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SESAMO PROJECT
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Sensors for structural monitoring
Analysis of Sensors Technology and test
bench report
SESAMO
Sensors for Structural Monitoring
EDA Contract N° A-0931-RT-GC
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INDEX
1.
INTRODUCTION ................................................................................................................................ 7
1.1
SCOPE ........................................................................................................................................ 7
1.2
GLOSSARY ................................................................................................................................ 8
1.3
REFERENCES ............................................................................................................................ 8
2.
PZT AND MEMS SENSORS ............................................................................................................. 9
2.1
INTRODUCTION ......................................................................................................................... 9
2.2
INTRODUCTION TO PZT................................................................................................................ 10
2.3
REQUIREMENTS FULFILLMENT............................................................................................. 11
2.3.1
Frequency ........................................................................................................................... 11
2.3.2
Pulse Shape........................................................................................................................ 11
2.3.3
Actuator Dimensions and weight ......................................................................................... 12
2.4
WORKING PRINCIPLE ............................................................................................................. 13
2.5
EMBEDDING TECHNIQUE ....................................................................................................... 14
2.6
SENSORS RESPONSE SIMULATION / A QUALITATIVE ANALYSIS ............................................. 14
2.7
TEST ......................................................................................................................................... 17
2.8
SUMMARY OF WORK DONE AND PLANNING ........................................................................ 19
3.
FIBRE OPTICS SENSORS ............................................................................................................. 20
3.1
INTRODUCTION ....................................................................................................................... 20
3.2
OPTICAL FIBER BRAGG SENSORS (CONTRIBUTION OF UNIPI) ............................................... 21
3.2.1
REQUIREMENTS FULFILLMENT....................................................................................... 24
3.2.2
Fiber Bragg Grating in silica fibres ...................................................................................... 25
3.2.3
FBG RELIABILITY AND DEGRADATION ........................................................................... 27
3.2.4
FBG SENSORS IMPLEMENTATION .................................................................................. 28
3.2.5
FBG SENSOR EMBEDDING TEST .................................................................................... 31
3.2.6
FBG INTERROGATION ...................................................................................................... 37
3.2.7
SOME NUMERICAL CONSIDERATIONS ........................................................................... 38
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OTHER PHYSICAL TECHNIQUES / FBG , BENDING LOSSES
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, MIXED TECHNIQUES (NHRF
CONTRIBUTION). ................................................................................................................................... 39
3.3.1
Conventional SMF-28 Fibres ............................................................................................... 39
3.3.2
Bending loss technique ....................................................................................................... 40
3.3.3
Test ..................................................................................................................................... 47
3.4
POLYMER OPTICAL FIBER (CONTRIBUTION OF NHRF) ................................................................... 51
3.4.1
Polymer Optical Fiber Sensors for Structural Health monitoring .......................................... 51
3.4.2
REFERENCES.................................................................................................................... 62
4.
BASIC INVESTIGATIONS ON OPTICAL FIBRE APPLICATION IN SOLID ROCKET MOTORS
FOR STRAIN MEASUREMENTS ................................................................................................................ 63
4.1
INTRODUCTION ............................................................................................................................ 63
4.1.1
4.2
REFERENCES.................................................................................................................... 63
ANALYSIS MODEL FOR FAILURE MODE ANALYSIS ........................................................................... 64
4.2.1
Axisymmetric debond .......................................................................................................... 65
4.2.2
Local debond ...................................................................................................................... 66
4.2.3
Radial-axial crack ................................................................................................................ 66
4.2.4
Radial-hoop crack ............................................................................................................... 67
4.3
RESULTS OF FAILURE MODE ANALYSIS ......................................................................................... 69
4.3.1
Configuration without defects .............................................................................................. 70
4.3.2
Axisymmetric debond .......................................................................................................... 71
4.3.3
Local debond ...................................................................................................................... 72
4.3.4
Radial-axial crack ................................................................................................................ 79
4.3.5
Radial-hoop crack ............................................................................................................... 95
4.3.6
Summary of results ........................................................................................................... 108
4.4
INVESTIGATIONS OF FIBRE STRAINS ............................................................................................ 110
4.4.1
General considerations ..................................................................................................... 110
4.4.2
specific configurations ....................................................................................................... 113
4.4.3
Mechanical interaction between fibre and propellant ......................................................... 124
4.4.4
Aspects of angular sensitivity and precision ...................................................................... 127
4.5
SUMMARY AND WAY AHEAD ....................................................................................................... 129
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SENSORS EFFECTS ON MAGNETIC MICRO WIRES FOR STRUCTURAL MONITORING IN
THE SESAMO PROJECT.......................................................................................................................... 131
5.1
AMORPHOUS GLASS COATED MICROWIRES ................................................................................. 131
5.1.1
Amorphous glass-coated microwires with negative magnetostriction ................................ 133
5.1.2
Amorphous glass-coated microwires with positive magnetostriction ................................. 133
5.1.3
Amorphous glass-coated microwires with low magnetostriction ........................................ 134
5.2
SENSORS BASED ON THE SWITCHING FIELD................................................................................. 135
5.2.1
Temperature dependence of the switching field ................................................................ 136
5.2.2
Stress dependence of the switching field .......................................................................... 136
5.2.3
Frequency dependence of the switching field.................................................................... 138
5.3
BISTABLE SENSOR OF TEMPERATURE USING TC .......................................................................... 139
5.4
BISTABLE SENSOR OF TEMPERATURE USING LOW MAGNETOSTRICTION ........................................ 140
5.4.1
References ....................................................................................................................... 140
5.5
EXPERIMENTAL RESULTS FOR SESAMO APPLICATIONS ................................................ 140
5.6
3 SENSORS DESIGN AND CONCLUSION ............................................................................. 143
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1. INTRODUCTION
This document fulfils part of the WP3 task; this work package covers aspects regarding sensors in
depth. The goal is to study the technologies to be explored: mainly MEMs and fibre optics. The first
family of sensors is essentially constituted of a grid of MEMS embedded in the structure to be
monitored. Such MEMs are capable of inducing waves in the structure and by evaluating their interaction
with the structure it is possible not only to gather data regarding the stress affecting the structure but
also the status of the material. The second family consists of appropriate fibre optics structures which
are sensible to stress: they require no power but light pulses to allow data readout and this allows
embedding of such sensors even inside energetic materials (e.g. solid rocket motors). Such sensors are
thus immune to electromagnetic interference which makes them suitable for complicated applications in
environments involving strong EM fields, explosion risky environments. Some details will be given also
about alternative technologies like the magnetic microwire.
1.1 SCOPE
This document addresses the following topics to be used as reference for study and development:

PZT and MEMS sensors studies

Fibre optics sensors studies
Signal conditioning and data extraction, devoted to the required signal processing both in the optical and
electrical domain, in order to allow the effective interconnection of the MEMS and Optical Fibre sensing
systems with the following system for the interpretation of the results and the actual diagnosis of the
structural health problems.
The present document will be organized as follows. Chapter 2 will present the PZT and MEM sensors
technology, analysing their possible application to our purposes (by TESEO). Chapter 3 is devoted to
present the different type of optical fiber sensors (by UNIPI and NHRF). A model of the failure mode
analysis in SRM and an analysis of the strains on fibre sensors embedded inside the rocket propellant is
presented in Chapter 4 (by Bayern-MBDA),. Finally, possible applications of magnetic nanowires are
the object of Chapter 5 (by EDIS).
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1.2 GLOSSARY
AP
Ammonium Perchlorate
CM
Composite Material
FBG
Fibre Bragg Grating
HTPB
hydroxyl-terminated poly-butadiene
MEMS
Micro electro-mechanical system
OF
Optical Fibre
SCM
Structural Composite Material
SiCN
Silica-Carbon-Nitrogen
SRM
Solid Rocket Motors
VOC
Volatile Organic Content
1.3 REFERENCES
[REF1] SESAMO Technical and Commercial Proposal
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2. PZT AND MEMS SENSORS
2.1
INTRODUCTION
Sensors are used to record variables such as strain, acceleration, sound waves, electrical or magnetic
impedance, pressure or temperature. Sensing systems can generally be divided into two classes:
passive or active sampling. Passive sampling systems are those that operate by detecting responses
due to perturbations of ambient conditions without any artificially introduced energy. The simplest forms
of a passive system are witness materials, which use sensors that intrinsically record a single value of
maximum or threshold stress, strain or displacement. Examples of this can be phase change alloys that
become magnetized beyond a certain stress level, shape memory alloys, pressure sensitive polymers,
or extensometers. Another type of passive sensing is strain measurement by piezoelectric wafers.
Lastly, several vibrational techniques can be performed passively, such as some accelerometers,
ambient frequency response and acoustic emission with piezoelectric wafers. Active sampling systems
are those that require externally supplied energy in the form of a stress or electromagnetic wave to
properly function. A few strain-based examples of active systems include electrical and magnetic
impedance measurements, eddy currents and optical fibers, which require a laser light source. Active
vibrational techniques include the transfer-function-based modal analysis and Lamb wave propagation.
Passive techniques tend to be simpler to implement and operate within a SHM system and provide
useful global damage detection capabilities, however generally active methods are more accurate in
providing localized information about a damaged area.
Vibrations are common phenomena of mechanical structures that can be detrimental to many systems.
The vibration monitoring is a key to insuring system robustness and enhancing overall performance.
Various methodologies have been developed to measure vibrations. Laser vibrometers compare the
frequency shift between the outgoing and reflected laser beam and the corresponding vibration velocity
is evaluated. These instruments can take very accurate measurement if the measured surface is
reasonably reflective and the laser beam is properly aligned. The non-contact nature of this
measurement is the major advantage over other types of vibration measurements. However, these
instruments are too bulky, require special attention during transport and are not easy to integrate into
the products covered by this research. Another methodology measures strain induced by the
mechanical vibrations of the structure. In this case sensors are applied to the mechanical structure to be
monitored. Assuming the presence of sensors has a negligible effect on the structure behaviour, the
true strain can be measured by monitoring the electrical signals over the sensors, and relate them to
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structure vibration. Piezoresistive and piezoelectric are the most common sensor family used in this kind
of measurement.
The piezoceramic is attached to the host structure using adhesives prior to measurement. This method
works well for relatively large structures such as specimens for tensile strength tests. In addition, it is
necessary to find an optimal sensor location to have a fine measurement of themechanical structure.
The optimal sensor location is a function of structure geometry and the portion of signal to be retrieved.
It is desirable to choose a sensor location such that only the vibrations of interests are detected by the
strain sensors. The sensors have to be much smaller than the host structure. Hence, the proper way to
implement these sensor schemes on small structures is to shrink sensors in all dimensions.
Micro-electro-mechanical systems (MEMS) fabrication techniques can be useful in these situations for
their ability to build very small sensors with precise geometries. Furthermore, many sensor technologies
have been developed using specialized sensor materials, such as high-quality semiconductors, and
piezoelectric thin films. Both silicon piezoresistors and piezoelectric thin films have been used to
measure vibrations at the micro-scale, in such applications as accelerometers and resonators.
2.2
INTRODUCTION TO PZT
Certain single crystal materials exhibit the following phenomenon: when the crystal is mechanically
strained, or when the crystal is deformed by the application of an external stress, electric charges
appear on the crystal surfaces; and when the direction of the strain reverses, the polarity of the electric
charge is reversed. This is called the direct piezoelectric effect, and the crystals that exhibit it are
classed as piezoelectric crystals.
Conversely, when a piezoelectric crystal is placed in an electric field, or when charges are applied by
external means to its faces, the crystal exhibits strain, i.e. the dimensions of the crystal change. When
the direction of the applied electric field is reversed, the direction of the resulting strain is reversed. This
is called the converse piezoelectric effect
Many of today's applications of piezoelectricity use polycrystalline ceramics instead of natural
piezoelectric crystals. Piezoelectric ceramics are more versatile in that their physical, chemical, and
piezoelectric characteristics can be tailored to specific applications. Piezoceramic materials can be
manufactured in almost any shape or size, and the mechanical and electrical axes of the material can be
oriented in relation to the shape of the material. These axes are set during poling (the process that
induces piezoelectric properties in the material). The orientation of the DC poling field determines the
orientation of the mechanical and electrical axes
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The direction in which tension or compression develops polarization parallel to the strain is called the
piezoelectric axis. In quartz, this axis is known as the "X-axis", and in poled ceramic materials such as
PZT the piezoelectric axis is referred to as the "Z-axis". From different combinations of the direction of
the applied field and orientation of the crystal it is possible to produce various stresses and strains in the
crystal
If, instead of the DC field, an alternating field is applied, the crystal will vibrate at the frequency of the AC
field. If the half wavelength of the AC field corresponds to the thickness of the crystal, the amplitude of
the crystal vibration will be much greater. This is called the crystal's fundamental resonance frequency.
The crystal will also have frequencies of large amplitude whenever the thickness of the crystal is equal
to an odd multiple of half a wavelength. These are termed harmonic, or overtone resonance,
frequencies (such as 3rd overtone, 5th overtone, etc.). The largest amplitude, however, occurs at the
fundamental frequency and as the harmonic number increases the vibration amplitude decreases. A
large percentage of energy loss occurs at the two faces of a crystal.
2.3
REQUIREMENTS FULFILLMENT
2.3.1
FREQUENCY
The first step is to select an appropriate driving frequency. For a given material under test thickness, it is
ideally necessary to choose the least dispersive driving frequency, which generally exists where the
slope of the phase velocity curve is equal to zero. This is because at low frequencies, the dispersion
curves have steep scope and thus are very sensitive to small variations in frequency. The higher the
frequency the smaller the slope of the dispersion curves. The wave velocities are also much faster at
higher frequencies, increasing data acquisition requirements.
The natural frequencies of the structure play a small role in the amplification or attenuation of the
transmitted wave, whereas the wave can travel with fewer disturbances at a resonant frequency.
In literature are available some frequency range for various materials. In particular for composite
materials a start-up range is from 15kHz to 40kHz. It's possible to expand this range up to ultrasonic
frequency (up to 400kHz).
2.3.2
PULSE SHAPE
The second set of variables explored was the actuation pulse parameters. These included the pulse
shape, amplitude and number of cycles to be sent during each pulse period. These parameters are
changed experimentally in relation to the generated effects.
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Below are the waveforms more commonly used for stimulation of actuator: sine wave, square wave,
exponential wave. About research of cracks, and Non-Invasive-Analysis of composite material, the most
suitable are sine wave and square wave. Sine wave, and sine wave with application of Hanning window
are particularly used with Lamb wave analysis. In other traditional analysis is commonly used a square
wave.
Once the driving frequency and signal shape have been selected, there is then a trade between the
number of waves that can be sent in an actuating pulse e and the distance from abrupt features in the
structure. The number of cycles of a periodic function desired to actuate the piezoelectric actuator is one
of the more complicated decisions to be made for all kind of analysis. An appropriate number of cycles
can be determined by the maximum number of waves that can be sent in the time it takes for the lead
wave to travel to the sensing PZT patch. It is also convenient to use intervals of half cycles so that the
sent sinusoidal pulse becomes symmetric. Research from the literature has used signals varying from
3.5 to 13.5 cycles per actuating pulse. Since the specimens in the current research are relatively short,
few cycles could be actuated without disturbing the received signal; thus 3.5 cycles were used to drive
the piezoceramic actuators. Lastly, by increasing the driving voltage, the magnitude of the strain
produced by the propagating Lamb wave proportionately increases. Driving voltage can vary from few
volt up to 100 V. Increasing the amplitude also increases the signal to noise ratio to yield a clearer
signal. Also, a SHM system should be as low power as possible, thus the voltage should be chosen to
be the minimum required to resolve the desired damage size. The response of sensor can be few mV
for PZT or ICP for small built-in accelerometer.
2.3.3
ACTUATOR DIMENSIONS AND WEIGHT
PZT piezoceramic actuators were chosen for the present research due to their high force output at
relatively low voltages, and their good response qualities at low frequencies. The generic shape of the
actuator should be chosen based upon desired propagation or reception directions. Several researchers
in various fields have examined the effects of piezoelectric wafer dimensions on the efficiency of their
actuation. Waves propagated parallel to each edge of the actuator, i.e. longitudinally and transversely
for a rectangular patch and circumferentially from a circular actuator. The width of the actuator in the
propagation direction is not critical, however the wider it is, the more uniform the waveform created.
For PZT sensor and actuator, the weight normally is not a problem because are very light (1 - 20
grams), and the maximum weight of the sensor/actuator must be less than 5% of the sample total
weight. For ICP accelerometer weight is a not negligible parameter, ICP sensor can weigh up to 200
grams.
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WORKING PRINCIPLE
The verification of the actuators and sensors will be approached initially in the following ways (see
Figure 1):
In plane pitch-catch
Receiver
Transmitter
Demaged Region
Drawings about different configurations:
Figure 1 Actuators and sensors setup scheme, in plane pitch-catch
Using an actuator (vibration generator) type PZT will be stimulated the specimen, and a vibration sensor
PZT will execute the relief of vibrations at the other end of the specimen. The sensor response will be
characterized as a function of actuator position, then the measurement result will be characterized as a
function of sensor position.
This characterization will be repeated using a MEMS accelerometer sensor in place of the PZT.
At the end of this first phase of characterization, this characterization will be repeated, using a single
PZT (actuator/sensor) (see Figure 2)
In plane pulse-echo
Transmitter
Receiver
Crack
Figure 2 Actuators and sensors setup scheme, in plane pulse-echo
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EMBEDDING TECHNIQUE
Typical embedding techniques are:

