{D} =
2150, for the dynamic cases compared to the static case with a
Reynolds number of Re_{D} = 7200. It can be seen that the
dynamics can completely compensate for the reduction of the maxima in
the cross-flow direction. Furthermore, the observed shift of the
maxima in convective heat transfer is also significantly reduced. The
normalization to the lower Reynolds number allows the identification
of the Reynolds number dependent increase of the local cooling effect.
This leads to the conclusion that the vortex systems created by the
dynamics are able to penetrate the cross-flow with reduced
interaction, allowing more cooling fluid to reach the
impingement plate. This can be attributed to the temporarily
increased exit velocity of each impingement jet compared to the steady
blowing case, which results in an increase of the impulse ratio
between the impingement jet and the cross-flow on a time average. The
effects derived from Figure 8 can also be observed in corresponding
PIV data.

- Figure 8: Effect of dynamic impingement cooling on the cross-flow as a function of the impingement distance H/D at a Reynolds number of ReD = 7200 and a nozzle distance of S/D = 5.
- [10]
- © SFB1029

Figure 9 shows the mean velocity fields for the
static case f_{D} = 0 Hz and for an exemplary dynamic case
f_{D} = 500 Hz. In both cases an identical cooling air mass
flow rate was used. From this it follows that the mean discharge
velocities of both cases are identical. As a result of the periodic
interruption, depending on the selected pulse width, a pulse increase
in the respective impingement jet results both temporarily and on a
time average, whereby the local pulse ratio between jet and cross-flow
is increased compared to the stationary case. This can be seen from
the increased jet velocities near the impingement plate and the
associated reduced jet deflection in the cross-flow direction. Due to
the increased velocities close to the wall, increased wall shear
stresses are also associated, which in turn mean increased convective
heat transfer at the wall [14]. After the basic classification
of the test rig and the determination of the cross-flow
influence, the convective heat transfer was maximized depending on the
frequency, the impingement distance and the Reynolds number. For this
purpose, all parameters are varied within the following limits:
f_{D} = 0...1000 Hz, Δf_{D} = 100 Hz, H/D = 2, 3, 5
and Re_{D} = 3200, 5200, 7200.

- Figure 9: Deflection of the imping jets by the cross-flow. PIV data, ReD = 7200, H/D = 2, S/D =5.
- [11]
- © SFB1029

Figure 10 shows the global effect of variation on
the convective heat transfer within the array, as a function of the
actuation frequencies f_{D}. A global averaged Nusselt number
Nu (see Figure 6 left) is formed and normalized to the global average
Nusselt number of the corresponding static case Nu_{0}. The
data shows that the global frequency-dependent course of all cases
shows a high degree of similarity. Each parameter combination examined
displays an increased convective heat transfer compared to the
static case and shows a growth maximum of 51% at f_{D} =
700Hz. Furthermore, the data show that the growth potential is reduced
with the impingement distance, whereas the increase in the Reynolds
number for the impingement distances of H/D = 3 and H/D = 5 shows a
potential increase in growth [15, 16, 21].

- Figure 10: Global influence of pulse frequency on convective heat transfer in Dependence of the Reynolds number ReD and the impingement distance H/D at S/D = 5.
- [12]
- © SFB1029

Figure 11 shows the vorticity fields averaged over
40 phase points of the last five nozzles in cross-flow direction for
the static (f_{D} = 0 Hz), the f_{D} = 500 Hz and for
the f_{D} = 700 Hz case. It can be seen that specific
actuation frequencies clearly influence the resulting vorticity in the
shear layer. An increased vorticity near the wall generally leads to
increased local wall shear stresses and thus to an increased
convective heat transfer. This effect can also be observed between the
impingement zones, whereby the spatial cooling effect within the
array can be significantly increased by the implementation
of a dynamic [22]. In addition to the global frequency influence, the
phase shift between the individual rows of nozzles is also a relevant
factor for optimizing the global convective heat transfer. In order to
analyse this, the phase position between the individual nozzle rows is
varied within the limits Φ = 0 ... 90 %. The variation takes place,
depending on the Reynolds number and the impingement distance, at
three representative actuation frequencies. These are f_{D} =
300Hz, f_{D} = 500Hz and f_{D} = 700Hz, which
corresponds to a weak, a moderate and the most effective actuation
frequency.

