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_________________________________________________________________
SLURRY SPRAYED THERMAL
BARRIER COATINGS FOR
AEROSPACE APPLICATIONS
_________________________________________________________________
Phuc Nguyen
A thesis submitted in fulfilment of requirements
for degree of Doctor of Philosophy
School of Mechanical Engineering
The University of Adelaide
May 2010
Appendices
191
APPENDICES
Appendices
192
Appendix A
Investigation of Thermo-Mechanical Properties of Slurry based Thermal Barrier Coatings under Repeated Thermal
Shock Phuc Nguyena, Andrei Kotousovb, Sook-Ying Hoc and
Stuart Wildyd
School of Mechanical Engineering, the University of Adelaide, SA 5005 Australia [email protected], [email protected],
[email protected], [email protected]
Keywords: thermal barrier coating, repeated thermal shock loading, thermo-mechanical properties, adhesion strength, thermal shock resistance, thermal stresses
Abstract. Thermal Barrier Coatings have existed for over 40 years, and within the last 15 years their use in industrial applications has dramatically increased. Thermal Barrier Coatings (TBCs) are currently used in gas turbines, diesel engines, throughout aerospace and nuclear power industries. The purpose of TBC is to reduce temperature and thermal stresses, and, as a result, increase the reliability and life of load-bearing components subjected to high temperature or temperature flux. However, TBCs often fail under thermal cyclic loading with reliability still being the major issue impeding their wide-spread applications.
The focus of this work is on experimental investigations of zirconia/nickel graded TBC system, subject to thermal shock loading. The graded TBC systems were fabricated utilising a recently developed slurry spray manufacturing technique. This is a robust technique, and is able to cover large and curved surfaces at low cost, and provides many advantages in comparison with its alternatives. This paper describes the developed technique and presents selected results of thermo-mechanical and fracture testing of the TBCs including graded coatings fabricated using this new technique.
Introduction Thermal Barrier Coatings (TBCs) represent a relatively thin layer of a material with high insulating properties, such as ceramics, that is bonded to a substrate, which is usually metal, to protect the metal load carrying structure during temperature excursions. The application of TBCs can significantly increase the operating temperatures up to 1400-1500ºC, increase efficiency and improve the durability of the components. There are many applications, which have benefited from adopting TBCs. These include the aeronautical, aerospace, automotive and nuclear industries and heavy-duty utilities such as diesel trucks [1].
Appendices
193
The development of TBCs has centred mostly on Partially Stabilised Zirconia (PSZ) due to its unique physico-mechanical properties and has been led by its use in aircraft-engine combustion-path components [2]. The significant advance in the development of an effective protective coating was associated with the development of Functionally Graded (FG) TBCs. FG-TBCs are multiphase composite materials that are engineered to a have a spatial variation of material constituencies. Using FG TBCs, as an alternative to joining directly together two dissimilar materials such as ceramics and metal, carries several advantages including: much lower thermal stress distribution across the thickness; minimisation of stress concentrations; and an increase in bonding strength. First, the new developed technique will be briefly outlined. The technique is also suitable for producing the FG-Coating. Examples of the FG-Coatings will be given in the paper. Experimental results on thermal cycling, adhesion strength, investigation of microstructure and effect of various manufacturing parameters on the quality, fracture and durability of the coating will be discussed. The paper will be concluded with a summary of major outcomes of the current experimental study and suggestions on future work.
Slurry Spray Technique The Slurry Spray Technique for manufacturing TBCs utilises traditional wet powder spraying methods to deposit sinterable coating materials onto target substrates to produce a functional coating. The process involves suspending the coating material within a fluid to form a slurry mixture that can be applied to a surface using common gravity fed spray guns. Successive layers are then sprayed onto the Inconel substrate and dried using varying slurry compositions. The optimal thickness of the layers to deter surface cracking during the drying process is approximately 100 µm (which can be seen in Fig. 1) and the drying time is approximately an hour, depending on ambient conditions. After the desirable number of layers of the TBC is deposited the multi-layered coating is loaded in a compression chamber to form a densified layer before being sintered with an acetylene torch or furnace. The applied pressure varies depending on the number of coating layers, typically between 10 and 40 MPa. Details of this technique can be found in [3] and [4].
Examples of Slurry based Thermal barrier coatings Below we describe several examples of TBC fabricated using the developed manufacturing technique, which can be seen in Fig. 1.