Beeswax

Cyanoacrylate glue (Loctite)

Mechanical locking

Resin
Beeswax is a simple way to lock the sensor/actuator, but with some temperature limit. Temperature
range for common beewax is +10 / +50°C, a tipically use is in the laboratory, because the temperature is
controlled and it’s easy to remove.
Cyanoacrylate glue had an extended range temperature but it’s harder to remove than beeswax.
Mechanical locking is the strongest way, but it’s also very invasive.
Resin is a good compromise between "blocking strength" and is removable. This system it will be the
first way in our tests.
2.6
SENSORS RESPONSE SIMULATION / A QUALITATIVE ANALYSIS
The results of a PZT sensors case study e Lamb-wave analysis were analyzed: “In-Situ Damage
Detection of Composites Structures using Lamb Wave Methods”
The first set of experiments was conducted on narrow composite coupons. The laminates were 25 x 5
cm rectangular quasi-isotropic laminates of the AS4/3501-6 graphite/epoxy system with various forms of
damage introduced to them, including matrix-cracks, delaminations and through-holes. PZT
piezoceramic patches were affixed to each specimen using 3M ThermoBond thermoplastic tape. Both
the actuation and the data acquisition were performed using a portable NI-Daqpad 6070E data
acquisition board, and a laptop running Labview as a virtual controller. A single pulse of the optimal
signal was sent to the driving PZT at 15 kHz to stimulate an A0 mode Lamb wave, and concurrently the
strain induced voltage outputs of the other two patches were recorded for 1 ms to monitor the wave
propagation. The results were compared by performing a wavelet decomposition using the Morlet
wavelet, and plotting the magnitude of the coefficients at the driving frequency.
This procedure was also carried out for beam specimens with various cores at a driving frequency of 50
kHz. Further experimentation examined damage in more complex built-up specimens. Laminated plates
were tested with ribs that were bonded across the center of each plate using Cytec FM-123 film
adhesive.
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Different configurations included 25 mm wide aluminium C-channel rib and a composite doubler with
and without a centre delamination, both using a driving frequency of 15 kHz. Next a sandwich
construction cylinder with a 40 cm diameter and length of 120 cm was tested. It had two face-sheets of
similar layup as the other tested specimens, and a 25 mm thick low-density aluminium honeycomb
bonded between them. Piezoceramic patches were placed down the length of the cylinder every 10 cm
for 60 cm in several control regions as well as in a region with visible impact damage. The driving
frequency for this test was 40 kHz because of the honeycomb core and slightly different lay-up.
Detailed results for each of the experiments described can be found in previous papers focusing on
Lamb wave experimentation. A few key results are shown below . Results from the narrow coupon tests
clearly showed the presence of damage in all of the specimens.
The most obvious method to distinguish between damaged and undamaged specimens was by
regarding the wavelet decomposition plots, show in Figure 3, where the control specimens retained over
twice as much energy at the peak frequency as compared to all of the damaged specimens. The loss of
energy in the damaged specimens was due to reflection energy, and dispersion caused by the microcracks within the laminate in the excitation of high-frequency local modes. Probably the most significant
result of the present research was the “blind test.” Four high density aluminium-core beam specimens
were tested, one of which had a known delamination in its centre, while of the remaining three
specimens it was unknown which contained the circular loosening and which two were the undamaged
controls. By comparing the four wavelet coefficient plots in Figure 4, one can easily deduce that the two
control specimens are the ones with much more energy in the transmitted signals, while the third
specimen (Control C) obviously has the flaw that reduces energy to a similar level to that of the known
delaminated specimen. This test serves as a testament to the viability of the Lamb Wave method being
able to detect damage in at least simple structures. Similar effects of damage were observed in each of
the built-up composite structure cases. By comparing the stiffened plates results, a reproducible signal
was transmitted across each of the intact portions of the composite stiffeners while it was obvious that
the signal travelling through the delaminated region was propagating at a different speed. Finally, by
comparing the axial wave propagation in the control and damaged regions of the cylinder, it could be
seen that the impacted region caused severe dispersion, which attenuated the received signal at each
sensor further down the tube. For all of the tested specimens, damage was easily perceived by
comparing the wavelet coefficient magnitudes for the control versus damaged signal.
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Figure 3: Wavelet coefficient for thin coupons; compares 15kHz Energy content.
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Figure 4: Wavelet coefficient for beam “blind test”; compares 50kHz energy content
2.7
TEST
The test sequence will be structured as describe in § 2.3.
This test sequence will be repeated for all kind of sensor/actuator. After this measures campaign, will
be analyzed all data to define the best way to find structural problems on a canister. Then will be
execute a series of measures on a canister.
Here are some data sheets of families of sensors and PZT actuators.
The sensor and the actuator will be chosen from one of these families.
PIEZOELECTRIC SINGLE SHEETS:
Thickness
:
0,127 – 2,03 mm
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Dimension
:
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72,4 X 72,4 mm
Capacitance :
40 – 650 nF
Composition :
PZT
PIEZOELECTRIC SINGLE DISK:
Thickness
:
0,191 mm
Diameter
:
3,2 X 63,5 mm
Capacitance :
0,65 – 265 nF
Composition :
PZT
2-PIEZO LAYER TRANDUCERS BENDERS / EXTENDERS:
Thickness
:
Total: 0,38 – 0,66 mm (Ceramic 0,13 – 0,27 mm)
Figure 5
Width
:
3.2 x 31.8 mm
Length
:
31.8 X 63.5 mm
Capacitance :
2.5 – 640 nF
Polarization
:
series / parallel
Frequency:
63 – 440 Hz Benders; 0.5 – 30 kHz Extenders
Composition :
A4:
Thin vacuum sputtered nickel electrodes produce extremely low current
leakage and low magnetic permeability. It operates over a wide temperature range and is relatively
temperature insensitive.
H4 :
It has a high motion/volt and charge/newton rating, which is useful when voltage or force is
limited. Thin vacuum sputtered nickel electrodes produce extremely low current leakage and magnetic
permeability. However, its temperature range is limited and its properties are more sensitive to
temperature
A3: Its has totally non-magnetic, fired-on silver electrodes, operates over a wide temperature range, and
is relatively temperature in sensitive
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2-PIEZO LAYER PIEZO BENDING DISKS:
Thickness
:
0,41 mm
Diameter
:
3,2 X 63,5 mm
Capacitance :
0,3 – 107 nF
Composition :
A4
Frequency
0.29 – 116 kHz
2.8
:
SUMMARY OF WORK DONE AND PLANNING
The first stage of the research was to define the technology of sensor/actuator and the survey
methodology. At the same time we have selected the target on which to search, the canister and then
the composite materials.
At this point a short list of possible sensors/actuators to be used has been defined. In parallel we have
started the procedures for the retrieval of specimens.
Once acquired the necessary equipment, the phase of laboratory tests on specimens will begin.
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3. FIBRE OPTICS SENSORS
3.1
INTRODUCTION
Field of applications of sensors (stress/strain measurement): structural health monitoring of composite
materials and Solid Rocket Motors .
The purpose of SESAMO is monitoring the health of both Composite Materials (CM) of aerospace
interest, employed also as missile canister, and of rocket solid propellant grain. At the present state of
the art, it is necessary to develop new kind of sensors to have a better understanding of the complex
nature of fatigue and ageing effects acting on CM and SRM. Sensors should allow to get data both
during material preparation (i.e. CM lamination, resin curing, and so on) and when the
structures/substances are in storage. The stability and reliability of these sensors must permit an
accurate tracking of the physical changes over the manufacturing procedure and during the operative
life of CM and SRM, which can be more than 10-20 years long. Currently missile canisters and
propellant grains are exposed to situations of poor controlled storage and to extreme environmental
variations, in terms of temperature, mechanical shocks and vibrations. Moreover, to reduce ownership
cost, SRM systems are often maintained in service for prolonged periods of time, even beyond their
initial designed life. Specially about SRM, accurate predictions are difficult to attain, being compromised
by both the complex loading and the ageing of the propellant grain. Knowledge of the environmental
load that could induce modifications in the SRM grain is required to verify structural integrity.
Manufacturing process, storage, transportation and handling, impose severe threats to the integrity of
the materials, whose characteristics depend on mechanical loads and thermal cycles. To achieve safe
CM and SRM operation, the recording of some physical parameters is essential. Parameters sampling
rate may vary in dependence of the process under investigation, ranging from a quite high rate needed
during both manufacturing and deployment, to a very low rate during warehouse storage. About the
information quality, the recovered data should provide not only a simple go/no-go threshold (intended as
the minimum level knowledge of the situation), but preferentially should allow a spatial reconstruction
and a quantitative estimation of the damage. The SESAMO project is in progress just to establish which
kind of embedded sensor can be better suitable to determine position and magnitude of such structural
defects, keeping into account the complexity of the stress distribution field that arises especially within
solid propellant grain because of environmental loads.
In the project, both the application of MEM and OF sensors will be considered and analysed.
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OPTICAL FIBER BRAGG SENSORS (CONTRIBUTION OF UNIPI)
Optical fibre sensors have been successfully employed during the last two decades in the field of
structural health monitoring. Various physical parameters (strain, temperature, humidity, pressure) can
be kept under investigation by using both localized single-point sensors and quasi-distributed network
arrays. Optical fibres can be easily integrated within structures because of their small dimensions (125
to 250 micron diameter for conventional Telecom silica single-mode fibres) and relative flexibility, which
makes possible to wind them down to small bending radius (~ 10 mm) to follow curve surfaces. The
small mass and immunity to electromagnetic interference, also make them attractive for sensing
applications.
Light is guided in optical fibres (OF) by means of total internal reflection. OF typically have a high
refractive index core surrounded by lower index cladding. Because of many different possible reflection
pathways for the travelling light rays, interference occurs at the arriving point. Depending on the fibre
diameter, only a finite number of transmission modes, called spatial mode, can be supported. Singlemode fibre (SMF) are fibres whose diameter is so small that just one transmission mode is allowed.
SMF C-band (1520-1580 nm) Telecom fibres have a silica core diameter of 9-10 m. It is conveniently
doped in order to increase its refractive index in comparison with the surrounding silica cladding. The
concentric duct of core plus cladding has a typical diameter of about 125 m. Finally, a protection
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Figure 6 Fibre structural composition
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Figure 7 Silica fibre attenuation vs. wavelength. ZBLAN is a heavy metal fluoride glass with greater IR
transmission.
coating (usually acrylate or polyimide) is applied to reach a total diameter of 250 m. Around 1550 nm,
the attenuation length is of many km (about 1 dB/km) and silica OF present a very wide information
bandwidth, of the order THz .
Polymer Optical Fibres (POF) is another fibre family which has become popular in the last years in some
kind of sensor applications. Plastic fibres can sustain much more large strain than silica fibres (up to 3040 % in comparison with a few percent only) and for this reason they can be applied to monitor large
scale deformation events with lower breakage risk. PMMA and Polystirene are used as fibre core, while
generally the cladding is made of silicon resin. A high refractive index difference is then maintained
between core and cladding, giving the opportunity of a high numerical aperture. Both single-mode and
multi-mode plastic fibres are available.
An obvious remark for all types of embedded OF sensor is that the fibre must maintain its integrity while
in use. Before considering whatever is the most suitable physical principle of measurement (elongation,
OTDR, FBG, bending loss), it is necessary to evaluate the maximum solicitation in terms of stress and
temperature to which the OF is subdued. The main drawback with POF is that red visible light source
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must be used, preventing the utilization of standard low cost C-band optical components. Moreover, the
attenuation length in POF is substantially shorter than silica fibres (of the order of 100 m).
3.2.1
REQUIREMENTS FULFILLMENT
To establish whether an OF sensor can be successfully employed in measuring CM and SRM structural
loads, a series of general requirements must be fulfilled. Maximum strain and temperature endurance,
strain and temperature operating ranges, life time, sensitivity (minimum detectable signal), responsivity
(magnitude of the physical observable corresponding to an unitary change of the parameter under
measurement), accuracy, reliability, sampling rate, number of sensors needed to obtain the information
on a defined volume within a specified spatial grid (sensor size – structure mesh size comparison ).
Besides, the dimension, the weight of the whole measurement equipment (a part from the sensors
themselves), its power budget and total cost may be parameters to keep under consideration.
The table underneath is a summary of the principal sensor requirements, relatively to a series of
sensitive parameters as it is indicated in the D1 report. The numerical values corresponding to the
various items must be integrated with those coming from the Bayern Chemie (BC) which are included in
the present document.
In particular, for the monitoring of solid propellant grains, as can be gathered from the BC report, large
strain arises close to the bore of the booster. Its amount can reach 20% in the hoop direction,
outperforming the strain capability of silica fibres which are known to posses a breakage point in the
region between 5-10% , that is 50-100 m, where 1 m (“millistrain”) is defined as the elongation of 1
part over 1000. Silica fibre breakage point is strongly dependent from the characteristics of the individual
fibre, and from the fibre handling. For example, the operation of removing the coating in order to expose
the fibre silica cladding, may weaken the fibre structure (ref. Limberger et al., Proc. of SPIE, vol. 2841,
pg. 84, 1996). It should be possible to overcome the problem of excessive load positioning the fibre
along directions of lower strain, under the hypothesis that this configuration is effective to measure
debonds and cracks within the grain. Another solution is to employ POF, if they can assure
performances of the same level and if their embedding can be carried out within the solid propellant
without integrity risk for the fibre itself and without affecting the properties of the rubber-like resin
constituting the propellant bond.
Besides the already cited advantages (small mass and size, immunity from electromagnetic
interference), silica OF sensors, and in particular polyimide coated silica OF, can operate in harsh
locations, and in situations where the use of electrical sensors would be impractical or absolutely
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impossible (acid environment, high temperature, high humidity). OF sensors can be multiplexed to
constitute an array suitable for a quasi-distributed monitoring, accessed by means of a single optical
link. In this way, if necessary, all the electro-optical equipment, but the sensors themselves, can be
remotely placed, far from the sample under test. OF sensor implementation to monitor CM ans SRM in
agreement with the expected performances, relies basically on two different physical principles and
techniques: Fibre Bragg Grating (FBG) and Bending Losses (BL).
3.2.2
FIBER BRAGG GRATING IN SILICA FIBRES
FBG sensors are in-fibre spectral filters based on the Bragg reflection law. A series of close parallel
lines (the grating) are printed by means of a lithographic method onto the fibre core, inducing a periodic
modulation of the refractive index. The production process is usually realized employing photosensitive
fibres and UV laser sources by means of the so called Phase Mask (PM) Technique. Fibre silica core is
lightly doped from the manufacturer adding a few percent of Ge and/or B (this doping procedure is
common in fibre manufacturing); in this way the fibre core sensitivity to absorb UV radiation in the 240260 nm band is greatly enhanced (photosensitivity). The UV laser beam is shaped in form of a tin strip
by a cylindrical lens and passes through a phase mask, which is a quartz substrate carrying a dielectric
layer representing the modulation pattern to be imaged into the fibre core.
When broadband light is sent through the fibre, only the fraction around the Bragg wavelength is
reflected, satisfying the Bragg condition:
 = 2 n 
where n is the average (effective) core refractive index, and  is the grating period (equal to half the
value of the PM period).
For some fixed grating length, the attainable peak reflectivity is essentially function of the amount of
refractive index induced variation, which in turn depends on the UV radiation fluence, that is laser
intensity and fibre exposure time. Practically it is not difficult to obtain peak reflectivity in excess of 95%
in few minutes of fibre irradiation. The reflection bandwidth (FWHM) is of the order of 0.1 – 0.3 nm in the
case of 10 – 12 mm grating length.
The Bragg wavelength can be set spectrally everywhere in the Telecom C-band by an appropriate
choice of PMs and/or applying a controlled strain to the fibre during the inscription process. This is just
the physical principle which makes possible using FBG as a strain sensor: when a mechanical strain is
applied to this structure, both the grating period and the refractive index undergo to a change (elastooptic effect). Thus a variation of the Bragg wavelength is induced, and a measurement of this change
allows a monitoring of the strain. The typical wavelength shift due to strain (which is defined as strain
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Figure 9 Fiber Bragg Grating
responsivity) is 1.2 pm/strain. It is important to point out that just for their nature, FBGs are also
strongly sensitive to temperature, not only through the expansion coefficient of the core/cladding
material (acting on the grating period  , but also because the thermo-optic effect induces variations on
the refractive index. Typical temperature responsivity is about 10.2 pm/oC.
3.2.3
FBG RELIABILITY AND DEGRADATION
Pristine silica fibre can sustain strains up to 10 m, however fibre handling during FBG inscription
process substantially reduces their mechanical strength. Most commonly used fibre coatings, as UVcurable acrylate or polyimide, do not transmit wavelengths where OF are photo-sensitive. Therefore,
these coatings must be mechanically or chemically removed in order to write FBGs into the fibre core.
Mechanical operations, carried out with stripping tools, are intrinsically more invasive than chemical
ones, and fibre damage is by far more probable. Thus chemical coating removal is generally preferred to
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maintain the integrity of the fibre, even if it can result less practical in certain operative situations. FBG
inscription itself can contribute to decrease the mechanical strength of the fibre (Limberger et al. 1996),
and for this reason it is preferable to employ photosensitive fibres that need shorter UV exposure time
for the same peak energy during grating writing. In any case the fibres specifically declared and sold as
photosensitive from the manufacturer generally require a level of dose irradiation which is safe for fibre
structural integrity. A lot of studies have been carried out about FBG performance under environmental
loads. At room temperature the grating is substantially stable, that is the reflectivity and the Bragg
wavelength remain unchanged for long time (10-20 years). Exposure to high temperature and/or to
repetitive high strains, induces variations on FBG parameters. Anyway, after few tens hours at 200 0C,
the grating reflectivity saturates to about 80% of its maximum, making the effect not dramatic over the