- Figure 11: Frequency influence on the z-vorticity averaged over 40 phase points for fD = 0Hz, fD = 500Hz and fD = 700Hz for ReD = 7200 and S/D = 5.
- [13]
- © SFB1029

Figure 12 shows the development of the global
Nusselt number for the representative impingement distance of H/D = 2
for the investigated Reynolds numbers and frequencies. Especially for
high Reynolds numbers a considerable additional potential for the most
effective frequency f_{D} = 700Hz can be observed. By varying
the phase shift, which symbolizes the time-controlled sequence of
adjacent ring vortices, the vortex interaction can be optimized with
respect to the resulting heat transfer at the wall. With a phase shift
of Φ = 50%, an additional increase in convective heat transfer of up
to 12% can be recorded. The two other frequencies shown also show
significant increases in convective heat transfer. Depending on the
phase offset, a further increase of up to 16% can be generated [16].
The increase in the cooling effect, especially at Φ = 50%, is one of
the driving factors for the use of fluidic actuators in the further
development of the system, since these have a functionally fixed phase
shift of Φ = 50%.

- Figure 12: Exemplary influence of phase shift on convective heat transfer as a function of frequency fD and Reynolds number ReD at an impingement distance of H/D = 2, at a nozzle distance of S/D = 5.
- [14]
- © SFB1029

An equally relevant parameter is the spatial
distance between the nozzles (see Figure 13). Investigations show that
the frequency-dependent convective heat transfer is influenced by the
reduction of the nozzle distance S/D in such a way that for small
impingement distances H/D = 2, a reduction of the nozzle distance to
S/D = 3 allows a potential increase of 60 % at specific frequencies.
However, when larger distances are investigated, a deterioration is
seen for almost all frequencies investigated. This trend continues
when the nozzle distance is reduced to S/D = 2.5. In this case, even
for the small impingement distance only a maximum increase of
35% is possible. In addition, the increase can only be generated in a
frequency range up to f_{D} = 400Hz. This indicates that with
the running length of the vortex and the associated vortex diameter
increase, there are increased interactions between adjacent vortices,
which can have a negative influence on the fluid transport properties
of the vortex structures. Corresponding PIV data as well as data
regarding the influence of phase shift at reduced nozzle distance have
already been recorded and are currently in the evaluation and
publication phase.

- Figure 13: Influence of nozzle distance S/D on convective heat transfer as a function of frequency fD and impingement distance H/D = 2 at a Reynolds number of ReD = 7200.
- [15]
- © SFB1029

### Variation of the 2D curved impingement jet array arrangement with dynamic cross flow

Following the flat test configuration, work package three focuses on a curved configuration. The selected radius of curvature of r = 1 m is based on a radius of curvature of an impingement-cooled inner surface of a NASA E3 turbine blade profile scaled to the nozzle outlet. Figure 14 shows the modified impingement plate, nozzle plate, and cross-flow frame. The curvatures take into account that the radius of the nozzle plate must be adjusted with the impingement distance to ensure a 90° angle between the target plate and the exiting impingement jet. The target plate is a curved glass plate allowing the use of both LCT and PIV test setups, which are equivalent to those shown in Figures 2 and 3.

First results regarding convective heat transfer
and the resulting cooling effect in terms of the global Nusselt number
are shown in Figure 15. Here, the general trend with regard to the
frequency-dependent increase of the cooling effect is comparable to
that of the plane configuration under the same geometric conditions
(see Figure 15 left). The difference between the two global
frequency-dependent Nusselt number curves of the curved and the plane
plate is shown in Figure 15 left. It can be seen that the global
growth maximum is at f_{D} = 800 Hz and that for most
frequencies below f_{D} = 600 Hz the difference between both
configurations is less than 5 %. The exceptions are f_{D} =
200 Hz and fD = 400 Hz, with the first frequency showing a Nusselt
number reduction of about 34% and the higher frequency a Nusselt
number improvement of 15%. This surprising and as yet unexplained
behaviour will be investigated in more detail by means of a finer
frequency resolution. In addition to the first effects, frequencies
above f_{D} = 600 Hz also show a Nusselt number reduction;
however, this is almost constant at 16%, independent of
frequency.