Appendices
194
(a) MonoLayer Coating 1 (b) MonoLayer Coating 2 (c) FG – Multilayered Coating
Figure 1. Cross-sectional view of a Slurry Sprayed TBC’s
An example of a monolayer coating, fabricated utilising the Slurry Spray Technique, can be seen in Fig. 1(a, b). The composition of the monolayered coating is 50% ZrO2 – 50% Ni, which can be distinguished by the different grain structure. The nickel substrate can be seen on the left hand side of Fig. 1(b, c), and the epoxy resin on the right hand side of the image; the epoxy resin was used to set the TBC specimens for SEM investigations. In Fig. 1(c) an example of a FG – Multilayered TBC can be seen. This coating consists of 2 layers, with the initial layer of the FG-TBC composition consisting of 50% ZrO2 – 50% Ni, and the top layer of the FG-TBC consisting of 100% ZrO2 – 0% Ni. In Fig. 2(c), the layer is more pronounce than the monolayer coating seen in Fig. 2(b), this is due to the mechanical densification of the coating during the fabrication of the FG-TBC [5].
Experimental Results Adhesion Test. Adhesion tests of the adhesion strength of the various TBC compositions were carried out with the PosiTest pull off adhesion tester. The results showed that FG-TBC were a 100% improvement from the monolayered TBC with 100% of ZrO2 and 25% improvement from the monolayered TBC with 50% ZrO2 - 50% Ni (Fig. 2). The maximum adhesion strength obtained through the experiment, for the FG-Coating was approximately 11 MPa. However in comparison with existing coating techniques such as the flame spray method, with adhesion strength of approximately 21 MPa, the adhesion strength of the FG-Coating produced is 50% lower than the flame spray method [6].
Substrate Ni Epoxy
MonoLayered TBC
Substrate Ni
MultiLayered TBC
Epoxy
Appendices
195
Figure 2. Adhesion strength range for various types of coatings
Thermal Cycling Test. The purpose of the thermal cycling tests was to give an indication of thermal fatigue behaviour of the Slurry based TBC [7]. The maximum temperature reached was 900°C, with a 30 minute heating/cooling cycle. The monolayered TBC with 100% of ZrO2, 50% ZrO2 – 50% Ni and FG-Coating with two layers of and 50% ZrO2 – 50% Ni and 100% ZrO2 were subjected to the thermal cycling.
The experimental study showed that the majority of the failures occur during the first few cycles (see Fig. 3). If the coating survived the first 4-5 thermal cycles, the coating integrity is preserved throughout the following thermal cycle. Fig. 3 demonstrates that FG-Coatings are normally much more durable and better resistant to the thermal cycling. The effect of the first few cycles can be explained by the manufacturing defects, which lead to the almost immediate failure of the coating. The FG-Coating has much lower probability of failure during the thermal cycling, which can be explained by the lower mismatch in material properties and the lower level of thermal stresses during the sintering of the TBC.
Figure 3. Ratio of the survived TBC to the total number of tested samples for FG-Coating and Monolayered Coatings
The results obtained through the experiment coincide with theoretical analysis presented in literature reviews where TBCs are expected to perform better in real life application if manufactured in a controlled sintering and cooling environment. Firstly, by applying constant heat flow, uniform heat expansion, and ideal boundary grain growth between particles is achieved thus reducing thermal stress induced
0
4
8
σad, MPa
50% ZrO2 – 50% Ni
100% ZrO2
50% ZrO2 – 50% Ni
100% ZrO2Delamination cracking
F F
TBC
0
0.5
1
0 5 10 15 20 N of cycles
Functionally graded coating
Monolayered
Appendices
196
during the thermal expansion process. Secondly, a slow cooling rate after sintering effectively reduces the strain and stress associated with rapid cooling. Thirdly, introduction of FG-Coating induces a temperature gradient across the coating hence minimising thermal mismatch due to cooling and the resulting residual stresses.
(a) (b) (c) Figure 4. Microstructure of TBC after fabrication (a), 10 thermal cycles (b)
and 20 thermalcycles (c), magnitification – 500x.
From SEM images taken after the fabrication, 10 and 20 thermal cycles (see Fig.4), it can be seen that the increase of porosity with the number of thermal cycles and formation of crack damage, eventually propagates through the thickness and lead to the failure of the coating.