required performance. FBGs are also sensitive to water exposition (humidity) and for this reason a recoating of the grating region is to be considered after fabrication, whenever it should be employed in wet
environment.
3.2.4
FBG SENSORS IMPLEMENTATION
A strain over the fibre induces variations of its refractive index n (elasto-optic effect) and of the grating
period  changing the FBG peak reflection wavelength according to the Bragg condition  = 2 n  In
general strain and temperature act simultaneously on the grating, in such a way that the following
relation holds:
T
where  ≈ 0.78 is the strain optic gauge factor, depending on the refractive index, on the Poisson's ratio
for the fibre, and on the strain optic tensor components.
In this relation n, where nisthe thermo-optic coefficient and  represents the expansion
coefficient for the fibre material.
However for silica  ≈ 10-6 / 0C, thus for  = 100 0C the induced equivalent strain is of the order of 100
 only, and the thermal expansion contribution can be neglected in most cases.
Considering the temperature T as a constant, 1 m causes a wavelength changing of about 1.2 nm at 
= 1550 nm. The typical FBG reflection bandwidth is of the order of few tenths of nm; however, in
projecting an array of FBG strain sensors, a suitable wavelength spacing must respected among
individual sensor channels, in order to distinguish their response without spectral overlap. Supposing for
each FBG sensor a dynamic range of 5 m, that is about 6 nm, it is evident that only about ten sensing
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channels can be allotted within the Telecom C-band. This number of channels can be increased by
using an optical switch, addressing in different times the interrogation of each sensor array. In
conclusion, a single equipment can perform the interrogation of tens of sensors at a scanning rate up to
several kHz.
Irrespective of the method chosen to interrogate the single FBG sensor, the problem of separating strain
and temperature effects arises. Several different solutions have been proposed to fix this item:
positioning of one reference FBG in a location where it could be insensitive to strain, while subjected to
the same temperature excursions sensed by the other gratings; using of two different wavelength bands
to interrogate the same FBG in order to benefit of the spectral dependence of the parameters; a-priori
knowledge of a different time scale (if any) for strain and temperature evolution; deployment within the
same specimen under monitoring of FBG coupled to Long Period Grating (LPG), having grating periods
of hundreds micron, and whose central reflection wavelength behaviour with temperature and strain is
different from the FBG response: temperature sensitivity 5-10 times higher and strain sensitivity 2 times
lower than FBG. That is because the LPG peak wavelength shift caused by temperature and/or strain is
proportional to grating period multiplied by the difference in refractive index between the core and the
cladding (ref. H.J. Patrick et al., IEEE Ph. Techn. Lett., vol. 8, n. 9,1223-1225; 1996). Use of FBG
written on Er:Yb co-doped fibres (J. Jung et al., App. Opt., vol. 39, n. 7, 1118-1120; 2000) in order to
benefit of the transmitted (or reflected) Amplified Spontaneous Emission dependence on temperature: a
simultaneous measurement of optical power and of wavelength shift by means of an Optical Spectrum
Analyser (OSA), allows to solve a linear matrix equation for strain and temperature values, with errors of
few tens of  and about 3 °C respectively.
Another method rests on enhancing the sensitivity to strain by chemical etching one half of the length of
the FBG (S.K. Mondal et al., Rev. of Sci. Instr., 80, 103106; 2009). The strain response for the pristine
and for the etched FBG parts is different because of the difference in diameter, while the temperature
sensitivity is the same. Maximum errors of ± 13  and ± 1 0C are reported over 1.7 m and 60 0C
measurement range. See Figure 10 .
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Figure 10 FBG-1 and FBG-2 denotes respectively the non-etched and the etched section of a 10 mm long
FBG. The pristine fibre diameter is 125 m; the etched fibre diameter is about 61.75 m.
It is worth noting that for the SESAMO goal of investigating the stress field within a solid propellant, we
are interested in knowing locally (with a spatial resolution of a few mm and a strain resolution of the
order of 100 corresponding to L = 0.5 m over L = 5 mm) the mechanical load which is principally
induced by thermal cycling. During the propellant resin curing, the temperature is held constant at 60 0C,
thus after a time transient, the whole volume reaches a steady temperature state. In the first step of the
process, FBG sensors are subdued to temperature changing within the material, and only after the
attainment of the thermal equilibrium they are effective in measuring the possible formation of debonds
and cracks. If the temperature transitory time is short enough in comparison with the generation of the
grain defects, it is not important to distinguish the separated temperature and strain contributions to the
Bragg wavelength shift. It is not trivial at all to apply the heat transfer equations to this case, in order to
deduce the time scale of the phenomenon, because we are dealing with a composite material whose
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parameters are largely unknown for the calculus under consideration. However it makes sense thinking
of the propellant compound as a rubber-like material, having a quite “low” thermal conductivity. Thus, if
we really want to perform pure strain measurements, it is not sufficient to employ conventional FBG
sensors. Differently, when considering long term measurement on warehouse stored propellant, it is
possible to use an environmental temperature sensing (nearby the material under investigation) as a
reference, and that is enough to extrapolate strain information from embedded FBG sensors. An
obvious remark to this last statement, is that no abrupt temperature variation should occur during the
observation time.
Another item to deal with, is the silica fibre stiffness. Silica Young modulus E is about 73 GPa, while for
rubber-like compounds this value is by far lower. Embedding silica OF within a soft host material two
kinds of issues should be addressed: whether the fibre coating could slip within the substance to
monitor, limiting the strain transduction from the material to the FBG, and whether the silica OF could
sense at all any strain effect because of the enormous difference in stiffness. Anyway, this last point
seems to have been already fixed in literature by several experimental works that demonstrated the
feasibility and effectiveness of silica OF sensors embedded in resins. (i.e. D. Karakelas, Rapid Prot. J.
14,2 (2008) 81-86). Embedding of OF sensors into plastic laminates or composite material is nowadays
a quite common practice, giving very good results in measuring shape deformation.
MAYTECH has already been able to perform this kind of process using carbon-fibre cylinders as host
elements. In the next months, UNIPI, NHRF and MAYTECH will cooperate to an experimental test
involving FBG sensors embedded within composite material canister.
3.2.5
FBG SENSOR EMBEDDING TEST
We performed a simple functional test embedding a section of silica fibre containing a FBG within a twocomponent filler resin used in building maintenance. The FBG (developed within a silica photo-sensitive
fibre) is 10 mm long and has a peak reflectivity of about 99 % centered at 1542.55nm. The OF section
is not disposed within the mould in straight line, but instead in the shape of a three turns helix, with a
pitch of about 30 mm and a curvature radius of about 25 mm. The helix axes is aligned with the axes of
the cylindrical mould.
The resin has an exothermic reaction during the very first period of curing, producing a temperature
increase up to 50-60 °C in few minutes. Afterwards, it reaches a steady state in about 24 hours with
nominally no volume variation. We illuminated the FBG by means of a broad-band diode and recorded
its transmission spectrum onto an optical spectrum analyser at various curing stages.
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Starting from the condition showed by Figure 11, the FBG is evidently subdued to dishomogeneous
strain along its length with the appearance of a multiple peak structure, which makes difficult to deduce
what could be the effective total wavelength shift of the process. It is worth noting that the FBG
reflection peak behaviour shows a compressive strain (Figure 12-Figure 14), indicating the raise up of
stress due to shrinkage. In the final state, (Figure 15) two main peaks are recognizable, making difficult
to address the overall wavelength (and strain) measurement shift through a comparison with the initial
condition of Figure 11.
However this test has been performed with a very peculiar resin, constituting a border line case because
no control on curing conditions was attempted. Besides, the formation within the resin compound of
quite large air bubbles and macroscopic cracks, indicates very poor homogeneity during solidification.
This test is anyway indicative of the FBG mechanical strength, because it could accommodate a
wavelength variation of several nm (equivalent to several me) without breakage. Moreover, the
experiment shows that a particular care must be taken in order to choose the best geometrical shape of
the OF which sustains the FBG to allow an optimal strain transduction from the host material to the
sensor.
Figure 11 FBG spectrum just before resin pouring
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Figure 12 Spectrum after 2 min resin curing. The peak wavelength shift is consistent with a 20 °C
temperature increase.
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Figure 13 After 20 min resin curing. Positive strain is growing up together with a spectrum deformation.
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Figure 14 After about 1 hour resin curing. Negative (compressive) strain is evident. Multiple peaks
appearance due to not homogeneous resin curing along FBG length.
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Figure 15 After about 24 hours resin curing (final steady state). It is not clear which peak could be
considered to measure the wavelength shift.
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FBG INTERROGATION
As extensively discussed, FBG strain sensor principle rests on a wavelength shift measurement (Bragg
wavelength). Several methods can be applied to this purpose, based on commercial instruments or
dedicated devices. Probably the most performing and reliable FBG interrogation system counts upon a
fibre-coupled continuous tunable diode laser source, which gives a monochromatic radiation with an
emission spectrum narrower than the FBG reflection band. The laser light wavelength is scanned
through the various FBG of an array, and the information is recovered provided that a spectral
calibration of the source itself has been previously done. The attainable wavelength resolution can be of
the order of 0.1 pm, well beneath the equivalent discrimination capability of a FBG strain sensor. This
devices are very expensive and more suitable to carry on laboratory experiments than field tests. The
employ of interferometric systems is quite cumbersome, although it allows very precise measurements.
An in-fiber Michelson or Mach-Zender interferometer must be built, respecting an arms unbalance not
exceeding the coherence length of the light source used to illuminate the FBG sensors. In this way any
wavelength variation can be converted into a phase changing, with a gain factor proportional to the
optical path difference (if any). Without going into detail, and aside from the opto-electronic complexity
of the apparatus, the system results inherently not affordable in measuring quasi-static phenomena, just
because the interferometer itself is particularly sensitive to mechanical vibrations, acoustic waves and
temperature excursions. Once again interferometric methods are largely used in laboratory test bench,
but seldom employed in current use. The most popular FBG strain sensors interrogation method is by
far based on low cost commercially available instruments, whose working principle consists in a spectral
dispersive element in conjunction with InGaAs photo-detector arrays. It is sufficient to illuminate the FBG
sensors with a in-fibre broad-band light source (as a super-luminescent diode), and to address the
reflected signals to the instrument, using the same fibre after backward decoupling by means of a
standard optic circulator. In terms of resolution, accuracy, repeatability, and handiness, these devices
are well suited to SESAMO requirements. In fact they allow multiple sensors monitoring with low budget
in a smart way and can be utilized by not-skilled people just after a short training. Half-way in terms of
easiness of usage, equipment complexity, cost, and attainable performance, there is another class of
instruments based on Electronically Tuneable Optical Filter (ETOF). In this case, the light from a broadband source is spectrally scanned by an ETOF device before to be sent to illuminate a FBG array. By
timing the arrival of the Bragg reflection peaks from each FBG with respect to the start of a filter scan,
the exact wavelength is deduced.
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SOME NUMERICAL CONSIDERATIONS
Considering a typical FBG length of 5 mm, its reflectivity attains a maximum of 90 - 95% at the Bragg
wavelength (within the 1525 – 1565 nm C-band), with a spectral band-width of about 200– 300 pm.
Because of the relation  = 1.2 pm/, and supposing 100 – 200 of spectral resolution in detecting
the centre of the FBG reflection band, we get a strain resolving power around 1 . The particle size
within the grain of a missile solid propellant, can reach 200 m and more, thus the spatial scale over
which a stress-strain definition makes sense for the whole compound is of the order of several mm. This
is just the suitable length to allow FBG strain sensor measurements. Besides, a strain resolution of the
order of 100  seems to be sufficient in order to monitor micron size debonds and cracks formation
nearby the position of an individual FBG.
It is worth noting that, whether the use of silica fibre is planned, the strain dynamic range should not
exceed about 5 m, to maintain a wavelength linear relationship and to remain well under any breakage
risk.
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OTHER PHYSICAL TECHNIQUES / FBG , BENDING LOSSES , MIXED TECHNIQUES
(NHRF CONTRIBUTION).
3.3.1
CONVENTIONAL SMF-28 FIBRES
Corning SMF-28 single-mode fibre is considered the “standard” optical fibre for telephony,
cable television submarine, and private network applications in the transmission of data, voice
and/or video services (Table I).
Table I
Geometrical Properties
Numerical Aperture
0.14
Cladding Diameter (μm)
125±0.7
Core Diameter (μm)
8.2
Core/Cladding concentricity error (μm)
5
Coating Diameter-Acrylate (μm)
145±5
Coating Diameter-Polyimide (μm)
155±5
Mechanical Properties
Operating temperature range-Acrylate (ºC)
-65 to +85
Operating temperature range-Polyimide (ºC)
-65 to 300
Coating Strip Force (N)
3
Fiber proof tensile test level (GN/m2)
Dynamic Fatigue Resistance Parameter (nd)
≥0.7
20
Optical Properties
≤1260
Cut-off wavelength (nm)
Mode-field Diameter at 1550 nm (μm)
Zero Dispersion Wavelength (nm)
10.4±0.8
1302-1322
Attenuation at 1310 nm (dB/km)
≤0.35
Attenuation at 1550 nm (dB/km)
≤0.22
Refractive Index Difference-Core/Cladding (%)
0.36
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Corning SMF-28 fibre is manufactured to the most demanding specifications in the industry.
SMF-28 fibre is optimized for use in the 1310 nm wavelength region. The information-carrying
capacity of the fibre is at its highest in this transmission window, and it is also where dispersion
is the lowest. SMF-28 fibre also can be used effectively in the 1550 nm wavelength region.
Corning’s enhanced, dual layer acrylate CPC6 coating provides excellent fibre protection and is
easy to work with. CPC6 can be mechanically stripped and has an outside diameter of 245 μm.
Other coating such as polyimide can also be used which makes the fiber extremely robust and
durable under harsh environments at elevated temperatures.
3.3.2
3.3.2.1
BENDING LOSS TECHNIQUE
BENDING LOSS PROPERTIES OF CONVENTIONAL SMF-28 FIBRES.
Optical fibres suffer from macro-bending loss at bends or curves on their paths. This is due to
the energy in the evanescent field at the bend exceeding the velocity of light in the cladding and
hence the guidance mechanism is inhibited, which causes light energy to be radiated from the
fibre.
This is shown in the following illustration together with the refractive index distribution of straight
and bent fibre.
Figure 16
The outside of the turn is strained (the refractive index decreases) and the inside of the turn is
stressed and (the refractive index increases). The part of mode which is on the outside of the
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bend is required to travel faster than that on the inside so that a wavefront perpendicular to the
direction of propagation is maintained.
So part of the mode in the cladding needs to travel faster than the velocity of light in that
medium. Because this is impossible, the energy associated with this part of the mode is lost
through radiation.
Typically, these losses rise very quickly once a certain critical bend radius is reached. This
critical radius can be very small (a few mm) for fibres with robust guiding characteristics (high
numerical aperture), whereas it is much larger (often tens of cm) for single-mode fibres with
large mode areas. In other words,
Generally, bend losses increase strongly for longer wavelengths, although the wavelength
dependence is often strongly oscillatory due to interference with light reflected at the
cladding/coating boundary, and/or at the outer coating surface. The increasing bend losses at
longer wavelengths often limit the usable wavelength range of a single-mode fibre. For
example, a fibre with a single-mode cut-off wavelength of 800 nm, as is suitable for operation in
the 1-μm region, may not be usable at 1500 nm, because they would exhibit excessive bend
losses. Note that even without macroscopic bending of a fibre, bend losses can occur as a
result of micro-bends, i.e., microscopic disturbances in the fibre, which can be caused by
imperfect fabrication conditions.
The power loss L (dB/m) coefficient can be expressed for the LP 01 mode in terms of familiar
fibre parameters as
where a is the fibre core radius, R is the bending radius, V number embodies the fibre structural
parameters and frequency V=a(2π/λ)(nco2-ncl2)1/2=(u2+w2)1/2, u and w are derived by solving the
Maxwell equation, β is the LP01 mode propagation constant at λ=1550 nm Kl are the Hankel
functions of lth order. The loss coefficient L is shown in Figure 17.
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(a)
(b)
Figure 17 Theoretical calculation of bending loss of SMF-28 fibre at λ=1500nm a) Linear scale, b)
Logarithmic scale.
Detailed bending loss measurements are shown in Figure 18 and Figure 19 for both acrylate and
polyimide coated SMF-28 fibres together with the modelling results. From the measured data on
Figure 18 and Figure 19, one can see the coherent coupling (oscillations) between the
fundamental propagation field and the reflected radiated field by the coating layer, i.e., so called
whispering-gallery mode, has an apparent effect on bend loss characteristics so that the
calculated results with the simplest model, i.e., treating the fibres as the core and infinite
cladding structure, are obviously different from the measured bend losses, although they are
generally in good agreement. A more elaborate modelling scheme is required taking into
account the multilayered high refractive index polymer coating structure.
measured bending loss results were very reversible.
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(b)
Figure 18 Bending loss measurements of acrylate coated SMF-28 fibres (blue line, circles: experimental
data. Red line: modeling) a) Linear scale, b) Logarithmic scale
(a)
(b)
Figure 19 Bending loss measurements of polyimide coated SMF-28 fibres (blue line, diamonds:
experimental data. Red line: modeling) a) Linear scale, b) Logarithmic scale
Table II
Fibre Bending Loss Summary
Measurement Dynamic Range
~1000
Displacement Resolution (mm)
0.2-0.7
3.3.2.2
PHOTONIC CRYSTAL FIBRES (PCFS)
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More recently, a fibre with a fundamentally new design has been demonstrated: the photonic
crystal fibre (PCF). This is made from a single material such as (undoped) fused silica. The
fibre incorporates a periodic array of air holes lying along the fibre, an example of a 2-D
photonic crystal. A missing hole leaves an extended solid region - a high-index "defect" - that
acts as the fibre's core. The surrounding material acts as the cladding ( Figure 20). This core is
index guiding (by total internal reflection) because the cladding with its holes has a lower
effective refractive index than the core. PCFs with a low-index defect have also made: an extra
or enlarged hole. These can only guide light by photonic band gap effects - PBG guiding.
Figure 20 SEM Photograph of an endlessly single mode PCF (ESM-12-01, Crystal Fibre)
PCFs have a number of remarkable properties:

