January 2014 

NASA/TM–2014-218154 

Adaptive Back Sheet Material for Acoustic 

Liner Applications—ARMD Seedling Fund 

Phase I Final Report 

Carl H. Gerhold and Michael G. Jones 

Langley Research Center, Hampton, Virginia 

Dawnielle Farrar 

John Hopkins University Applied Physics Laboratory, LLC, Laurel, Maryland 
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January 2014 

NASA/TM–2014-218154 

Adaptive Back Sheet Material for Acoustics 

Liner Applications—ARMD Seedling Fund 

Phase I Final Report 

Carl H. Gerhold and Michael G. Jones 

Langley Research Center, Hampton, Virginia 

Dawnielle Farrar 

John Hopkins University Applied Physics Laboratory, LLC, Laurel, Maryland 
Available from: 

NASA Center for AeroSpace Information 

7115 Standard Drive 

Hanover, MD 21076-1320 

443-757-5802 

The use of trademarks or names of manufacturers in this report is for accurate reporting and does not constitute an 

official endorsement, either expressed or implied, of such products or manufacturers by the National Aeronautics 

and Space Administration. 
1 

Abstract

A recently developed piezo-electric composite film is evaluated for its usefulness in application in acoustic 

liners. Researchers at the NASA Langley Research Center Liner Technology Facility developed 

experiments to measure the electrical response of the material to acoustic excitation and the vibrational 

response of the material to electrical excitation. The robustness of the piezo-electric film was also assessed. 

The material’s electrical response to acoustic excitation is found to be comparable to a commercial 

microphone in the range of frequencies from 500 to 3000 Hz. However, the vibrational response to 

electrical excitation in the frequency range of interest is an order of magnitude less than may be necessary 

for application to acoustic liners. Nevertheless, experimental results indicate that the potential exists for the 

material to produce a measurable change in the impedance spectrum of a liner. Work continues to improve 

the authority of the piezo-electric film. 

 

Nomenclature 

A = amplitude of the incident wave in the duct 

B = amplitude of the reflected wave in the duct 

pt = total sound pressure in the duct 

vt = total acoustic particle velocity in the duct 

f = frequency 

c = sound speed in air 

t = perforate thickness 

d = perforate hole diameter 

h = honeycomb chamber depth 

M = Mach number of flow in the duct 

k = wavenumber (=2*π*f/c) 

L = distance from the piezo-electric film sample to the rigid termination 

Z = acoustic impedance = R +i*I 

R = Resistance, real part of the acoustic impedance 

I = Reactance, imaginary part of the acoustic impedance 

ρ = density of the air 

ζ = acoustic impedance, Z, normalized by the characteristic impedance, ρc

θ = phase angle 

Cα

 = absorption coefficient 

 i = −1 

Subscripts 

1 = location of microphone 1 for impedance evaluation 

2 = location of microphone 2 for impedance evaluation 

s = indicates a piezo-electric film 

l = indicates an SDOF perforate-over-honeycomb liner 

m = mass reactance term of impedance 

tl = transmission line term of impedance 

I. Introduction 

The purpose of this research project is to demonstrate the applicability of a newly discovered smart 

composite material to serve as a compliant back wall for an acoustic liner. A compliant back wall allows 

the liner to respond to changes in the aircraft engine shaft speed related tone noise by actively modifying 

the liner impedance. The material being investigated consists of an α-Helical Polypeptide, which has been 

developed by researchers at the Johns Hopkins University Applied Physics Laboratory1

, and has shown 

promise as an actuator capable of being excited at frequencies in the audible range. This project is the first 

phase of an Aeronautics Research Mission Directorate (ARMD) Seedling Fund project. 

The basic parameter used to describe the acoustic properties of the piezo-electric composite is the 

impedance. The impedance is defined as the ratio of the sound pressure to the acoustic particle velocity 

normal to the surface. The impedance consists of a real term that describes the resistance to sound 
2 

propagation and an imaginary term that describes the frequency-dependent reactance. The impedance 

quantifies the boundary condition presented by the wall of a duct and is used to complete mathematical 

models of sound propagation in waveguides. Derivation of the expression for impedance will be presented 

in the next section. 