Conclusion This paper presents results of an experimental study on the thermo-mechanical properties of TBCs fabricated using a new method based upon the Slurry based TBC technique. The main advantages of this technique are the low costs and the ability to cover large and curved surfaces, which are critical for a number of important practical applications. This technique allows the fabrication of the multilayered FG-TBC’s, which can significantly reduce the thermal stresses and, and as a rule, have higher durability and lower failure rates in comparison with monolayered TBCs.
The test results demonstrate a satisfactory adhesive strength of the coating, which is comparable with the adhesive strength of other coatings, fabricated using traditional techniques such as the Flame Spray method. The outcomes of the experimental study also showed that the FG-TBC fabricated using the Slurry Spray technique are able to survive low-cycle thermal excursions, when the temperature increases up to 1000°C. Further work will focus on real-life applications, such as high temperature burner tests and leading edge of scramjet propulsion systems.
Acknowledgments The authors acknowledge, with thanks, the financial support from the U.S Air
Force, without their support this research would not have been possible.
Appendices
197
References
[1] Koizumi, M. (1997), Composites Part B; Engineering, 28(1-2), 1-4. [2] Martena, M., Botto, D., Fino P., Sabbadini S., Gola M.M., Badini C. (2006), Engineering Failure Analysis, 13 (3), 409-426. [3] Nguyen, P., Harding, S., Ho, S-Y. (2007) ACAM, 1, 545-550. [4] Ho, S-Y., Kotousov, A., Nguyen, P., Harding, S., Codrington, J., Tsukamoto, H., (2007) Scientific and Technical Information Network, Defense Technical Information Centre. [5] Dahl, P., Kaus, I., Zhao, Z., Johnsson, M., Nygren, M., Wiik, K., Grande, T. & Einarsrud, M.A. (2007) Ceramics International, 33, 1603-1610. [6] Davis, J. R. (2004) Handbook of Thermal Spray Technology, Thermal Spray Society and ASM International, United States of America. [7] Zhu, D., Choi, S.R., Miller, R. A. (2004) Surface and Coatings Technology, 188-189, 146-152.
Appendices
198
Appendix B
Induction heating apparatus for high temperature testing of thermo-mechanical properties
by J. Codrington*, P. Nguyen, S.Y. Ho, and A. Kotousov
School of Mechanical Engineering
The University of Adelaide, SA 5005, Australia
*Tel: +61 – 8 – 83033177; Fax: +61 – 8 – 83034367 E-mail: [email protected]
ABSTRACT— A low cost high temperature test facility designed and built for
the purpose of thermo-mechanical testing is described. An induction heater
provides variable heating rates, simple operation and easy access for
temperature and strain measurement. Specially designed high temperature
specimen grips with water-cooling allow for testing over long periods of time.
Contact temperature and strain measurements are utilised to provide
accurate and reliable results. Detail is given on the experimental procedure
including calibration of the thermocouple temperature measurement. A
validation study of the thermal expansion and tensile Young’s Modulus of
carbon steel 1020 at temperatures up to 850°C prove s the accuracy of the
test set-up and procedure. Results are given for the stress-strain curves of
aluminium alloy 7000 T4 at various temperatures to further demonstrate the
capabilities of the test facility. The measured thermo-mechanical properties
Appendices
199
of these materials were used to develop high temperature constitutive
models for implementation in finite element thermal-structural analysis of
hypersonic structures.
KEY WORDS— High temperature, Mechanical test, Thermo-mechanical
properties, Induction heating
1 Introduction
Advanced applications are emerging that require high temperature materials and
structures that are able to withstand conditions above 1000°C. These applications
include hypersonic and supersonic aircraft, anti-terrorist measures, welding
technologies, the mining industry and many others. At elevated temperatures
significant changes occur in the thermal and mechanical properties of materials.
These properties are the basis of thermo-structural design calculations and allow for
development of accurate finite element (FE) models. High temperature material
properties that are suitable for development of constitutive models in FE analysis are
not readily available in the literature. The primary motivation for the present study is
to develop a relatively low cost, simple and reliable method to measure thermo-
mechanical properties in the high temperature regime for use in thermal-structural
analysis of hypersonic structures.