They can be single mode at all wavelengths, unlike conventional fibres that become
multimode at sufficiently short wavelengths.

They can also be single mode at all scales, making large mode area single-mode fibres
possible without having to control the concentration and distribution of dopants. This makes
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them interesting for optical power delivery, an application that further benefits from the
absence of dopants.

They can confine light to a hollow single-mode core: the ultimate high-power fibre!

Their (normal) chromatic dispersion can be tailored by simply changing the hole size and the
pitch, making them suitable for dispersion compensation and optical nonlinear applications.

Multi-core PCFs can be made as easily as single core PCFs simply by stacking additional
solid rods.

The ability to interact optically with gases in the holes, and to guide modes with effective
indices to phase-match with water and even vacuum, makes PCFs significant as sensor
elements.

They have unusual bending loss properties at the low wavelength range.
3.3.2.3
BENDING LOSS OF ENDLESSLY SINGLE MODE (ESM) PCF
We tested, an ESM-PCF (ESM-12-01, Crystal Fibre) with a core diameter of about 12 μm in
which the air holes of diameter 3.5 μm are arranged in a hexagonal pattern with a pitch of
7.7 μm, as shown in Fig. 1. The measured output power for various bend radiuses, r,
applied in a circular loop with a single-turn of fiber at 473, 633 and 1550 nm wavelengths
with arbitrary angular orientation, as shown in Fig. 5. Standard ESM-PCF was remained
unaffected to bend perturbations for small enough bend diameters at 1550 nm. However,
high bend loss levels were reached for below 5 mm radius. As the wavelength decreases to
633 and 473 nm, the fiber obtained same loss for bend radius as large as r = 50 mm. Shortwavelength bend loss edge of PCFs
has been described as
from a simple
effective index method. The high bending loss at large bending radius measured at 473 nm
can potentially sense large displacements in bulky structures. It must be also noted here
that, partially peaks of the bend losses presented in our experimental results originated due
to coupling between the fundamental core mode and the gallery of cladding modes.
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Figure 21 Bending loss measurements of PCF (ESM-12-01, Crystal Fibre) at different
wavelengths (green line, diamonds: 1550nm. Red line squares: 633nm. Blue line,
circles: 473 nm). For comparison (black line, triangles) bending loss measurements
of acrylate coated SMF-28 fibre.
3.3.2.4
BENDING LOSS OF HETERO-CORE FIBRES
Figure 22 shows an alternative macro-bending loss sensor based on spliced fibres with different
cores. It consists of a transmission portion and a spliced hetero-core portion in the centre. This
structure uses optical fibres with different core diameters. The transmission portion has a 9-μm
core diameter single-mode fibre (SMF) into which a hetero-core consisting of a 5-μm SMF with
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a length from 5 mm to 1 cm is inserted. When light in the transmission portion passes into the
hetero-core portion, a small amount of the light is lost by leakage. We believed that this loss
would increase with gentle bending. By a simple conversion mechanism, in which linear
displacement is converted into moderate bending, it has been found that the displacement is
sensitively reflected by a change in loss with sufficient reproducibility even at very large bending
radiuses.
Figure 22 Hetero-core fibre bending loss sensor scheme
3.3.3
3.3.3.1
TEST
TEST OF EMBEDED BENDING LOSS SENSORS
Figure 23 shows the proposed set-up for the bending loss measurement. It is a simple light
power measurement. Light is launched from a diode laser to a single loop pre-bent optical fiber
embedded in the specimen. The output is monitored with a hand held fiber coupled power
meter. Any change in power is directly linked to a change in the bend diameter of the fiber and
hence to strain for a specific specimen. Also, directionality of the pre-bent fiber can be directly
linked to strain directionality. A standard acrylate, polyimide coated SMF-28 fiber or PCFwith
external diameter (with the coating) of 250 μm can be used.
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Figure 23 Schematic diagram of a single loop bending sensor embedded inside a
dogbone for power transmission measurements.
Figure 24 Schematic diagram of a half loop bending sensor embedded inside a
dogbone for power transmission measurements.
Figure 24 shows an alternative set–up scheme where the light is launched to a half loop pre-bent
optical perpendicular to the direction of strain. The measurement principle is exactly the same
with Figure 23 with an increase in transmission power in this case since the bending radius
increases.
The bending loss measurement set-up can be further simplified and at the same time improved
with the use of an introduction of a Optical Time Domain Reflectometer (OTDR), as shown in
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Figure 25. An OTDR injects a series of optical pulses into the fibre under test. It also extracts,
from the same end of the fibre, light that is scattered back and reflected back from points in the
fibre where the index of refraction changes. (This is equivalent to the way that an electronic
TDR measures reflections caused by changes in the impedance of the cable under test.) The
intensity of the return pulses is measured and integrated as a function of time, and is plotted as
a function of fibre length. An OTDR may be used for estimating the fibre’s length and overall
attenuation, including splice and mated-connector losses. It may also be used to locate faults,
such as breaks or macro-bends in the optical fibre. The main advantage of OTDR over the
scheme in Figure 23 and Figure 24 is that a single fibre end is used for both input and output. It
can also measure and resolve multiple bending losses along the same fibre length.
In order to have a proof of principle of the single loop bending sensor, we embedded a looped
fibre inside a PDMS elastomeric block as shown in Figure 26. PDMS (poly-dimethylsiloxane)
elastomer is a polymeric silicone material widely used in the area of photonics, particularly in
opto/microfluidics, having very optical properties. It also exhibits very good mechanical
properties due to low Young's modulus; it is soft and deformable with no shrinkage and
combined with its low cost, and ease fabrication procedure is a potential active material for
tunable devices and sensing applications.
Figure 25 Schematic diagram of a single loop bending sensor embedded inside a
dogbone for Optical Time Domain Reflectometry (OTDR) measurements.
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Figure 26 Photographs of a single loop bent fibre embedded inside a PDMS block.
Red light is launched in order to high light the loop.
The PDMS block dimensions Length x Width x Height are 162x30x11 mm and the single loop
diameter is 18mm. Figure 27 shows detailed transmission measurements of ~1.2mm
displacement along the length of the block.
(a)
(b)
Figure 27 Measurements of optical transmission vs. linear displacement of a single
loop bent fibre embedded inside a PDMS block. a) Absolute power measurements, b)
Normalized power measurements. Red line: Linear fitting.
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A linear displacement along the length of the block of ~1.2mm (~0.8% strain) induced more
bending of the fibre and as a result a drop in the transmitted power of ~33%, proving that the
that the embedded single loop bent fiber can be used as a strain sensor.
3.4
POLYMER OPTICAL FIBER (CONTRIBUTION OF NHRF)
3.4.1
POLYMER OPTICAL FIBER SENSORS FOR STRUCTURAL HEALTH MONITORING
Monitoring strain in SRMs will need probably to be divided in two distinct parts and areas of application,
the first for measuring strain in the actual composite tube hosting the propellant material and second for
measuring the strain conditions in the actual energetic propellant material. The strains anticipated in the
first case in this stiff filament wound composite tube are much smaller than the expected strain in the
propellant and thus different and suitable solutions and architectures should be proposed and
implemented for successful and effective monitoring in both parts of the SRM. Monitoring the composite
tube has been partially investigated in the literature by incorporating suitable FBG architectures or
suitably embedded piezoelectric elements –PZT creating a sensors’ network along the tube.
Monitoring of strain condition inside the actual energetic material of SRMs requires solutions and
architectures able to monitor quite high values of strain. While Fiber Bragg Gratings is a very well known
and widely applied solution for localized SHM the required strain range above 5-10% makes necessary
the identification of alternative solutions able to monitor such large strains.
As such a solution is proposed the use of Polymer Optical Fibers –POF. Polymer Optical Fiber sensors
are currently attracting a quite intense interest in Structural Health Monitoring applications, being low
cost solutions and also robust and highly reliable. Among the different types of optical fibers (such as
traditional silica fibers) POF offer the additional advantages of low cost, easy handling, high elastic
strain limits, high fracture toughness, and high sensitivity to the amount of strain when they are stressed
or pulled. Moreover the experimental methodology that can employed with POFs is much easier and
cheaper, enabling also the autonomous operations and wireless networking of the sensors when this is
required. While in SESAMO project FBGs are suitable for monitoring the composite tube, POF solutions
could be used for both the monitoring of the composite and monitoring also of the propellant material
behavior.
POFs are typical optical fibers made of optically transparent polymeric materials. The cross section is
typically circular with three distinct layers, core, cladding and protective jacket. The jacket could be
easily customized and could be, polyethylene, polyvinylchloride, chlorinated polyethylene depending on
the actual end application.
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Figure 28
There is also relatively large variety and options of polymeric materials and also geometrical fibers’
characteristics. Typical and widely used material for POFs is PMMA (Polymethyl-Methacrylate Resin)
with a typical refractive index 1.49. Typically the cladding is Fluorinated Polymer of lower refractive
index. The standard 1mm diameter POF has a 980 micrometer core and 20 micrometer cladding (for
example ESKA CK40, Mitsubishi Rayon). The graph below shows the attenuation diagram for the typical
polymeric material, together with the used wavelengths.
Figure 29
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Depending on the required operating characteristics, different POFs could be used. For example for
relatively elevated operating temperatures, thermal stability, and for low loss (increased used POF
length) applications the preferred solution is perfluorinated polymer – CYTOP.
However it should be stressed and should be taken also into account in our future studies that since the
actual physical and mechanical characteristics of POFs are strongly dependent on the actual
manufacturing and drawing method of each industrial manufacturer the used POF should be identical or
should be recalibrated if are from different suppliers.
Until recently the used POFs for SHM were almost exclusively highly multimoded with a typical 1 mm
diameter as described above. Comparative studies have demonstrated that POF offer good stability and
repeatability and the measuring performance is very similar to traditional Bragg gratings at a reduced
cost and complexity. Moreover another feature useful for certain SHM applications is the fact that the
POF provide stress information for all the area the POFs are located where FBG are suitable for rather
localized strain measurements. Since PMMA elastic limit is 10% in contrary to silica which is 1-2%,
POFs provide largest strain measurement range. Through mechanical testing at various strain rates, it
was determined that the average POF failure nominal strain was between 30% and 40%. The yield
strain, however, was between 5.3% and 8.0%, depending upon the applied strain rate. In contrast, the
brittle failure limit for silica optical fibers is approximately 3%–4%. Multimode POF have been
successfully used for SHM in constructions and buildings, as seen below in the detection of a crack in
concrete. Is obvious the large deformation of the POF still below its crack point and able to monitor the
concrete crack.
Figure 30
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The strain measurement is almost linear to the attenuation of transmitted light providing a direct