II. Results 

The two key deliverables of this Phase I include: 1) Built-up liner samples with piezo-electric actuators 

bonded as a back wall, and 2) Mathematical model of the impedance of the liner incorporating the active 

back wall. The test coupons that were fabricated and evaluated consist of the piezo-electric film sample 

bonded to a sleeve with a small air gap behind. The evaluation of the samples identifies the impedance of 

the piezo-electric film plus air gap and the impedance is then incorporated into a mathematical model of the 

built-up honeycomb liner with compliant back wall. 

1. Description of Samples Tested 

The researchers at Johns Hopkins University Applied Physics Laboratory have fabricated five test coupons 

mounted on plastic sleeves for testing. An example coupon is shown in Figure 1. The piezo-electric sample 

is clear. A metal electrode is attached to the sample in order to deliver voltage to the piezo-electric (actuator 

mode) or to sense voltage from the sample (sensor mode). Three such samples with electrodes of various 

metals and design (1.0 in x 1.0 in Cr-Au, 1.75 in x 1.75 in in Cr-Au, Conductive adhesive) were evaluated 

in the test. Two other samples were fabricated and mounted on the plexiglass sleeves. These samples do not 

have electrodes bonded to them, but rather are encapsulated in a plastic sheathing for improved durability. 

Figure 1. Photograph of piezo-electric sample mounted on plexiglass sleeve sample holder. 

It became necessary to encapsulate the samples when it was discovered that the original piezo- electric 

samples are too fragile to be subjected to sound pressure levels in excess of 120 dB. Table 1 briefly 

describes the characteristics and summarizes the current condition of the five mounted samples. In addition 

to the mounted samples, two unmounted samples of the encapsulated piezo-electric composite have also 

been fabricated. 

2. Characterization Of The Composite Samples 

The normal incidence impedance is evaluated based on measurements performed in the Normal Incidence 

Tube in the Liner Technology Facility at NASA Langley Research Center. 

An example set up is shown in Figure 2. The piezo-electric film sample, which is affixed to the plexiglass 

sleeve, is located approximately 1.25 inch above the rigid plate, shown at the bottom of the photograph. An 
3 

empty sleeve approximately 3.0 inch length is attached above the sample and connects to the Normal 

Incidence Tube. Sound is generated by four loudspeakers whose output is adjusted to achieve a specified 

sound pressure level at the reference microphone, which is indicated at the top of the photograph. The 

sound pressure level is measured at two microphones z1 and z2 that are located 5.50 and 6.75 inch, 

respectively, above the sample. These microphones are not visible in the photograph. 

Table 1. Piezo-electric sample configurations evaluated to date

Sample Description Electrode Electrode size Status

JHS1 Composite film, no 

encapsulate 

Cr-Au, evaporated 1.00”x1.00” Functional

JHS2 Composite film, no 

encapsulate 

Cr-Au, evaporated 1.75”x1.75” Destroyed 

JHS3 Composite film, 

encapsulated 

No electrode N/A Functional 

JHS4 Composite film, no 

encapsulate 

Conductive 

adhesive 

1.75”x1.00” Destroyed 

JHS5 Composite film, 

encapsulted 

No electrode N/A Functional 

Figure 2. Photograph of impedance test set up. 

The total sound pressure, pt in the hard wall duct at any location x above the sample is given by: 

pt(x) = Ae−ikx + Beikx 

The ratio of sound pressure at microphone locations x1 and x2 is: 

Piezo-electric sample 

Reference microphone 
4 

pt(x2 )

pt(x1 )

= p2

p1

e

iθ = Ae−ikx2 + Beikx2

Ae−ikx1 + Beikx1





 where: 

p2

p1

 is the ratio of the magnitudes of the sound pressures 

θ is the phase of signal 2 relative to signal 1 

The ratio of amplitudes, B/A is determined by rearranging: 



B

A =

e−ikx2 − p2

p1

e

iθ e−ikx1

p2

p1

e

iθ eikx1 − eikx2



The location of the piezo-electric film sample is defined as x = 0. The sound pressure at this location is: 

	

and the particle velocity is: 




The impedance, Z at location x = 0 is the ratio of the total sound pressure in the duct to the total particle 

velocity. The impedance is often expressed in a form in which it has been normalized by the characteristic 

impedance: 



ζ = Z

ρc

= R

ρc

+ i I

ρc

= pt(0)

vt(0) =

1+

B

A

1− B

A



The impedance that is evaluated in the set up of Fig. 2 combines the impedance of the piezo-electric film 

and the impedance of the air gap beyond it: 