The main constraints placed on high temperature mechanical test facilities are
usually based on the high cost and limited availability of test material. As a result
small test specimens are favoured. However, at high temperatures external forces that
are usually considered negligible can greatly affect the mechanical behaviour of
small specimens. For example, contact forces from temperature or strain
measurement equipment, such as thermocouples or mechanical extensometers, can
produce large stress concentrations. Additionally, the loads required to hold the
sensors against the specimen surface are often sufficient enough to cause distortion
or bending of the specimen. This places restrictions on the chosen test equipment in
Appendices
200
particular the temperature and strain measurement techniques. A review of various
high temperature mechanical test methods was presented by V�lkl and Fischer [1].
Details were also included of their own specially designed facilities [1,2] for testing
metallic materials at temperatures up to 3000°C using non-contact temperature and
strain measurement. A further review and summary of current techniques is provided
in this paper.
An important requirement of the test facility is to be able to provide a
controllable heat distribution over the test specimen gauge length, while still
allowing access for temperature and strain measurement. A means of gripping and
applying loads to the high temperature specimen is also necessary. The grips must
not affect the specimen’s mechanical behaviour and should also prevent heat damage
to any load equipment. The main methods of heating the specimen include single and
multi-zone furnaces, induction heating and ohmic heating. Reppich et al. [3]. used a
single zone furnace with the specimen grips placed inside the heated zone. The grips
were made from alumina (Al2O3) and experienced temperatures up to 1400°C. On
the other hand, Ho and MacEwen [4] utilised a three-zone furnace, which allowed for
low cost stainless steel grips to be used without creating significant temperature
gradients over the specimen gauge length. As an alternative to furnaces, induction
heating and ohmic heating both provide localised heating within the test specimen
and allow for fast heating and cooling rates. Policella and Pacou [5] made use of an
induction heater for tension and torsion testing of metals at temperatures of up to
1000°C. However, Ohmic heating was chosen by V�lkl and Fischer [1] for its
simplicity and for full access to the specimen gauge length. The localised heating
also meant that low cost copper grips could be used.
Both contact and non-contact temperature measurement techniques are
commonly used in high temperature test facilities. Contact temperature methods
include thermocouples and resistance temperature detectors (RTD). Lee et al. [6]
used thermocouples located at the top and bottom of the steel tensile specimens for
temperatures of up to 800°C. A platinum resistance thermometer was chosen by Aria
and Yamazawa [7] for high stability control of their furnace to temperatures up to
1000°C.
Appendices
201
For precise temperature measurement intimate contact is required with the
specimen throughout testing. This can affect the temperature distribution and
mechanical behaviour of the test specimen. Non-contact temperature measurement
techniques avoid these problems and include radiation thermometers and optical
pyrometers. Temperature measurement can also be made indirectly by using
thermocouples, for example, to measure environmental conditions such as the air
temperature in a furnace. V�lkl and Fischer [1] used an infrared pyrometer operating
at wavelengths of 0.7 to 1.1 �m for temperature measurement between 750 to
3000°C. For accurate radiation temperature measurement spectral emissivity data is
required as function of temperature for the materials to be tested. This problem was
overcome by Potdar and Zehnder [8] who calibrated an infrared detector against
thermocouples spot welded to a test specimen’s surface. Calibration can also be
achieved by using a reference material of known spectral emissivity. Neuer and
Jaroma-Weiland [9] recommended the use of various Pt/Rh alloys for which they
measured the total and spectral emissivity at temperatures of up to 1350°C.
Contact strain measurement techniques include the use of strain gauges and
mechanical extensometers. Lei et al. [10] compared the use of resistance strain
gauges and PdCr wire strain gauges, along with various attachment methods, at
temperatures up to 800°C. Alternatively, mechanical extensometers allow strain
measurement at temperatures of up to around 1500°C by using extension rods to
remove the displacement sensor from the heated specimen. Cooling of the sensor can
provide a further increase in the range of operation. Reppich et al. [3] used an axial
extensometer with Al2O3 extension rods and an inductive linear position sensor for
tensile tests at temperatures up to 1400°C. A low cost method of modifying room
temperature extensometers was offered by Quesnel and Tsou [11] for use at
temperatures up to 500°C. Lee et al. [6] also used a modified room temperature
extensometer for tensile tests of steel at temperatures up to 800°C. Contact strain
measurement techniques can be utilised as non-contact methods by measuring
displacement outside of the heated gauge length. This however requires knowledge
of the mechanical behaviour of the entire region between the measurement points.