interrogation method. The figure below illustrates different POFs embedded in a composite material for
monitoring the stress behavior.
Figure 31
The embedded POFs could monitor the behavior of the composite material during its curing process
also. The graph bellow shows the strain gauge readings relatively to the normalized attenuation,
exhibiting a linear relation.
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Figure 32
It is well known that embedding fibers in a composite material or even worse in a laminate could disturb
the geometry or even its structural integrity. So is required to select appropriate POF dimensions for
inducing minimal or even compatible to the existing geometry disturbances. In similar problems specific
suitable POF with appropriate diameters have been selected from a variety of dimensions such as 1mm,
0.5mm. 0.25 mm providing thus successful solution in monitoring glass-fiber epoxy laminates. Single
mode PMMA POFs have for comparison a diameter of 115µm
If it is required, the sensitivity of the POF for measuring strain locally could be enhanced by careful
partial removal of the cladding (by polishing the POF) making thus the fiber more sensitive to bend loss
or to elongation by allowing more trapped light to escape through the cladding.
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Figure 33
Alternatively those points of increased sensitivity can be produced in a highly controllable way, by laser
micromachining of the POF as we have already demonstrated, and as can be seen in next figure where
a standard 2mm diameter POF was etched with 193 ArF Excimer pulsed Laser radiation.
Figure 34
This way we could ad-hoc induce along the POF certain points of increased sensitivity in areas for
example in the propellant where there is an increased risk of excessive stresses or cracks. This way we
could create a quasi distributed sensing system allowing the average monitoring of strain along the fiber
together also with the examination of discrete points.
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The typical interrogation system for a multimoded POF is based on power transmission measurements,
as described above. If we furthermore need an increased spatial resolution for distributed sensing along
the POF we could use special high sensitivity Optical Time Domain Reflectometry in order to monitor the
backscattered light of each perturbation due to stresses. Such a method to monitor very low
backscattered light is the Photon Counting OTDR and could provide a spatial resolution in the order of
10 cm.
NHRF has already developed an autonomous and wireless enabled POF based sensor system
demonstrated so far for bend loss measurements as well as for liquid level and distributed flood
monitoring. The transmitter (a LED of about 200 microwatts at 650 nm) and the receiver are embedded
in the same Fiber Optic Driver Circuit Board (FODCB). The receiver is connected to a phototransistor
which is serially connected to an amplifier for amplifying the phototransistor’s signal. The FODCB
together will all the appropriate electronic driving circuits can be seen in the figure below. The output of
the amplifier is connected to the A/D converter of the wireless sensor node for further processing. The
FODCB is supplied by 2 AA batteries of 3V in total.
Figure 35
Further to the typical
Multimoded POFs lately there is an increasing amount of research on single-
mode POFs that have been demonstrated and used in some specific SHM applications. Single-mode or
few-mode POFs have much smaller dimensions (even smaller than silica fibers) and can allow the
manipulation of their optical propagating properties (optical modes) in a way similar to silica fibers. So is
possible the inscription of Bragg gratings serving thus the same applications as with silica fibers, but
with having the additional advantage of wider tuning and strain range measurement as the elastic limit of
POF is much larger than that of silica. Experimentally has been reported a 10 times higher tuning range
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(~75nm) in Bragg grating POF compared to the tuning range of a typical silica fiber BG. Similarly Long
Period Gratings –LPG have been demonstrated in single mode POFs serving as strain sensors.
The first PMMA single-mode POF demonstrated at an operating wavelength of λ = 1300 nm. These
fibers demonstrated attenuation levels of approximately 0.3 dB/cm, close to the intrinsic material
attenuation for PMMA at that wavelength. Later fabricated single-mode PMMA optical fibers with
attenuation levels of 0.25 dB/cm at λ=1550 nm and 0.05 dB/ cm at λ = 850 nm, and finally improved
single-mode POFs at λ = 1060 nm with an attenuation level of 0.18 dB/ cm. Currently, doped, singlemode POFs are still experimental, with only very few commercial manufacturers. An additional solution
to decrease the attenuation of single-mode POFs is the fabrication of microstructured POFs (mPOFs)
which have air holes in the fiber cross-section.
Figure 36
The air fraction decreases the material in a given cross-section and therefore decreases the intrinsic
losses. Additionally, lightwave guidance in the mPOF can be controlled by the hole arrangement and
can be considerably stronger than that in solid core, single-mode optical fibers. These two features
decrease the attenuation in mPOFs significantly. Such fibers can also function as endlessly singlemode optical fibers, meaning that they are single-moded for a wide range of wavelengths. This singlemode, low attenuation operation could enable POF sensors in the near-infrared wavelength range where
many telecoms components are commercially available. Those mPOFs could be employed very
effectively for Bragg grating and long period grating inscription. Preliminary studies using mPOF sensors
based on LPGs demonstrate that the use of polymer fiber increases the range of repeatable strain
measurements by several times and the yield limit by an order of magnitude, compared to a silica-based
sensor. The following graph demonstrates the strain results in an mPOF where the strain removed
rapidly after the application
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Figure 37
The viscoelastic properties of the polymer means there are time dependent effects relating to strain rate
and magnitude. These effects are small when the sensor is intermittently strained up to 2%, and are
relatively small at strains of up to 4%–5%. Further testing is ongoing to characterize these effects at very
high strains. The effect of stress relaxation has a small effect on the change in the wavelength of the
loss features used in the measurement of strain. The use of high strains complicates response of the
sensor due to the viscoelastic properties, requiring careful calibration, this technology is the only one
that feasibly allows the development of fiber strain sensors that can operate at strains of up to 30%–
45%. These are very high strain sensors are currently investigated by different research groups.
Tensile testing of POFs requires fixtures designed for small diameter fibers with low rigidities. Extensive
tensile testing reports of single-mode PMMA POFs under various strain rates and cyclic load conditions
are available for the literature. A summary of measured properties, including Young’s modulus (E), yield
strain (εY) and ultimate strain (εU), are listed in the following table
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The following figure presents typical true stress–strain curves for a single-mode POF measured in a
universal testing machine. The applied strain rate was varied from 0.01 to 3.05 min−1. The failure strain
for the POF was around 30% for most of the samples. The yield strain increased with the applied strain
rate, while the initial slope of the curve also increased, although by a much lower amount.
Figure 38
Additional recent studies have allowed measurements of large deformations by measuring phase shift
changes changes in SM POF. A recent study by Kiesel examined whether the maximum usable strain
for a POF sensor is determined by the yield strain of the POF, as in the case of POF Bragg gratings. It
was demonstrated that the maximum usable strain range in a pristine POF is in fact much larger than
the yield point. A simple experimental setup for stress loading the POF between two glued points is
shown below.
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Figure 39
Experiments achieved to successfully measure the phase shift in a POF Mach–Zehnder in-fiber
interferometer up to 15.8% nominal strain. This result indicates that interferometric-based POF sensors
can be used for strain ranges over twice that of POF Bragg grating sensors and three times that of silica
in-fiber interferometers. The next graph shows the linear relation of phase shift as a function of applied
nominal strain, demonstrating the possibility to measure such low strains.
Figure 40
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REFERENCES
1. X. Chang, X. He, J. Hu, J. Li, “Experimental Research on Embedded Fiber Bragg Grating Sensors Network for
Solid Rocket Motors Health Monitor”, IEEE First International Conference on Intelligent Networks and Intelligent
Systems
2. X. P. Qing , S. J. Beard, A. Kumar, H.-L Chan, R. Ikegami, “Advances in the development of built-in diagnostic
system for filament wound composite structures”, Composites Science and Technology 66, pp. 1694–1702,
2006
3. K.S.C. Kuang, W.J. Cantwell, P.J. Scully, “An evaluation of a novel plastic optical fibre sensor for axial strain
and bend measurements”, Meas. Sci. Technol. 13 (2002) 1523-1534.
4. X. Chang, M. Li, X. Han, “Recent developments and applications of polymer optical fiber sensors for strain
measurement”, Front. Optoelectron. China, 2(4) 362-367, 2009.
5. Y-C. Chen, L-W. Chen, P-C. Chen, “Combined effects of bending and elongation on polymer optical fiber
losses” Opt. Lett. 30, 3, 230-232, 2005.
6. K.S.C. Kuang, S.T. Quek, C. G. Koh, W.J. Cantwell, P.J. Scully, “Plastic optical fibre sensors for structural health
monitoring: A review of recent progress”, J. Sensors, Article ID 312053, 2009.
7. J. Gomez, J. Zubia, G. Aranguren, J. Arrue, H. Poisel, I. Saez, “Comparing polymer optical fiber, fiber Bragg
grating, and traditional strain gauge for aircraft structural health monitoring”, Appl. Optics, 48, 8, pp. 14361443, 2009.
8. C. Riziotis, D. Dimas, S. Katsikas, A.C. Boucouvalas, “Photonic sensors for autonomous wireless sensing
nodes”, In Proceedings of the 23rd International Congress on Condition Monitoring and Diagnostic
Engineering Management (Nara, Japan, June 28-July 2, 2010). Sunrise Publishing Limited, Hikone, Shiga,
ISBN:978-4-88325-419-4, pp. 669-676.
9. D. Dimas, S. Katsikas, A.C. Boucouvalas, and C. Riziotis, “Wireless-enabled photonic sensor for liquid level
and distributed flood monitoring” 24th International Congress on Condition Monitoring and Diagnostic
Engineering Management, Stavanger, Norway, 30 May-1 June, 2011.
10. K. Peters, “Polymer optical fiber sensors-a review”, Smart. Mater. Struct. 20, Article ID 013002, 2011
11. S. Kiesel, K. Peters, T. Hassan, M. Kowalsky, “Large deformation in-fiber polymer optical fiber sensors”, IEEE
Photon. Technol. Lett. 20, 6, pp. 416-418, 2008.
12. S. Kiesel, K. Peters, T. Hassan, M. Kowalsky, “Behaviour of intrinsic polymer optical fibre sensor for largestrain applications” Meas. Sci. Technol. 18, pp. 3144-3154, 2007.
13. X.-L, Chang, M. Li, X.-F. Gu, “Research on strain transfer of embedded polymer optical fiber sensors based on
linear viscoelasticity”, Journal of Solid Rocket Technology (Guti Huojian Jishu), Volume 33, Issue 3, June 2010,
Pages 353-359
14. K.S.C. Kuang, Private Communication and unpublished material
15. M.C.J. Large, D. Blacket, C.-A Bunge, “Microstuctured polymer optical fibers compared to conventional POF:
Novel properties and Applications”, IEEE Sensors Journal, 10, 7, pp.1213-1217, 2010
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4. BASIC INVESTIGATIONS ON OPTICAL FIBRE APPLICATION IN SOLID ROCKET
MOTORS FOR STRAIN MEASUREMENTS
4.1
INTRODUCTION
The scope of the SESAMO activities covers the application of fibre optics sensors for the monitoring of
structures. One envisaged field of application is the use of optical sensors in solid rocket motors. The
propellant grain, as a key component critical to safety and reliability, is intended to be monitored
throughout its service life in order to assess the impact of thermal and mechanical loads on its structural
integrity.
Both thermal and mechanical loads induce stress and strain fields, which need to be monitored to
assess the risk of failure of the grain due to mechanical fatigue. Failure could also occur because of
exposure beyond the design limits of the grain or because of manufacturing defects. From the
phenomenological point of view, the hereby created failure scenarios are categorised in

Debond of the grain from the insulation or case (both axisymmetric and local)

Longitudinal crack in the bore (radial-axial)