- Figure 15: Nusselt number difference between plane and curved configuration as a function of frequency fD at a Reynolds number of ReD = 7200, an impingement distance of H/D = 3 and a nozzle distance of S/D = 3.
- [17]
- © SFB1029

At this point it is also to be noted that the
cooling maxima found in this configuration at f_{D} = 350 Hz
and f_{D} = 800 Hz are similar to the numerically determined
results of subproject B04. In subproject B04 the most receptive modes
were found at Strouhal numbers of Sr_{D} = 0.46 and
Sr_{D} = 0.92. If the found most effective frequencies of this
configuration are converted to the Strouhal numbers Sr_{D}
related to the nozzle outlet, Strouhal numbers of Sr_{D} =
0.46 and Sr_{D} = 1.05 result for both the plane and the
curved surface despite different exit Mach number and impingement
distance. Thus, for the first preliminary results of the curved test
configuration it can be stated that the global behavior of the
frequency-dependent Nusselt number development is similar to the plane
configuration. However, a comparison of the absolute growth rates,
especially at high frequencies, shows a lower increase in cooling
effect in the case of the curved test body. Furthermore, it can be
seen that the experimental and numerical results show a similarity
with regard to the expected frequencies for the cooling
maxima.

### Optimization and control of the dynamically forced impingement cooling

For the optimization and control of the dynamic impingement cooling, cooperation with subproject B06 is underway. Thereby it is necessary to be able to evaluate the changes of the cooling capacity in real-time. This is possible by using Pt-100 sensors instead of liquid crystal thermography. Figure 16 on the left shows the modification of the experimental setup presented in detail in Figure 2. Here, the LCT foil is replaced by 15 Pt-100 sensors, which are positioned at an even distance below the center section of the nozzle array. Compared to the LCT film, the sensors have a much higher bandwidth, which means that the dynamics of the surface temperature are better resolved by the measurement. The heat conduction from the plate surface to the rest of the experimental setup is much slower. It can therefore be assumed that this only slightly distorts the determination of convective heat transfer from the plate using the measured surface temperature. In order to guarantee the real-time capability of the measurements, it is not possible to wait for a thermal equilibrium of the entire system. Therefore, only general qualitative trends can be detected and not quantitative ones. Within this work package, in cooperation with subproject B06, an innovative Fast Extremum Seeking algorithm has been developed in such a way that it is possible to estimate the thermal response of the investigated overall system by using pertubated parameter ramps. The acquisition of a ramp takes only five minutes, which represents an elementary time saving compared to the measurement methods used so far [17]. The experiments performed show that the developed algorithm is able to detect the frequency-dependent course of the cooling capacity with good accuracy (see Figure 16 right). Thus, it is possible for an impingement cooling system to detect maxima and minima in the heat transfer due to changed operating conditions in a comparatively short time and to store the thermal behavior as a data matrix [17].

- Figure 16: Left: Modification of the experimental setup. Right: Comparison of LCT data to Pt-100 data with Fast Extremum Seeking Algorithm at a Reynols-number of ReD = 7200, an impingement distance of H/D = 3 and a nozzle distance of S/D = 5.
- [18]
- © SFB1029

## Literature used (extract)

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M. Bauer, J. Lohse, F. Haucke, and W. Nitsche. High-lift
performance investigation of a two-element configuration with a
two-stage actuator system. AIAA Journal - Technical Notes, 2014.

**[2]** L. W. Florschuetz, C. R. Truman, and D.
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acoustic excitation positions on heat transfer and flow in
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excitation. Int. J. of Heat and Fluid Flow, 24:199–209, 2003.