ζ = ζ s − i cot(kL)

The sample is subjected to broadband noise from 400 to 3000 Hz with overall sound level at the reference 

microphone set at 110, 120, 130, or 140 dB. Single tones are also generated from 400 to 3000 Hz in 200 Hz 

increments. Figure 3 shows a typical impedance spectrum evaluated from the measured data for the sample 

JHS3 with broadband noise input. This sample is described as an encapsulated, mounted composite. The 

test set up includes the sample backed by a cavity that terminates in a rigid wall. The figure shows that the 

sample has a resonant frequency at approximately 525 Hz, which is indicated when the reactance goes from 

negative to positive. This is the lowest order resonant frequency of the membrane consisting of the 

encapsulated piezo-electric film backed by an air gap. The impedance becomes quite large from 2000 to 

3000 Hz, indicating the sample acts as a rigid surface at these higher frequencies. 

The effect of the piezo-electric film with the air gap is to change the overall impedance of a liner. This is 

shown in the following example, in which the rigid back wall of a standard, single degree of freedom 

perforate-over-honeycomb liner is replaced with a back wall consisting of the composite backed by a 

cavity. Jones, et al.2

 use a Two Parameter Impedance Prediction model to define the impedance parameters, 

pt(0) = A + B

vt(0) = 1

ρc

(A − B)
5 

the resistance and reactance, of a single-degree-of-freedom perforate over honeycomb core liner. The 

expression for impedance, normalized by the characteristic impedance is given as: 

 

Zl

ρc

= Rl

ρc

(d,t,M )+

i

ρc

Im ( ) ( f ,t,d + Itl(k,h))


Figure 3. Impedance of test set up with sample JHS3 evaluated with two-microphone method using 

broadband noise at 130 dB overall sound level. 

The resistance term, 

Rl

ρc

, is defined by the thickness of the perforated facesheet, the hole diameter, and the 

Mach number of the grazing flow in the duct. The Mach number is 0 in the Normal Incidence Tube 

experiment. The reactance, 

1

ρc

*(Im + Itl), depends on the perforate thickness and hole diameter, the 

core depth of the honeycomb, and the frequency. The reactance component of the impedance consists of a 

mass reactance, Im , which varies linearly with frequency and a transmission line, Itl . The transmission 

line term is evaluated as follows. 

Figure 4a shows a typical liner cross-section with perforate face sheet, honeycomb core, and rigid back 

wall. The depth of the honeycomb core is h. For the purposes of this discussion, the location of back wall is 

designated x = 0. In this case then, the face sheet is located at x = -h. 

The sound transmission through the honeycomb core characterizes the impedance seen at the face sheet: 

Itl

ρc

(−h) = Aeikh + Be−ikh

Aeikh − Be−ikh



The expression for the impedance at the termination of the honeycomb is: 

Z

ρc

(0) = A + B

A − B 

If the honeycomb is terminated by a rigid wall, Z = ∞ or, A=B, and the impedance seen at the face sheet 

is: 
6 

Itl

ρc

(−h) = −i cot(kh)

 a b 

Figure 4. Schematic representation of a standard perforate-over-honeycomb liner with honeycomb of depth 

h terminating; (a) at a rigid wall, and (b) with compliant back wall at h. 

The transmission term of equation 1.11 is incorporated into the impedance equation (1.8) in order to 

estimate the sound absorption. 

When the rigid back wall is replaced by the piezo-electric film with the air gap, as shown in Figure 4b, the 

impedance at x = 0 is finite and is evaluated from: 

Itl

ρc

(0) = A + B

A − B = ζ



from which: 

B = ζ −1

ζ +1

A



The transmission term, equation 1.9 is written: 

Itl

ρc

(−h) = ζ cos(kh)+ isin(kh)

iζ sin(kh)+ cos(kh)

	

It is assumed that a small gap exists between the honeycomb cells and the piezo-electric film such that the 

motion of the film is not constrained by contact with the edges of the honeycomb cells. The impedance in 

equation 1.8 incorporating the transmission term of equation 1.14 is used to estimate the sound absorption,

Cα: 