Circumferential crack in the bore (radial-hoop).
All these scenarios may adversely affect the missile from performing its mission or may even lead to a
catastrophic failure at the instance of ignition or during operation of the motor. Typically, the thermal
loads, especially at low temperatures, are more severe than the mechanical loads with respect to
propellant debond or cracking.
Prior to establishing a monitoring concept with optical sensors, it is essential to have an idea of what
secanario creates what occurence that can be captured by the sensors. In today’s view, the fibres can
only be applied at the surface of the propellant grain. In the case of a cylindrical, case bonded
configuration, this is the bore surface, i.e. the central conduit. Therefore, this report is focused on the
induced strains at that location. The investigation is carried out with the Finite Element method,
exemplified for a realistic motor grain design for tactical applications and for a typical cooldown scenario.
Furthermore, an analysis method has been established which basically allows to investigate the induced
fibre strains in helical and other application patterns. A summary of some selected configurations
exemplifies the model.
4.1.1
REFERENCES
[Ref. 1] SESAMO Technical and Commercial Proposal
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[Ref. 2] Requirements and future improvements report
[Ref. 3] Stephen Wolfram, The Mathematica Book, 4th ed.,
Wolfram Media/Cambridge University Press, 1999
[Ref. 4] S. Kiesel, K. Peters, T. Hassan, M. Kowalsky, Behaviour of intrinsic polymer optical fibre sensor
for large-strain applications, Measurement Science and Technology, North Carolina State
University, 2007
4.2
ANALYSIS MODEL FOR FAILURE MODE ANALYSIS
The analyses have been carried out using a Finite Element model of a realistic SRM, applying a typical
cool-down profile (cool-down from curing temperature to -54 °C in 1 day by application of an external
convective boundary condition). The analysis code is MSC/Marc, a tool for performing coupled thermomechanical analysis. For the propellant, a visco-elastic constitutive material model with Herrmann
element formulation has been applied, while the other components are of type linear elastic with
temperature-dependent properties. For the axisymmetric debond analysis, an axisymmetric model (2-D)
is used, whereas for the local debond, axial crack and circumferential crack analysis, a 3-D model
(axisymmetric expanded to 3-D) has to be applied. The element types are QUAD4 and some TRIA3 in
the axisymmetric case, the 3-D-model is HEXA8 dominated with some PENTA5 elements where
needed. By symmetry reasons, only a half model has to be accounted for. The applied symmetry
conditions at the x-y plane make the model be fully equivalent to a 3-D structure.
Some components at the nozzle end side of the motor are touching each other. Therefore, contact
conditions between the relevant parts are established. Figure 41 gives an overview of the model.
All defects (debond, cracks) have been implemented in the centre portion of the grain, where the
induced strains are highest.
The CPU time for one thermo-mechanical run amounts to 12 minutes for the axisymmetric model and
approximately 17 hours for the 3-D case..
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Nozzle
Propellant grain
Bore
Insulation
Steel case
Figure 41 – Finite Element analysis model of example SRM
4.2.1
AXISYMMETRIC DEBOND
Debond is a local detachment of adjacent structural components. Figure 42 shows the implemented
debond between propellant and insulation for the axisymmetric case. Two debond lengths l were
chosen: 17.9 mm and 35.8 mm.
Insulation
Propellant
l
Figure 42 – Axisymmetric debond
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LOCAL DEBOND
For local debond, the 3-D model is used (see Figure 43). The debonded area has the length l in axial
and the width  in hoop direction (/2 in the symmetric model). Two different local debond defects are
considered: l=35.8 mm with =8.35 mm and 16.7 mm.
Propellant
2
l
Figure 43 – Local debond
4.2.3
RADIAL-AXIAL CRACK
Crack here is defined as a local detachment within one material domain. For classification reasons, a
crack is orientd radially and extends in either axial or in hoop direction. Figure 44 shows the dimensions
of the radial-axial crack.
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l
Propellant
r
Figure 44 – Crack in radial and axial orientation
4.2.4
RADIAL-HOOP CRACK
The radial-hoop crack is shown in Figure 45. Its detached surface is normal to the motor axis. It has the
radial depth r and extends over 180 degrees in hoop direction (90 degrees in the symmetric model). In
the post-processing, results for various angular positions along the motor axis are given. The positions
are shown in Figure 45.
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Propellant
Position 3
r
Position 2
90 °
Position 1
Figure 45 – Crack in radial and hoop orientation; Location of output positions
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RESULTS OF FAILURE MODE ANALYSIS
The resuts of the FEM failure mode analyses are given in terms of induced axial and hoop strains vs.
motor axis for the soaked temperature condition of -54 °C at the end of the cooldown process (t=1 day).
All strains are given in terms of natural (Cauchy) strains. A value of 0.05 for instance means 5 % strain.
A cylindrical coordinate system is introduced with r being the radial,  the hoop and z the axial direction.
Figure 46 shows the path (along bore surface in axial direction of the motor) of the presented results.
Path of results plots (z)
Figure 46 – Path of results plots
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CONFIGURATION WITHOUT DEFECTS
Figure 47 – Hoop and axial strain in bore for original configuration without defects
Except in a small area of the nozzle regime, the axial strain is negative. It is governed by the axial
contraction of the steel case at cooldown. In the present case, the length-to-diamter ratio (l/d) of the
motor approaches 6. The lower the l/d ratio, the higher (in negative direction) the axial strain would be
due to the increasing influence of boundary effects. The hoop strain is positive throughout the cylindrical
bore and goes up to approx. 10% in the present case. It is governed by the different coefficients of
thermal expansion (CTE) of the steel case on the one hand and the insulation layer and the propellant
grain on the other. Typically, the CTE of the polymer-based insulation and propellant components
exceed the CTE of steel by a factor of 5 to 10.
At the end of this paragraph a concise summary in table-form is given for the local debond and cracks.
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AXISYMMETRIC DEBOND
The effect of axisymmetric debond on the induced strains in the bore are illustrated in Figure 48 and
Figure 49. Two lengths of delamination at the centre portion of the motor were chosen: 17.9 mm and
35.8 mm.
Debond
Figure 48 – Fringe plot of hoop strain for axisymmetric debond (l=17.9 mm)
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Figure 49 – Effect of axisymmetric debond
Significant effects are observed in both hoop and axial strain. The hoop strain at the centre of the
debonded area drops from 10.1 % to 8.9 % and 6.5 % respectively. A remarkable result is found for the
axial strain. From an originally slight negative value (-0.18 %) it reverts to positive ones (+0.92 % and
+3.0 %). This is a coupled effect: the missing bond to the insulation makes the propellant in the vicinity
of the defect radially lag behind the rest of the grain, thus resulting in positive bending strain in the bore.
4.3.3
LOCAL DEBOND
Local debond is analysed using the 3-D model (see Figure 50).
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L=35.8 MM, =8.35 MM
Debond
Figure 50 – Fringe plot of hoop strain for local debond (l=35.8 mm, =8.35 mm)
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Figure 51 – Effect on axial strain for local debond (l=35.8 mm, =8.35 mm)
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Figure 52 - Effect on hoop strain for local debond (l=35.8 mm, =8.35 mm)
Figure 53 - Effect on axial and hoop strain for local debond (l=35.8 mm, =8.35 mm) vs. perimeter
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L=35.8 MM, =16.7 MM
Debond
Figure 54 - Fringe plot of hoop strain for local debond (l=35.8 mm, =16.7 mm)
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Figure 55 - Effect on axial strain for local debond (l=35.8 mm, =16.7 mm)
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Figure 56 - Effect on hoop strain for local debond (l=35.8 mm, =16.7 mm)
Figure 57 - Effect on axial and hoop strain for local debond (l=35.8 mm, =16.7 mm) vs. perimeter
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RADIAL-AXIAL CRACK
For the radial-axial crack scenario, four different crack sizes are under consderation:
l=35.9 mm and 107.9 mm with r=3.9 mm and 9.06 mm.
4.3.4.1
L=35.9 MM, R=3.9 MM
Figure 58 - Effect on axial strain for radial-axial crack (l=35.9 mm, r=3.9 mm)
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Figure 59 - Effect on hoop strain for radial-axial crack (l=35.9 mm, r=3.9 mm)
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Figure 60 - Effect on axial strain for radial-axial crack (l=35.9 mm, r=3.9 mm), detailed view
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Figure 61 - Effect on hoop strain for radial-axial crack (l=35.9 mm, r=3.9 mm), detailed view
Figure 62 - Percentual strain deviation in bore vs. axis for radial-axial crack (l=35.9 mm, r=3.9 mm) at lower
and upper angular positions
4.3.4.2
L=107.9 MM, R=3.9 MM
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Figure 63 - Effect on axial strain for radial-axial crack (l=107.9 mm, r=3.9 mm)
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Figure 64 - Effect on axial strain for radial-axial crack (l=107.9 mm, r=3.9 mm), detailed view
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Figure 66 - Effect on hoop strain for radial-axial crack (l=107.9 mm, r=3.9 mm), detailed view
Figure 67 - Percentual strain deviation in bore vs. axis for radial-axial crack (l=107.9 mm, r=3.9 mm) at
various angular positions
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L=35.9 MM, R=9.06 MM
Figure 68 - Effect on axial strain for radial-axial crack (l=35.9 mm, r=9.06 mm)
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Figure 69 - Effect on axial strain for radial-axial crack (l=35.9 mm, r=9.06 mm), detailed view
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Figure 70 - Effect on hoop strain for radial-axial crack (l=35.9 mm, r=9.06 mm)
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Figure 71 - Effect on hoop strain for radial-axial crack (l=35.9 mm, r=9.06 mm), detailed view
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L=107.9 MM, R=9.06 MM
Figure 72 - Effect on axial strain for radial-axial crack (l=107.9 mm, r=9.06 mm)
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Figure 73 - Effect on axial strain for radial-axial crack (l=107.9 mm, r=9.06 mm), detailed view
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Figure 75 - Effect on hoop strain for radial-axial crack (l=107.9 mm, r=9.06 mm), detailed view
Figure 76 - Percentual strain deviation in bore vs. axis for radial-axial crack (l=107.9 mm, r=9.06 mm) at
various angular positions
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RADIAL-HOOP CRACK
Three different crack sizes are accounted for in the radial-hoop case: r=3.9 mm, 14.2 mm and 29.7
mm. All these cracks are assumed to extend by 180 degrees in hoop direction.
4.3.5.1
R = 3.9 MM
Figure 77 - Effect on axial strain for radial-hoop crack (r=3.9 mm)
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Figure 78 - Effect on axial strain for radial-hoop crack (r=3.9 mm), detailed view
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Figure 79 - Effect on hoop strain for radial-hoop crack (r=3.9 mm)
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Figure 80 - Effect on hoop strain for radial-hoop crack (r=3.9 mm), detailed view
Figure 81 - Effect on axial and hoop strain for radial-hoop crack (r=3.9 mm) vs. perimeter
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4.3.5.2
R = 14.2 MM
Figure 82 - Effect on axial strain for radial-hoop crack (r=14.2 mm)
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Figure 83 - Effect on axial strain for radial-hoop crack (r=14.2 mm), detailed view
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Figure 85 - Effect on hoop strain for radial-hoop crack (r=14.2 mm), detailed view
Figure 86 - Effect on axial and hoop strain for radial-hoop crack (r=14.2mm) vs. perimeter
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R = 29.7 MM
Figure 87 illustrates the appearance of a radial-hoop crack with r=29.7 mm.
Radial-hoop
crack
Figure 87 – Fringe plot of axial strain for radial-hoop crack (r=29.7 mm)
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Figure 88 - Effect on axial strain for radial-hoop crack (r=29.7 mm)
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Figure 91 - Effect on hoop strain for radial-hoop crack (r=29.7 mm), detailed view
Figure 92 - Effect on axial and hoop strain for radial-hoop crack (r=29.7 mm) vs. perimeter
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SUMMARY OF RESULTS
Table 1 gives a concise summary of the failure mode analysis. The indicated values refer to the
extremal points of strain (hoop and axial) at the centre of the bore. The maximum change of induced
strains  is given in absolute values, the adjacent column indicates the percentual change relative to
the faultless motor.
Location of
strai
Defect
n
Max. change in
Max. change in
hoop strain
axial strain
 [%]
[%]
 [%]
[%]
outp
ut
Axisymm.
l=17.9 mm
perimeter
-1.19
-11.8
1.11
-602
Figure 49
l=35.8 mm
perimeter
-3.60
-35.6
3.15
-1710
Figure 51,
l=35.8 mm,
Upper side
6.60E-3
0.0653
0.171
-92.7
Lower side
-0.0762
-0.754
4.39E-3
-2.38
l=35.8 mm,
Upper side
0.107
1.06
0.473
-257
=1
Lower side
-0.0944
-0.934
0.0123
-6.60
l=35.9 mm,
Pos. 1
-0.101
-0.997
5.95E-3
-3.23
to
r=3.
Pos. 2
-0.122
-1.204
0.0156
-8.47
Figure 62
9 mm
Pos. 3
-8.55
-84.6
0.240
-130
l=107.9 mm,
Pos. 1
-0.0672
-0.665
6.42E-3
-3.48
to
r=3.
Pos. 2
-0.143
-1.42
0.0279
-15.1
Figure 67
9 mm
Pos. 3
-8.16
-80.8
0.141
-76.4
Pos. 1
-0.0708
-0.701
9.92E-3
-5.38
Pos. 2
-0.166
-1.65
0.0526
-28.5
de
bo
nd
Figur
=8.
e 52,
Local
Figur
de
bo
e 53
Figure 55,
Figur
nd
35
mm
e 56,
Figur
e 57
Figure 58
Radialaxi
al
cra
Figure 63
ck
Figure 68
to
6.7
mm
l=35.9 mm,
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Pos. 3
-0.148
-146
1.00
-542
l=107.9 mm,
Pos. 1
-0.0792
-0.784
0.0292
-15.8
r=9.
Pos. 2
-0.427
-4.23
0.128
-69.4
06
Pos. 3
-0.141
-140
0.603
-327
Pos. 1
-0.0652
-0.645
2.02E-3
-1.10
Pos. 2
-0.365
-3.61
-2.10
1140
Pos. 1
-0.0835
-0.826
0.0204
-11.1
Pos. 2
-2.94
-29.1
1.73
-936
Pos. 1
-0.131
-1.29
0.100
-54.2
Pos. 2
-5.42
-53.6
4.80
-2600
06
mm
Figure 72
to
Figure 76
mm
Figure 77 to
Figur
Radial-
r = 3.9 mm
e 81
ho
Figure 82 to
op
Figur
cra
e 86
ck
r = 14.2 mm
Figure 88 to
Figur
r = 29.7 mm
e 92
Table 1 – Summary of effects on peak strains in bore
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INVESTIGATIONS OF FIBRE STRAINS
4.4.1
GENERAL CONSIDERATIONS
The following analytic investigation was performed with Mathematica (Wolfram Research), a fully
integrated environment for scientific and technical computing (Ref. 3).
The strain (like the stress) in any point of a solid body may be expressed in tensor-notation. In general,
when no presumption is made on a particular component, it is of triaxial nature. Expressed in an
orthonormal cylindrical coordinate system with the components r,  and z, the strain then is written as
  rr  r  rz 




 triaxial   r     z  .


  zr  z  zz 


Like the stress tensor, the strain tensor generally is symmetric to its main diagonal, i.e. ij = ji.
Therefore, when referring to an arbitrary coordinate system, six different components in every point of
the structure are needed to fully characterize the state of strain. In the special case, when using the
principal coordinate system, only the three main diagonal terms apply, all other (deviatoric) terms, by
definition, then are zero (however, now three additional pieces of information are needed that define the
orientation of the principal system, such as the three Eulerian angles).
The individual components of the tensor depend on the selected coordinate system. A given strain
 
tensor  , expressed in the system {x,y,z}T, changes its appearence  * when expressed in a different
system {x’, y’, z’}T according to the transformation rule
 *   . . T ,
with  being the transformation matrix1. It can be expressed in terms of the direction cosines of the new
and the original coordinate axes:
 cosx' , x  cosx' , y  cosx' , z  


   cos y' , x  cos y' , y  cos y' , z   .
 cosz ' , x  cosz ' , y  cosz ' , z  


1
The three invariants, which form the base of any failure criterion, remain unchanged under a transformation.
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As mentioned in the introduction, the strain measurements with fibre optics on the solid propellant at
present is believed to be practically confined to the surface of the grain. That means, the state of strain
under consideration is of biaxial nature. In particular, when considering the bore, the radial component r
drops out and the two components  (hoop direction) and z (axial direction) remain.
The strain tensor then reduces to
     z 
,
 biaxial  



z

zz


and the transformation matrix reduces to
 cos ' ,   cos ' , z 
 .
  




cos
z
'
,

cos
z
'
,
z


Analogous to the triaxial state, a biaxial state of strain is fully determined by three independent variables.
In our case these are   ,   z and  zz .
In a cylindrical grain configuration at high l/d ratio however, the principal directions are known to be the
hoop and the axial direction. Thus, the cylindrical system is identified as a principal system, and the
number of unknowns is reduced by 1.
The strain tensor simplifies to
   0 
.
 biaxial   princ   0

zz 

The unknowns now are the (principal) strains   and  zz .
A fibre captures the strain component fibre in a given direction  (fibre longitudinal sense) at a specific
point P of the strain field (Figure 93).
 zz
P

 fibre

Figure 93 – Fibre orientation in a strain field
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The transformation rule for the strain field expressed in principal coordinates yields
2
2
 fibre    cos     zz sin    .
Figure 94 shows the induced strain in a fibre vs. inclination angle In order to get insight into the
sensitivity of potential angular misalignments d, the values for d=±1° and ±3° are plotted in addition.
This topic will be discussed in more detail in Chapter 4.4.4.
Figure 94 – Fibre strain vs. fibre orientation in a principal strain field
In our example of a long cylindrical grain it follows that, at any point, two different, independent
measurements are needed to characterize the state of strain (see Figure 95). When assuming that the
induced strains are axisymmetric, it is sufficient to have the measurements at the same axial position,
but not necessarily at the same point (strains then are independent of hoop position ).
Another result is that the inclination angles must differ by their absolute values, i.e. | 1|  | 2|.
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
r
z
hoop
Fibre 1
1
2
Axial ()
P
Fibre 2
Figure 95 – Fibres in a cylindrical configuration
4.4.2
SPECIFIC CONFIGURATIONS
In these general considerations, a set of assumptions is made on the fibre behaviour.
These are:

The fibre is assumed to perfectly follow the propellant matrix deformation and there is no
interaction of the fibre on the carrier material. In the ideal case, this means
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Stiffness of fibre << stiffness of propellant material.