**[5]** T. Janetzke and W. Nitsche and J. Täge.
Experimental investigations of flow field and heat transfer
characteristics due to periodically pulsating impinging air jets. Heat
and Mass Transfer, 45:193–206, 2008. DOI: 10.1007/s00231-008-0410-8
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resolved investigations on flow field and quasi wall shear stress of
an impingement configuration with pulsating jets by means of high
speed piv and surface hot wire array. Int. J. Heat and Fluid Flow,
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## Own publications relevant to the project

**[14]** **A. Berthold
(B03) **and **F. Haucke (B03)**. Experimental
investigation of dynamically forced impingement cooling. In
Proceedings of ASME Turbo Expo 2017, volume 5A: Heat Transfer. ASME
Turbo Expo 2017: Turbomachinery Technical Conference and Exposition,
2017.

**[15]** **A. Berthold (B03) **and
**F. Haucke (B03)**. Experimental study on the alteration
of cooling effectivity through excitation-frequency variation within
an impingement jet array with side-wall induced crossflow. Active Flow
and Combustion Control 2018, Ch: Springer International
Publishing.:(S.339–354), 2018.

**[16]** **A. Berthold (B03) **and
**F. Haucke (B03)**. Influence of excitation frequency,
phase-shift and duty cycle on cooling ration in a dynamically forced
impingement jet array. In ASME Turbo Expo 2019, number GT2019-90695.
ASME Turbo Expo 2019: Turbomachinery Technical Conference and
Exposition, 2019.

**[17]** **B. Fietzke (B06)**,
**M. Kiesner (B06)**, **A. Berthold (B03)**,
**F. Haucke (B03)**, and **R. King (B06)**.
Map estimation for impingement cooling with a fast extremum seeking
algorithm. Active Flow and Combustion Control 2018, Ch: Springer
International Publishing.:(S.367–378), 2018.

**[18]** **F. Haucke (B03)**, H.
Kroll, **I. Peltzer (B01)**, and **W. Nitsche
(B03)**. Experimental investigation of a 7 by 7 nozzle jet
a array for dynamic impingement cooling. In
Active Flow and Combustion Control 2014, Editor R. King,
dx.doi.org/10.14279/depositonce-39 [20]. Springer, 2014.

**[19]** **F. Haucke
(B03)**,** W. Nitsche (B01)**, and **D.
Peitsch (B03)**. Enhanced convective heat transfer due to
dynamically forced impingement jet array. In Proceedings of ASME Turbo
Expo 2016, number GT2016-57360, 2016.

**[20]** **F. Haucke (B03) **and
**A. Berthold (B03)**. Experimental investigation of a
dynamically forced impinging jet array. New Results in Numerical and
Experimental Fluid Mechanics XI, Notes on Numerical Fluid Mechanics
and Multidisciplinary Design 136, 136:209–218, 2016.

**[21]** B. Ataseven, **A. Berthold
(B03)**, and **F. Haucke (B03)**. Influence of
system resonance on a dynamically forced impinging jet. In ASME Turbo
Expo 2019: Turbomachinery Technical Conference and Exposition,, number
Poster No. GT2019-92266, 2019.

**[22]** **A. Berthold (B03), F. Haucke
(B03), **and J. Weiss. Flow field analysis of a dynamically
forced impingement jet array. In 2020 AIAA SciTech Forum, 2020.

**[23]** **F. Haucke
(B03)**,** W. Nitsche (B03)**, **R. Wilke
(B04)**, and **J.L. Sesterhenn (B04)**.
Experimental and numerical investigation regarding pulsed imipingement
cooling. In Deutscher Luft- und Raumfahrt Kongress Rostock,
2015.

## Speaker

**Prof. Dr.-Ing. Dieter Peitsch**

e-mail query [21]

## Office

Steffi Stehrsec. F 2

Room 107

Marchstr. 12

10623 Berlin

+49 (30) 314 22954

e-mail query [22]

haucke/

parameter/en/id/122150/?no_cache=1&ask_mail=Xzg3RQA

J4hM2a%2FshjHpmLzc20b3qlHn9ItZ7RJDA2v3C%2BHsF5LTOVA%3D%

3D&ask_name=FRANK%20HAUCKE

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