Cα = 1−

Zl

ρc

−1

Zl

ρc

+1

⎛

⎝

⎜

⎜

⎜

⎞

⎠

⎟

⎟

⎟

conj

Zl

ρc

−1

Zl

ρc

+1

⎛

⎝

⎜

⎜

⎜

⎞

⎠

⎟

⎟

⎟




Figure 5 shows the absorption calculated for the liner with the rigid back wall, configuration a in Figure 4, 

compared with the absorption for compliant back wall, configuration b in Figure 4. In this case, the 

honeycomb core is 1.5 in deep, the perforate is 0.040 in thick with 0.040 in diameter holes. The perforate 

open area is 8.7%. The figure shows that the hard wall backed liner has maximum absorption of ~80% at 

















	
	




		
 









7 

2200 Hz. This is the resonance of the rigid wall backed liner, where cot(kh) ~ 0. The compliant back wall 

adds another absorption peak near 500 Hz. This was noted as the resonant response of the encapsulated 

piezo-electric film backed by an air gap in the normal impedance curve, Figure 3. So replacing the rigid 

back wall with a compliant back wall gives rise to an increase in absorption when the back wall goes into 

resonance. It was also noted that the piezoelectric impedance is very high in the range from 2000 to 3000 

Hz. This means that the piezoelectric appears as a hard wall from 2000 to 3000 Hz and thus the liner 

behaves as the rigid back wall liner, with the high absorption in the vicinity of 2200 Hz. 

The analysis indicates that the frequency range for high liner absorption can be extended with the passive 

compliant back wall. 

Figure 5. Calculated absorption of a perforate-over-honeycomb liner comparing rigid back wall to 

compliant back wall, JHS3. 

3. Further Characterization Of The Piezo-Electric Composites 

The impedance of the passive sample can be measured and the resulting impedance incorporated into the 

analytical model for the overall impedance of a liner with compliant back wall, as was shown above. 

Refinement of the analytical model requires that the response of the sample to acoustic excitation and the 

response of the sample to electrical excitation be known. Experimental procedures were developed in order 

to obtain these relationships. The experimental techniques were developed using the film sample JHS1, 

which has electrodes for electrical excitation. 

A. Piezo-electric sample as sensor. 

While the primary purpose of this project is to assess the piezo-electric material as an active device, it may 

be useful as a sensor. This is particularly the case in a feedback control loop where one piezo-electric film 

is used as a sensor whose signal drives an actuator piezo-electric film. In order to evaluate the piezoelectric film as a sensor, it is subjected to a known acoustic signal and the voltage generated at the electrode 

is measured. Figure 6 shows a typical plot of voltage out vs. sound pressure in, where the pressure is 

expressed in Pascals. The sample shows good linearity between the voltage generated by the piezo-electric 

film and the pressure input. The sound pressure level at the sensor was 97 < SPL < 124 dB. 

Table 2 shows the sensitivity of the sensor at different frequencies. The sensitivity is higher at 1500 Hz 

than at other frequencies. This indicates that the sensor resonance is in the vicinity of 1500 Hz. The sample 

JHS4 was also evaluated before it failed. The sensitivity of that sample at 1000 Hz was found to be 0.00070 
8 

v/Pa, which is comparable to the sensitivity determined for sample JHS1. The sensitivity of these piezoelectric devices, which is on the order of 1 mV/Pa, is comparable to the sensitivity of a 1/8 in diameter 

condenser microphone, although the variation of sensitivity with frequency is higher than is expected in a

precision microphone. Thus, this piezo-electric device could be suitable in sensor mode if it is calibrated. 

Figure 6. Voltage generated at the electrode of sample JHS1 when subjected to sound at 1000 Hz. 

Table 2. Measured sensitivity of sensor JHS1, voltage output for acoustic signal input 

Frequency (Hz) Input sensitivity (v/Pa) 

500 0.00101 

1000 0.00082 

1500 0.00266 

2000 0.00082 

2500 0.00122 

3000 0.00107 

B. Piezo-electric as an actuator. 

The primary purpose of this project is to evaluate the piezo-electric sample as an acoustic actuator. For this 

reason, an experiment was devised in which the out-of-plane velocity at the center of the sensor is 

measured as a function of voltage excitation to the electrode. The test set-up is shown in Figure 7. The 

velocity was measured using a Polytec PDV 100 portable digital laser vibrometer, on loan from the 

Systems Integration and Test Branch at NASA Langley Research Center. The sample was excited by a tone 

from an HP model 3325A function generator, amplified using a PiezoSystems, Inc, model EPA-104-115 

piezo-electric amplifier. The amplifier is on loan from the Advanced Sensing and Optical Measurement 

Branch at NASA Langley Research Center. The amplifier generates a dynamic signal of up to +/- 200 volt 

to the electrode. 