The fibre is assumed to have no apparent thickness and the strains induced in transverse
direction have no effect on the measured values.

The fibre is assumed to be temperature compensated, i.e. the temperature has no impact on the
measurement.

No specific location of grating along the fibre is assumed. The positions, lengths and numbers of
gratings can be adjusted to the needs of the application.

The propellant is treated as a continuum (strain of matrix and AP particles compound is
considered as uniform average strain).
In reality there will be an interaction to the greater or lesser extent. Currently, the fibres are being tested
with respect to their stiffness and their adhesive capability to the propellant. Once these data are
available, the interaction topic will be investigated in addition by FEM.
The motivation behind the following considerations is to gain insight into possible fibre configurations
and to learn how geometric parameters affect the measurements.
For the mathematical treatment of a particular fibre configuration, a function f() is introduced, with 
being aligned with the z-axis.  may not necessarily start at z=0. This function then is ‘wrapped’ along
the axis to cover the surface of the bore, which, as a cylinder, is a ruled surface.
The local coordinate system {’,z’}T is aligned with the fibre (’ in longitudinal direction, z’ = r × ’).
For the transformation matrix


f
of the fibre strain tensor we findì
f
 f   / 


 1  f  /  2
 

1


 1  f    /  2







2
1  f    / 
.
 f   / 

2
1  f    /  

1




An important aspect in the assembly of fibres is that the bending radius is limited by mechanical
constraints. It is determined by factors as the fibre core radius, the coating thickness and the stress
capability. The bending radius  in a given configuration is linked with the curvature K by  
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For the determination of the curvature, a cartesian coordinate system {x,y,z}T, attached to the motor, is
introduced, where x denotes the horizontal, y the vertical and z the axial component.
Hence, x   r cos   , y   r sin    and z    , with    being the circumferential
position (see Figure 95). The fibre may start at any initial position  0 in the bore of radius r:
     0  f   / r .
With the derivatives
d x 
d y  
d z  
 1
, y' 
, z' 
d
d
d
d 2 x 
d 2 y  
d 2 z  
 0
x' ' 
, y' ' 
, z' ' 
d2
d2
d2
x' 
we formally find the spatial curvature K of the fibre as
K2 
4.4.2.1
x'2  y'2  z'2x' '2  y' '2  z' '2  x' x' '  y' y' '  z' z' '2 .
x'2  y'2  z'23
SINUSOIDAL ARRANGEMENT
In the first examined configuration a sinusoidal shaped fibre arrangement (Figure 96) is looped around
the bore at the central position of the motor (x=500 mm).
Surface of
bore
Fibre
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Figure 96 – Sinusoidal fibre arrangement
Figure 97 shows the induced strains in fibre longitudinal and transverse direction. These are plotted
versus the fibre longitudinal coordinate.
Figure 97 – Induced fibre strains in longitudinal and transverse direction in sinusoidal configuration
The curvature radius vs.hoop angle is given in Figure 98
Figure 98 – Curvature radius in sinusoidal configuration
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It turns out that the sinusoidal configuration offers some major drawbacks:

The induced strain is locally high and has the full hoop strain as a maximum, which may raise a
problem with respect to the maximum strain capablity of the fibre.

The curvature radius is locally low, which may conflict with the minimum bending capability.
With a decreasing number of waves around the perimeter the latter conflict could be mitigated, whereas
the problem with the high strains is immanent to that configuration.
4.4.2.2
HELICAL ARRANGEMENT
A more promising configuration was identified in the helical arrangement. It offers more flexibility in
adjusting the maximum induced strain, has lower bending radii and a single fibre can span over the
entire length of the motor. Both helices with constant and variable pitch have been considered.
4.4.2.2.1 Constant pitch
At constant pitch (either positive or negative) the shape of the fibre is a helicoid.
Figure 99 gives an example with two fibres, starting close to the head end of the grain (x=20 mm).
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Figure 99 – Example of constant pitch helices
The induced fibre strains vs. fibre longitudinal coordinates are given in Figure 100.
Figure 100 - Induced fibre strains in longitudinal and transverse direction in constant pitch configuration
A graphical overview provides Figure 101 with the longitudinal fibre strains displayed in colours
(blue=low strain, red=high strain).
Figure 101 – Quick-look on induced fibre strains in longitudinal direction in constant pitch configuration
As by definition, the bending radii in that configuration are constant (Figure 102).
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Figure 102 – Bending radii of 2 fibres in constant pitch configuration
As mentioned above, for an axisymmetric strain field, two fibres with different inclination angles at every
axial position are required to fully characterize the state of strain. In order to verify the developed
analysis procedure, a back-calculation has been done with the induced fibre strain values in the attempt
to reproduce the underlying strain field. The fibres extended over the full length of the bore. Figure 103
shows the result. It provides evidence for the correctness of the procedure.
Figure 103 – Original and back-calculated strains
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4.4.2.2.2 Variable pitch
As long as the strain field can be assumed to be axisymmetric, it may be sufficient to work with two
constant pitch helices. However, as pointed out in paragraph 4.3, at the presence of local debonds or
cracks, the assumption of axisymmetry no longer applies. The strain field then is affected and it varies in
hoop direction. In this case it would be desirable to measure discrete strain values at different axial and
circumferential positions at the overlap points of two fibres. In principle, this can be achieved in constant
pitch arrangement, albeit to a limited extent.
More flexibility is accomplished by the application of variable pitch helices. In some cases there may be
locations which are expected to be more prone to defects than others, and then it would be preferable to
have the checkpoints clustered in these areas.
Two exemplary configurations with variable pitch have been examined: one with a cluster at the central
position and another with the checkpoints clustered at the nozzle side of the propellant grain. For the
sake of brevity, results of only the former configuration are presented here.
Figure 104 – Example of variable pitch helices clustered at the central position
The bending radii now are not constant (see Figure 105).
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Figure 105 – Bending radii of variable pitch helices clustered at central position
The induced fibre strains vs. fibre longitudinal coordinates are given in Figure 106.
Figure 106 - Induced fibre strains in longitudinal and transverse direction in variable pitch configuration,
clustered at central position
It should be noted that at increasing degree of clustering the inclination angle to the hoop direction
decreases and the induced fibre strains approach the hoop strain. For a given type of fibre, the shape of
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the helix has to be adjusted as to meet the strain capability, i.e. a fibre with low capability would require
for larger angles (flat helix) than one having higher capability.
The developed analysis procedure allows to find and visualise the overlap points (see Figure 107 and
Figure 108) which might be used to gauge local values in a non-axisymmetric strain field.
Figure 107 - Overlap points for centrally clustered variable pitch helices
Figure 108 – Distribution of overlap points in z- plane for centrally clustered variable pitch helices
Figure 109 shows an example of a configuration clustered at the nozzle side.
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Figure 109 – Distribution of overlap points in z- plane for nozzle side clustered variable pitch helices
When the overlap points are to be used for the evaluation, both fibres have to have a grating at these
locations. It remains to clarify on if and how the fibres interact and what impact this might have on the
measurements.
Figure 110 - Inclination angles and intersection angles for centrally clustered variable pitch helices
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At the overlap points it is of interest to get the inclination angles  and the intersection angles 
cross
of
the fibres. Figure 110 gives an overview for the presented example.
For the axisymmetric debond scenario of paragraph 4.2.1, an investigation has been made with a
centrally clustered configuration similar to Figure 104. The results are shown in Figure 111.
Figure 111 – Fibre strains in longitudinal and transverse direction for the axisymmetric debond scenario
A pronounced decrease in fibre longitudinal strains compared to the undebonded motor is found (dotted
lines).
4.4.3
MECHANICAL INTERACTION BETWEEN FIBRE AND PROPELLANT
The mechanical interaction between fibre and propellant is governed by the stiffness of the fibre. Ideally,
a fibre with zero-stiffness would perfectly resolve that problem. At the moment, only limited information
is available on the mechanical characteristics of the individual fibre types. In [4], which deals with POFs,
initial moduli of approx. 5 GPa are reported (silica fibres are expected to have significantly higher
moduli).
In order to get insight into the interaction, a helical shaped fibre with the mentioned stiffness, 0.125 mm
diameter, an assumed Poisson ratio of 0.49 and a constant CTE of 7*10-6 K-1 has been implemented in
the 3D FEM model. It extends over 870 mm from the front end of the bore and consists of four coils (see
Figure 112).
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Fibre
Figure 112 – Finite Element model with integrated helical shaped POF
For better visibility, only half of the model is posted. In addition, the fibre thickness is scaled by the factor
20. The fibre is modelled by linear Bar-elements and is fixed on the propellant by a glued contact
condition. Potential influences of adhesive material have been disregarded.
As can be seen from the results of the induced strains and stresses, the interaction both during
cooldown phase and at soaked condition (i.e. the temperatures have become stationary) is insignificant
wth the considered POF. This is due to the low fibre stiffness. If there were siginificant interactions,
these would appear as disturbances of the strain and stress field in the vicinity of the fibre. To
demonstrate this, the induced hoop strains and stresses are plotted in Figure 113 and Figure 114
respectively for t=2h, 4h and 7.5 h. The plot ranges are individually adapted to the corresponding
conditions in the bore to get a high resolution.
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t=2 h
t=4 h
t=7.5 h
Figure 113 – Distribution of induced hoop strains in bore at t=2h, 4h and 7.5 h (deformation scale factor=5)
t=2 h
t=4 h
t=7.5 h
Figure 114 - Distribution of induced hoop stresss in bore at t=2h, 4h and 7.5 h (deformation scale factor=5)
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ASPECTS OF ANGULAR SENSITIVITY AND PRECISION
As mentioned in paragraph 4.4.1, any angular deviation of the fibre orientation angle from its nominal
value leads to deviations in the measured strain. As can be seen from Figure 94, where misalignment
angles d of ±1° and ±3° are regarded, the maximum sensitivity is encountered at angles between the
two principal directions. Close to the principal directions, the sensitivity on angular deviations
approaches zero.
An associated aspect is the precision of the fibre measurement itself. A given relative or absolute
precision may be related to an equivalent sensitivity angle. For the exemplary motor and load condition,
Figure 115 illustrates this relation for a number of selected relative and absolute gauge precisions. Both
criteria (angular sensitivity and precision) suggest one should keep close to the principal directions. In
an extreme case, one fibre would be placed circular, whilst the other one would be aligned in axial
direction. However, there are some restrictions on doing so:
The principal directions exhibit the highest strains (positive or negative). This may conflict with the strain
capability of the fibre. Further on, a fibre in hoop direction only covers one particular axial position, while
an axially aligned fibre covers only one hoop position. For monitoring the entire bore, a large quantity of
fibres would be needed.
Another aspect arises when a measurement is going to be performed at the overlap point of two fibres.
In this case it would be preferable to have that point at a position with high angular sensitivity in order to
achieve a broad range of possible intersection angles. In contrast to the former consideration, this would
suggest fibre inclination angles in between the principal directions.
In search of the optimum configuration, one has to consider aspects like the maximum (minimum)
expected strains in the carrier material, the fibre strain capabilities, whether the strain field is
axisymmetric or not, the intersection angles at the overlap points of two fibres (when a local
measurement is needed), where areas prone to debonds or cracks are expected, and, last but not least,
the way of installation.
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Figure 115 – Equivalent sensitivity angles for the example motor and load condition
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SUMMARY AND WAY AHEAD
A failure mode analysis was performed to investigate the impact of typical defects in a SRM grain on the
surface strains of its cylindrical bore. A realistic motor grain design for tactical applications, subjected to
an exemplaric thermal load, was used for demonstration purposes. The defects under consideration are
debonds of the grain from the insulation and cracks within the propellant. Purpose of the analysis was to
clarify whether the installation of OF sensors on the free-standing surface of the grain (bore) would allow
to monitor the strain field and/or discover grain defects when they are generated throughout its service
life.
The flawless configuration has been analysed and taken as a reference. The results also show what
crack creates what deviations in the strains (hoop and axial) compared to the intact motor (Table 1).
An analysis tool has been developed which allows to configure analytically the installation pattern of
fibres on the surface of the grain, thereby showing through which configuration any of the
aforementioned defects could be observed best.
The analysis shows that in principle, typical macroscopic defects of sufficient size in a motor can be
reliably detected by properly embedded OF sensors and indicates what characteristics this sensor
should have.
Firstly, the sensor should not alter (e.g. decrease) the mechanical load state of the propellant. A first 3D
analysis, using a POF in helical configuration, shows that the POF would not significantly disturb the
mechanical response of the propellant. Silica optical fibres would probably disturb the strain field
because of their comparatively high stiffness.
Secondly, the fibre should have itself sufficient strain capability to survive the strain induced in the
propellant even in absence of cracks. In general, POF systems potentially offer a larger strain range
measurement capability than silica optical fibres can do. In [4], 6 % strain before failure is reported, and
strains up to 13 % are considered as a potential for good quality material.
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The effect of cracks on the strain field could easily be detected for all crack configurations, depending
on the orientation and position of the POF. The formation of a crack could even break a surfaceembedded POF if the fibre was bridging the crack mouth.
Finally, to fully observe the strain field on the bore of the SRM grain, one needs to apply POFs in two
different directions, i.e. at a mutual angle. This assuming that the principal straining directions are known
through a previous analysis.
As it appears, fibres with low stiffness and large strain capability are required for applications in SRMs.
The POF (polymer optical fibre) type seems to be a promising candidate. Full monitoring of both the
undisturbed strain field and the presence of defects would require the application and interrogation of
several POF sensors.
For the time being, the fibres are assumed to perfectly follow the carrier material. Accompanying work
has to be done on the mechanical and chemical characterisation of various types of fibres, i.e. adhesive
capability on the propellant, stiffness, strain capability, measurement precision and chemical
compatibility with the propellant.
With these data available, further analysis works are necessary to study the mechanical interaction of
the fibres embedded in propellant specimens and even within the propellant grain, when being subjected
to thermally and mechanically induced strains.
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5. SENSORS EFFECTS ON MAGNETIC MICRO WIRES FOR STRUCTURAL MONITORING
IN THE SESAMO PROJECT.
5.1
AMORPHOUS GLASS COATED MICROWIRES
Amorphous glass coated microwires are novel materials that are very perspective for technical
applications. They consist of metallic nucleus (of diameter 1-30 μm) that is covered by the glass-coating
of thickness 2-20µm (see Figure 116). They are prepared by Taylor-Ulitovski technique [1-3] by a rapid
quenching of molted master alloy and drawing.
Their biggest advantage of glass-coated microwires is their production costs. From 1 kg of
master alloy, it is possible to prepare 40 000 km of microwire. Practically, 10 km of microwire is
produced from 0.25 g of master alloy in one shoot. It is possible to prepare different chemical
composition: Cu wire, Pt wire, Ag wire for electrical purpose, or magnetic wire based on Fe, Co and Ni.
Figure 116 TEM image of amorphous glass-coated microwire [1].
Another advantage is the glass coating that prevents from electrical short-cuts up to 600 oC. Moreover, it
prevents metallic nucleus from corrosion.
And finally, their dimensions allow production of sensors on micrometric scale.
However, the biggest interest (at present) is focused on amorphous magnetic microwires based on Fe
and Co compositions. They allow for build up microsize contact-less sensors of temperature, mechanical
stress, frequency etc.
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Due to their amorphous structure, the amorphous microwires are characterized by very low anisotropy
(since the magnetocrystalline anisotropy is missed). Hence, their magnetic properties are determined by
magnetoelastic and shape anisotropy:


(1)
Where s is saturation magnetostriction and  represents the mechanical stresses. Magnetoelastic
anisotropy arises from magnetoelastic interaction of magnetic moments with the mechanical stresses
introduced into the microwires during their production process. Mainly there are axial stresses
introduced by drawing, radial and circular stresses introduced by quenching (fig.2). Moreover, additional
stresses are introduced by different thermal expansion coefficient of metallic nucleus and glass-coating.
It was shown [4] that axial stresses prevail in the center of metallic nucleus, whereas the radial stresses
are strongest just below the surface. Such a stress distribution is extremely important for magnetic
properties of glass-coated microwires (as it will be shown later).
Figure 117 Distribution of mechanical stresses introduced into the microwire during their production.
Shape anisotropy is given mainly by the dimensions of microwires (small diameter of metallic nucleus –
few µm- compared to its length – few mm). Hence the demagnetizing factor in the axial direction is zero
and demagnetizing factor in the direction perpendicular to the wire’s axis is 0.5. Therefore, the easy axis
for magnetization will be identical with the wire’s axis.
According to eq.1, the amorphous magnetic microwires can be divided into next three main groups
based on their magnetostriction.
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AMORPHOUS GLASS-COATED MICROWIRES WITH NEGATIVE MAGNETOSTRICTION
These are mainly microwires based on high concentration of Co. Negative magnetostriction tries to
rotate magnetic moments perpendicularly to the applied mechanical stresses. Based on the abovementioned stress distribution (with axial direction in the center of the wire and radial below the surface),
the easy axis for magnetization will have circular direction (fig.3). The domain structure of such
microwires consists of domains with circular magnetization. As a result, the hysteresis loop in axial
direction will show no hysteresis and constant permeability within wide range of the fields (up to the
anisotropy field). Such a material is suitable for applications, constant permeability is required
(microtransformers, etc..).
Figure 118 Schematic figure of magnetization easy axis and hysteresis loop for glass-coated magnetic
microwires with negative magnetostriction. [1].
5.1.2
AMORPHOUS GLASS-COATED MICROWIRES WITH POSITIVE MAGNETOSTRICTION
These are mainly microwires based on the Fe alloys, having high and positive magnetostriction.
Positive magnetostriction prefers magnetic moment to be oriented in the direction of applied
mechanical stress. Based on the stress distribution given above (fig.2, fig.3), the easy axis of
magnetization have mainly axial direction (fig.4). As a result, the domain structure of such
microwires consists of big single domain with axial magnetization that is surrounded by the radial
domains (because radial stresses prevail below the surface- see fig.4). Magnetization process in
axial direction runs through the depining and subsequent propagation of single closure domain
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along entire microwire in a single Barkhaussen jump at the level of critical field. Hysteresis loop of
such microwire is perfectly squared (fig.4) and such microwires present magnetic bistability (e.g.
magnetization has just two values +Ms and - Ms).
Figure 119 Schematic figure of magnetization easy axis and hysteresis loop for glass-coated magnetic
microwires with positive magnetostriction. [1].
Bistable behavior of magnetic microwires can be employed in different applications for coding of person
and goods as well as for different sensors of temperature, stress, rotations, etc…
5.1.3
AMORPHOUS GLASS-COATED MICROWIRES WITH LOW MAGNETOSTRICTION
These are mainly microwires based on CoFe (3-5% of Fe) composition that has low but a little bit
negative magnetostriction. Since there is no crystalline structure (no magnetocrystalline anisotropy), low
magnetostriction will results in a lack of anisotropy at all (see eq.1). Easy axis is not well defined;
however it is circular just below the surface. The domain structure consists of lot of domains. Hysteresis
loops is very narrow (fig.5), it looks almost square-shaped, however magnetization can achieve any
values within the interval <-Ms ; + Ms >. Sharp change of magnetization at low fields points to a high
magnetic permeability that can be used in various sensors of magnetic fields, position, micro
transformers,
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Figure 120 Schematic figure of hysteresis loop for glass-coated magnetic microwires with low
magnetostriction. [1].
Having magnetostriction low, but a little bit negative, the easy axis will have circular directions, but low
anisotropy. As a result, the domain structure of such microwire consists of lot of circular domains, antiparallel oriented (Figure 121). Such a configuration is ideal for application based on new effect- Giant
Magnetoimpedance effect. It can be employed for very precise and sensitive microsensors of magnetic
field, electric current, stress, temperature, etc…
Figure 121 Schematic domain structure of amorphous microwires with low and negative magnetostriction.
5.2
SENSORS BASED ON THE SWITCHING FIELD
Monodomain structure of magnetic microwires with positive magnetostriction is ideal for
construction of various simple sensors and devices for coding, transfer and treatment of binary
information. Due to the shape and magnetoelastic anisotropy, the magnetization can have only
two values (+Ms and -Ms). Hence, one directly has two binary states. The change of the state
appears at the applied field that is called switching field. Switching field can be measured by very
simple contact-less induction method (Figure 122), that consists of excitation coil being fed by a
triangular shape current. It induces the linearly increasing magnetic field. When excitation field
reaches the value of the switching field, the closure domain wall depins and propagates along
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the wire, inducing the emf maximum at the end of the pick-up coil. The advantage of the method
is that the switching field is directly proportional to the time between the point when U 1=0 and
the time at which U2= Uswitch. Such a system can easily be digitized and automatized.
Figure 122 Schematic picture of induction method for switching field measurement.
5.2.1
TEMPERATURE DEPENDENCE OF THE SWITCHING FIELD
The several mechanisms for magnetisation reversal through the domain wall displacements have
been previously reported [5]. In particular, several sources of pinning mechanisms for the
magnetisation process by wall displacements have been identified contributing to the total
coercivity. They are:
1. Volume pinning of the domain walls, DW, by structural defects, which is particularly important in
magnetostrictive alloys,
2. Relaxation effects due to local structural rearrangements,
3. Clusters of chemically short-range ordered regions,
4. Surface irregularities
5. Intrinsic fluctuations of exchange energies and local anisotropy.
The effects are ordered in decreasing importance, the first one is assumed to have the biggest
contribution to the coercivity.
5.2.2
STRESS DEPENDENCE OF THE SWITCHING FIELD
Since the magnetoelastic anisotropy is the one that determines the magnetic properties of amorphous
microwires, they are very sensitive to the stress applied. Firstly, it is the stress dependence of the
switching field as a result of the change of the stress distribution. Secondly, the stress dependence of
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the switching field arises also from the relaxation contribution. The sum of these two contributions gives
a complex stress dependence of the switching fields (Figure 123)
Figure 123 The stress dependence of the switching field is given as a sum of two contributions:
magnetoelastic and structural relaxation.
However, both contributions to the stress dependence of the switching field can be tailored by the
properly selecting thermal treatment. Either stress distribution can be changed that influence
magnetoelastic contribution, or the structural defects concentration and its sensibility to the stress can
be tailored by thermomagnetic treatment. Figure 124 shows the stress dependence of the switching field
for amorphous and heat treated FeNiMoB microwire. The weak and non-monotonous stress
dependence of the as-cast microwire can be improved by thermal treatment at 450oC, after which the
switching field is very sensible and monotonous with the stress even in a very narrow range of the stress
applied. Another possibility shown in Figure 124 arose from annealing at 375 or 425 °C, after which the
switching field does not depend on the applied stress. It is because the partial crystallization of the
microwire. Precipitates of FeNi phase play role of strong pinning centers and the hysteresis
mechanism is no more given by the magnetoelastic non structural relaxation contributions. In this case
the switching field is given by inclusion theory model and depends only on the size and concentration of
the inclusions. Such thermal treatment is useful for sensors, where stress dependence must be avoiding
(e.g. temperature, frequency sensors).
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Figure 124 Stress dependence of the switching field Hsw in amorphous glass-coated FeNiMoB microwire
after different thermal treatments [9].
5.2.3
FREQUENCY DEPENDENCE OF THE SWITCHING FIELD
The switching field of amorphous magnetic microwire is also dependent on the frequency of the
applied magnetic field. Such dependence arises from the frequency dependence of the relaxation
function G(T,t) given by eq.6. as well as from the magnetic relaxation of magnetic moments that is
hindered by the Gilbert damping. As a result, complex non- monotonous frequency dependence of the
switching field appears that decreases at low temperatures as a result of the structural relaxation and
increases at higher frequencies because of the frequency dependence of magnetoelastic contribution.
The frequency dependence of the switching field in amorphous magnetic microwires can also be
tailored by thermal treatment and applied mechanical load according to desired conditions. In the ascast state, the frequency dependence of the switching field is not sensitive (or very weakly) to the
applied mechanical stress (Figure 125). Annealing at 300 oC introduces strong stresses into the
metallic nucleus of the glass-coated wire and the frequency dependence changes. Its slope increases
with the stress increased. At strong stresses (97 MPa), even frequency independent range (750-2000
Hz) can be found as a result of the change of the domain wall structure.
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Figure 125 - Frequency dependence of the switching field in as-cast amorphous FeNiSiB microwire.
Applied tensile stress as a parameter.
Figure 126 Frequency dependence of the switching field in amorphous FeNiSiB microwire annealed at
300oC. Applied tensile stress as a parameter.
5.3
BISTABLE SENSOR OF TEMPERATURE USING TC
Among continuous sensors described above, another type of threshold sensors can be based on
disappearance of the magnetic bistability. This happens, when the temperature of measurement
exceeds the Curie temperature. Hence, by properly choosing the microwire’s composition, the Curie
temperature can be adjusted according to desired conditions.
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BISTABLE SENSOR OF TEMPERATURE USING LOW MAGNETOSTRICTION
The disappearance of magnetic bistability can be observed also in the microwires with low
magnetostriction. Low-magnetostrictive microwires do not show bistability. However, when high stress is
applied on the wire axis, such bistability can appear as a result of increase of magnetoelastic interaction
(according to eq.1). This is the way, how to construct the threshold sensors for the stress application.
5.4.1
REFERENCES
[1] M. Vazquez, Physica B, 2001, vol.299, 302.
[2] Chiriac H., Ovari T.A., Progress Mater. Science, 1996, vol.40, 333.
[3] Zhukov A., Gonzalez J., Vazquez M., Larin V., Torcunov A., Nanocrystalline and amorphous
magnetic microwires, in: H.S. Nalwa (Ed.), Enciclopedia of Nanoscience and Nanotechnology,
American Scientific Publishers, 2004 p.23
[4] Chiriac H., Ovari T. A., Pop G., Phys. Rev. B, 1995, vol.52, 10104.
[5] H. Kronmüller, Phys. Stat. Sol. (b) 127, 531 (1985).
[6] Varga R, Garcia KL, Vazquez M, Zhukov A, Vojtanik P, PHYSICAL REVIEW B 70 (2004), 024402.
5.5
EXPERIMENTAL RESULTS FOR SESAMO APPLICATIONS
Four alloys were tested for stress and temperature measurements - Fe73Nb3Si11B13, Fe71Nb5Si11B13,
Fe49.6Ni27.9Si7.5B15 and Fe77.5Si7.5B15. Measuring system is very simple. One just needs to excite the
microwire with magnetic field and to sense by a single coil the induced maximum when the domain wall
propagates along the wire. This happens at a switching field. The switching field is sensible to external
parameters such a mechanical load, temperature, etc…
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Figure 127 Schematic diagram showing the experimental setup for the measurement of switching field
distribution in magnetic microwires.
The most promising results we get with alloy Fe73Nb3Si11B13. Figure 128 shows almost linear
temperature dependence of the switching field in the temperature range from -150oC up to 150oC. The
switching field varies from 380A/m down to 170A/m. It decreases more than half of its value allows for
high sensitive measuring of temperature.
Figure 128 Temperature dependence of Fe73Nb3Si11B13 microwire at 400Hz (metallic diameter d=8m and
total diameter =16m, length = 4cm).
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Figure 129 Temperature dependence of Fe73Nb3Si11B13 microwire at 280Hz
At lower frequency of applied field one can get less linear dependence of the switching field on the
temperature, however, the switching field increases almost 3 times from 140 A/m at 150oC up to 380
A/m at -200oC. The slope is higher, and the temperature dependence is still monotonous and well
defined. (It can be calibrated).
Figure 130 Temperature dependence of Fe73Nb3Si11B13 microwire at 20Hz
Figure 130 shows that by simply changing the frequency to 20Hz of the applied field one can get the
switching field invariant to the temperature in wide range- what is necessary for the stress sensor that is
invariant to the temperature. No temperature compensation is necessary. The change of the frequency
of applied field can be made by software – we do not need to change hardware of the sensor =>we can
get intelligent sensor for various application
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Figure 131 Stress dependence of the switching field for various frequencies of applied magnetic field.
Figure 131 shows that the highest stress dependence is for high frequency. On the other hand, the
temperature dependence decreases with the frequency .
THE BEST CANDIDATES FOR STRESS SENSOR
Fe77.5Si7.5B15 as cast
Fe73Nb3Si11B13 as-cast
THE BEST CANDIDATES FOR TEMPERATURE SENSOR
Fe38,75Ni38,8Si7,5B15 As cast
5.6 3 SENSORS DESIGN AND CONCLUSION
The design follows the requirements of the SESAMO project end users for contactless measuring of
strain up to 20% .
Physical principles of strain and temperature sensors were tested and suitable materials were found.
Some of the results were listed in the previous chapter. Sensors are designed of a short 1cm long
microwire embedded in hard rubber VUKOL. It is polyurethane casting substance (further PUR ZL)
serve for filling of voids of every type, especially at construction of electrical devices, where casting
resins without internal tension with good elastic and electric insulating properties are required. They
have good adhesion to glass, PVC and rubber. They are suitable for casting of transformers, electronic
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components to covers, cables ends and joints, for moisten prevention. Due to high elongation they
manage thermo dilatation stress and it is suitable to cast materials with various thermal expansions.
Thermal endurance temperature is 120 – 130 °C. They do not contain solvents neither other volatile
substances. Their curing is characterized by low exothermal temperature. After curing they are
characterized by flexibility of hard gum.
Shape strain anisotropy of the sensor is provided by the carbon fiber to determine axis of strain
sensitivity. Other materials for embedding of microwires could be used to achieve good compatibility
with composite materials under strain test. The dimension of sensor prototype is about 10 x 10 x 0.1mm.
After placing the sensor into the test object, it is possible to sense the switching field by the excitation
and sensing coils outside, i.e. contactless.
Result of experiment confirmed that electromagnetic interaction is also in this configuration quite strong
and sensoric phenomena derived from the movement of domain walls occurs. Test result is shown in the
Figure 132. The output of the sensing coil are impulses, so the result of measurement is independent of
distance from the sensor implemented in the object under test up to couple of centimeters.
Figure 132 Signals of contactless nonconcetric sensor, U1 is voltage on excitation and U2 on pick-up coils
of the sensor. Measured parameter which is proportional to the applied stress is time t of impulse caused
by domain wall motion.
Based on the success of the first experiments, the excitation and sensing coil were designed as U shape
as shown in Figure 133.
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Figure 133 Excitation coil of contatless sensor, the pick up coil is transversal and located inside,
sensitive elements are out of a flat part of coil.
Sensitive elements, which can meet the requirements of the SESAMO project have been designed in
several prototypes – bended, linear and random, follow these basic images (Figure 134):
Figure 134 Basic designed prototype of sensing elments 10x10x0.1mm
The entire layout has hard rubber flexible base, carbon fibers and magnetic microwires are arranged for
measuring the horizontal axis, even for large values of strain.
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Figure 135 Arrangement of excitation coil, sensing coil and embedded sensor
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