Ho and MacEwen [4] measured displacement by attaching an extensometer at the
Appendices
202
shoulders of a heated tensile specimen under the assumption that the total measured
deformation is mainly due to the significant plastic flow within the gauge length.
Non-contact strain measurement techniques include laser speckle, laser
extensometers and computer-vision systems such as digital-image correlation (DIC)
or video extensometers. Lyons et al. [12] utilised DIC by coating test specimens with
BN or alumina to create black speckles on the surface. Strain was then measured at
temperatures of up to 650°C by comparing images of the specimen’s surface. The
performance of DIC methods can be affected by various factors such as image
distortion, increased illumination and changes in the specimen surface. These
decorrelation effects were reduced by Anwander et al. [13] who combined both DIC
and laser speckle techniques. Two laser diode beams created speckle patterns on the
specimen surface and the displacement of these patterns was measured. This allowed
strain to be measured at temperatures up to 1200°C. V�lkl and Fischer [1] measured
strain with a digital camera and the computer software SuperCreep [1] at
temperatures of up to 3000°C. The software determined the distance between
physical markers on the test specimen to provide a strain measurement. Optical fibre
sensors provide another means of strain measurement that combines both contact and
non-contact methods, and can also be used for temperature measurement. Elster et al.
[14] used Fabry-Perot sensor elements to measure strain during fatigue tests at
temperatures of up to 1100°C.
The design constraints placed on high temperature test facilities are usually in
contradiction with the desired characteristics of low cost and versatility. High
precision and reliability generally means expensive and specialised test equipment. A
balance between the requirements placed on test equipment, including cost, accurate
results, and simple set-up and operation therefore needs to be found.
This paper describes a low cost high temperature test rig that was developed in
this investigation. The versatile facility can be used for a range of thermo-mechanical
tests and is simple to set-up and operate. Details of an experimental validation study
are given for the thermo-mechanical properties of carbon steel 1020 at elevated
temperatures up to 850°C. This study includes a comparison of the experimental
results obtained with already published data to verify the test setup. An experimental
Appendices
203
study of aluminium alloy 7000 T4, a material used in aerospace applications, is also
presented to further prove the reliability of the test rig. This material is tested at
temperatures of 260°C and 480°C and demonstrates the rapid deterioration in
mechanical properties with temperature. The thermo-mechanical properties
determined in both experimental studies are the linear coefficient of thermal
expansion and the tensile modulus of elasticity, yield strength and ultimate strength.
These properties are used to develop high temperature constitutive models for
thermal-structural analysis of hypersonic structures.
2 Description of the Apparatus
A low cost test rig has been designed and built for the investigation of high
temperature thermo-mechanical properties. The test rig consists primarily of an
induction heater, high temperature specimen grips and a control system. Loading of
the test specimens is via an Instron (1342) test machine, although the design of the
high temperature grips allows the use of any standard test machine. Strain is
measured by a high temperature mechanical extensometer with ceramic extension
rods. The test specimen temperature is measured by multiple type K thermocouples,
which provide the feedback signal for a PID temperature controller. The test rig is
shown in Fig. 1 and is described in further detail in the following sections.
Appendices
204
Fig. 1. Test rig
2.1 Induction Heater
A 3 kW induction heater provides an effective method of heating the test
specimen to temperatures up to around 1500°C, depending on the material and
specimen size. Induction heating allows for rapid heating and cooling rates and can
be used with a range of specimen designs and materials. Metallic materials can be
heated directly, while ceramics and other materials can be heated indirectly using a
metal radiation tube, or similar device. The design of the induction heating coil
ensures localised and uniform heating over the specimen gauge length and provides
access to the specimen for temperature and strain measurement. Induction heating
Mechanical Extensometer
Thermocouples
Extensometer - Displacement Sensor
High Temperature Grip
Test Specimen in Heating Coil
High Temperature Grip Extensometer - Extension
Appendices
205
was also chosen for its ease of operation. The test specimen is heated in air with the
heating rate and temperature controlled by a feedback control system.