Figure 8 shows the rms velocity response of the sample JHS1, measured at the center of the sample, for 

excitation at 1500 Hz. The curve is linear over the range of excitation voltage 40 < v < 140 volt. It was 

noted throughout the data acquisition that the velocity response at the first harmonic of the excitation 

frequency was quite strong. It was the same order of magnitude as the response at the excitation frequency

and, in some cases, was even higher. 

	


















 
  
  
  







9 

Figure 7. Test set up for measurement of piezo-electric sample velocity response to AC voltage excitation. 

Figure 8. Velocity response of sample JHS1 to dynamic signal excitation at 1500 Hz. 

Table 3 summarizes the measured response of the piezo-electric sample JHS1 to voltage excitation over the 

range of frequencies from 500 to 3000 Hz. The sensitivity of the piezo-electric sample to voltage excitation 

is higher at 1500 and 2000 Hz than it is at the other frequencies. This behavior is consistent with higher 

response at 1500 Hz noted with acoustic excitation. 











 







       


!""#


$
10 

Table 3. Measured sensitivity of piezo-electric sample JHS1 to dynamic voltage input. 

Frequency (Hz) Output sensitivity (mm/s/volt) 

500 .00027 

1000 .00040 

1500 .00077 

2000 .00087 

2500 .00032 

3000 .00034 

The velocity response is less than what is expected to be required as an actuator. For example, the acoustic 

particle velocity at 130 dB sound level is approximately 150 mm/sec. The velocity of the piezo-electric 

film, when actuated at 150 volt at 1500 Hz, is approximately 0.12 mm/sec, or 0.1% of the particle velocity. 

However, since the velocity of the rigid back wall of the liner is zero, even a small motion of the back wall 

may be sufficient to change the liner response near the resonant frequency of the liner. 

 

4. Measured Actuator Performance Under Acoustic Load

In order to assess the change in performance with voltage excitation of the piezo-electric sample JHS1 

under acoustic load, the impedance is measured first without and then with power to the sample. The 

sample, with no power to the electrode, is subjected to sound at 110 dB in tones from 400 to 3000 Hz in 

200 Hz increments, and the impedance determined. The experiment is repeated with the electrode powered 

by a signal at the same frequency as the acoustic excitation and amplified to approximately 150 V (0-peak). 

No attempt is made to control the phase of the excitation in this experiment. The impedance is determined 

under this condition. The absorption of the piezo-electric film with the air gap is determined from the 

resulting impedance measurements and plotted in Figure 9. The most significant effect of the excitation of 

the piezo-electric occurs near the resonance, where the absorption is greatest. The peak absorption of the 

unpowered piezo-electric sample occurs at 1600 Hz, which is the unforced resonance frequency. 

Energizing the sample shifts the peak absorption from 1600 to 1400 Hz. Replacing the 0 velocity at the 

back wall with a finite velocity makes the chamber appear to be deeper and thus lowers the resonant 

frequency. 

Figure 9. Calculated absorption coefficient for sample JHS1, compares power off to power on. 

5. Predicted Performance Of Active Back Wall In A Liner 
11 

The effect of installing the piezo-electric film in a liner configuration and energizing it is estimated. The 

impedance of the piezo-electric film without excitation is used to determine the transmission term of the 

impedance, equation 1.14 and the overall impedance is estimated using equation 1.8. The energized 

impedance is then used to determine the transmission term and the overall liner impedance is determined. 

This overall impedance is used to estimate the sound absorption for the two conditions. Figure 10 shows 

the calculated absorption of the liner with a passive compliant back wall (power off) and an energized 

compliant back wall (power on). When the power to the piezo-electric film is off, the absorption curve 

shows two broad peaks, at 900 Hz and at 2200 Hz. This is two degree of freedom behavior typical of a liner 

with a buried septum, where the passive piezo-electric film acts as a septum. When the piezo-electric is 

energized, the absorption increases at frequencies between the two peaks. If the velocity response of the 

piezo-electric film were greater, it is expected that the gap between the two peaks would be filled in even 

more, such that the liner would be a broadband liner, absorbing sound over the frequency range from 900 to 

2200 Hz. 

The current design of the electrodes lacks authority to affect the absorption to a greater extent than 200 Hz 

from the resonance, as seen in Figure 9. Despite this lack of authority, it is seen in Figure 10 that a 

measurable improvement in the absorption is possible in a liner configuration. Modifications to the 

electrode design have started in order to increase the control authority by at least 1 order of magnitude. 