2.2 High Temperature Specimen Grips
Heat transfer from the heated specimen to the Instron test machine is prevented
by specially designed high temperature grips (shown in Fig. 2). Using an induction
heater means only the specimen is directly heated with the grips experiencing a lower
temperature then at the specimen gauge length. The grips therefore consist of a low
cost stainless steel main body with pull rods for connection to the test machine. A
small section of higher cost Inconel 601, a nickel-chromium alloy with a low thermal
conductivity of 11.2 Wm-1K-1, reduces the initial heat transfer from the specimen.
Excess heat is removed from the grips by water circulation through the main body
via a pump and cooling unit. The combination of a low thermal conductivity section
and water-cooling allows the grips to be used for testing over long periods of time.
The Inconel section of each grip is made up of two removable pieces to allow for
attachment of the specimen by a variety of methods, e.g. threaded end piece, wedge
blocks. A range of specimen shapes and sizes can therefore be tested without the cost
of having to remanufacture the grips.
Appendices
206
Fig.2. High Temperature Grip
2.3 Test Specimens
The test specimens used for these investigations were round tensile specimens in
accordance with ASTM E21 [15] and E8M. [16] A diameter of 6 mm and gauge
length of 30 mm was chosen as a compromise between the limited amount of test
material and the increased specimen size requirements for contact strain
measurement. Round specimens were used for simplicity to manufacture and to
ensure uniform heat distribution from the heating coil. A threaded end section was
used to attach the specimens to the high temperature grips.
2.4 Measurement and Control
Temperature is measured by three type K thermocouples along the length of the
specimen reduced section (Fig. 1). Thermocouples were chosen over optical methods
based on low cost and simplicity of use. Furthermore, the use of thermocouples
means that spectral emissivity data for the test materials need not be determined.
Temperature measurements are taken from the specimen surface and are calibrated
against measurements from a thermocouple placed inside a test specimen. A further
discussion of this is provided in section 3.1. This method allows specimen
temperatures of around 1700°C to be measured, depending on the material being
tested, with the thermocouples only reaching temperatures of up to 1300°C. Multiple
thermocouples along the specimen length ensure a uniform and steady heat
distribution throughout testing.
Strain is measured by a high temperature mechanical extensometer with high
purity Al2O3 (99.7%) ceramic extension rods and a strain gauge based displacement
sensor (see Fig. 1). The extensometer has a gauge length of 30 mm with a 50%
displacement range for tension and can be left in place throughout specimen failure.
For temperatures up to 1200°C the extensometer can be used uncooled and with
Appendices
207
cooling can be used at temperatures above this value. A mechanical extensometer
was chosen due to low cost and straightforward operation, enabling strain
measurement in real-time without the need for complex software. However, being a
contact measurement method limitations are placed on the size of the specimens that
can be tested.
A laptop computer and 16-bit microcontroller are used to monitor and control all
test equipment. The microcontroller runs a C based code written by the authors and
interfaced through the computer. A PID feedback control system is included in the
code to control the induction heater and hence specimen temperature and heating
rates. For these investigations a sample rate of 100 Hz was used to ensure that the
digital PID implemented with the microcontroller best emulated the analogue PID
used for its design. All test variables and measurements are output from the
microcontroller to the computer for data logging. The Instron test machine used for
these investigations has its own control computer and software which allows for both
strain and load control. The output from the load cell, however was data logged
through the microcontroller.
3 Experimental Procedure
3.1 Thermocouple calibration
Temperature is determined by the measuring junctions of the thermocouples
being held against the specimen’s surface. The thermocouples apply only a minimal
contact force and are free to move with the strained specimens. This is to ensure that
any affect on the specimen’s mechanical behaviour will be negligible. Heat is then
transferred to the thermocouples mainly via conduction and radiation, depending on
the material being tested. At the same time heat is lost from the thermocouples due to
radiation and convection to the surroundings. Heat is also lost by the non-perfect
contact between the thermocouples and the test specimen. A theoretical study of the
heat transfer to the specimen and to the thermocouple and heat loss to the
Appendices
208
surroundings was performed by finite element (FE) transient thermal analysis of a
3D model of the specimen with a thermocouple on the surface (Fig. 3a). A heat flux
of 4.8 x 105 W/m2K from the induction heater was used for the initial analysis and
was varied by a factor of 0.5 - 2 to investigate the effect of heating rate. The results
from the FEA confirmed that the steady state temperature was achieved very quickly,
because of the rapid heating rate, and the temperature distribution across the
specimen was fairly uniform (see Figure 3b). The heat loss from the thermocouple to
the surroundings by radiation and convection was quite significant and can result in a
lower reading than the actual temperature on the specimen surface. Figure 4 shows
the theoretical temperatures, from the FEA, at the surface and centre of the specimen
and at the surface of the thermocouple. The temperature of the thermocouple at the
surface of the specimen was 100 – 250 degrees lower than the actual temperature of
the specimen surface. The thermocouple placed inside the centre of specimen,
however, would have much better contact with the specimen and heat loss by
convection is expected to be minimal.