Table 4 shows the material properties of the two encapsulated samples as supplied by Applied Physics 

Laboratory researchers. The process of encapsulation greatly improves the durability of the composite 

sample and it is expected that all future samples will be encapsulated. 

Figure 10. Calculated absorption spectrum of liner sample as depicted in Figure 9. Compares power off to 

power on to piezo-electric film. 

Table 4. Properties of Encapsulated Composite Samples 

Type Designation Color Durometer 

hardness 

Tensile @ break 

MPa (psi) 

Elongation @ 

break % 

Elasticity Modulus 

MPa (psi) 

SM JHS5 pink D40 6.8 (980) 160 3.9 (560) 

9318 JHS3 yellow A55 3.0 (440) 130 2.0 (310) 
12 

III. Summary 

Researchers have successfully completed the first phase of the project. They have developed liner samples 

with piezo-electric actuators bonded as a back wall and they have demonstrated that the energized piezoelectric can change the normal incidence impedance. The researchers found that the original design of the 

piezo-electric samples was too delicate to be used in acoustic environments greater than 120 dB. They 

revised the design to improve robustness. The encapsulate used to strengthen the piezo-electric film affects 

the stiffness of the film and, therefore its impedance. Further research is planned to assess whether the 

encapsulate has an adverse impact on control authority of the piezo-electric film and how to configure the 

encapsulate to minimize the impact. The researchers found that, while excitation of the piezo-electric with a 

dynamic voltage noticeably affects the acoustic response, the sample lacks authority to alter the impedance 

significantly. The researchers have begun plans to redesign the electroding to increase control authority. It 

should be noted that the change of impedance seen by exciting the piezo-electric was accomplished without 

control of the phase of the electrical excitation relative to the phase of the acoustic excitation. The 

researchers expect that the control authority will improve not only with better designed elctroding but also 

with control of the phase of the electric excitation. The researchers have developed mathematical models of 

the impedance of the liner incorporating the active back wall. The mathematical models use measured data, 

for which the researchers developed experimental procedures. 

Acknowledgements 

The researchers wish to acknowledge the technical staff of the NASA Langley Research Center Liner 

Technology Laboratory, in particular Carol Harrison and Jerry Lyle, for their assistance in data acquisition 

and data reduction. The researchers are grateful to the NASA Aeronautics Research Mission Directorate for 

the funding to support this research. 

References 

1. Farrar, D., Ren, K., Cheng, D., Kim, S., Moon, W., Wilson, W. L., West, J. E., and Yu, S. M., 

“Permanent Polarity and Piezoelectricity of Electrospun α-Helical (α-Amino Acid) Fibers”, Advanced 

Materials, vol 23, pp 3954-3958, 2011 

2. Jones, M. G., Parrott, T. L., and Watson, W. R., "Uncertainty and Sensitivity Analyses of a TwoParameter Impedance Prediction Model", AIAA paper 2008-2928, May, 2008 
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*C-416-10156

16. SECURITY CLASSIFICATION OF:

The public reporting burden for this collection of information is estimated to average 1 hour per response, including the time for reviewing instructions, searching existing data sources, 

gathering and maintaining the data needed, and completing and reviewing the collection of information. Send comments regarding this burden estimate or any other aspect of this 

collection of information, including suggestions for reducing this burden, to Department of Defense, Washington Headquarters Services, Directorate for Information Operations and 

Reports (0704-0188), 1215 Jefferson Davis Highway, Suite 1204, Arlington, VA 22202-4302. Respondents should be aware that notwithstanding any other provision of law, no person 

shall be subject to any penalty for failing to comply with a collection of information if it does not display a currently valid OMB control number. PLEASE DO NOT RETURN YOUR FORM TO THE ABOVE ADDRESS.

1. REPORT DATE (DD-MM-YYYY)

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