Appendices
209
(a)
Appendices
210
(b)
Fig. 3 — 3D Finite element thermal analysis: (a) 3D model and (b)
Temperature distribution across the specimen
Appendices
211
0
500
1000
1500
0 500 1000 1500
SurfaceThermocouple
Temperature (°C)
Cen
tre
Tem
pera
ture
(°C
)
Fig. 4 — Comparison of the theoretical specimen temperature and
thermocouple measurement
Calibration was achieved by comparison of the surface temperature readings with
that of a centre temperature measurement. The centre thermocouples were placed
inside a 30 cm deep hole drilled along the specimen centreline. Any direct heating
affects of the induction heater on the thermocouples were neglected as the resulting
temperature levels were well below that of the actual specimen and surface
temperature measurements. Fig. 5 shows the calibration data for carbon steel 1020
with a third-order best-fit curve to allow for extrapolation above the centre
thermocouple’s operational limit. A similar overall trend can be seen to that of the
theoretical curve presented in Fig. 4. The differences in the curve shapes below
approximately 800°C can be partially explained by the non-linear heating effects of
the induction heating process, which were not considered in the FE analysis. The test
Appendices
212
results were also used to calibrate the induction heater power level for use in the PID
control system. Over time oxidation of the thermocouple and specimen surfaces can
lead to small changes in calibration. Average measurements over extended test
periods were therefore used to account for any variations with time.
0
250
500
750
1000
1250
1500
0 200 400 600 800 1000
experimental databest-fit curve
Surface Measurement (°C)
Cen
tre
Mea
sure
men
t (°C
)
Fig. 5—Thermocouple calibration curve for carbon steel 1020
Thermocouple calibration was verified by experimentally determining the Curie
and melting temperatures for carbon steel 1020. The Curie temperature was found
during heating the specimen by the point at which magnetic heating affects, or
hysteresis, ceased. A Curie temperature of 790°C was determined compared to a
value of 760°C given by Smithells [17]. The Curie point can also be observed in Fig.
5 where the calibration curve changes shape. A melting temperature of 1420°C was
determined compared to a value of 1455°C given by the material’s manufacturer
Onesteel [18]. This gives an error of ±4% in the temperature measurement compared
Appendices
213
to the published values. Taking into account inherent differences of the magnetic and
thermal properties between the various carbon steel 1020 specimens; this level of
accuracy was considered to be very good.
3.1 Test Procedure
The described high temperature test rig was used for thermal expansion and
tensile testing of metals. Although, it can also be used for a range of other
mechanical tests such as compression, fatigue and creep. The general test procedure
involves preparation of the test specimen, placement of the specimen in the high
temperature grips and set-up of the heating and measurement equipment. The desired
test temperature is set with the PID controller and the specimen is heated until at
steady state. For these investigations a heating rate of approximately 10°Cs-1 was
used with a total heating time of 3 mins to ensure uniform temperature distribution
through the specimen thickness and over the gauge length.
The high heating rate from the induction method is appropriate to that seen in
hypersonic flights where the heat flux from stagnation heating, shock/boundary layer
interaction, etc. can be of the order of 106 – 109 W/m2K. Faster, or slower, heating
rates are also possible using the induction method. The affect that the heating rate has
on the mechanical properties was not investigated in this study. Instead the same
heating rate was employed for each of the specimens and the final tests were carried
out once the temperature distribution was uniform in the gauge section. However, the
effect of heating rate is not expected to be significant because the induction method
rapidly heats up the specimen and FEA shows that a near uniform temperature
distribution is reached within 20-60 s. Finite element heat transfer analysis of the
specimen showed that a heating time of 3 mins was sufficient for uniform
temperature distribution through the specimen thickness and gauge length. This is
also supported by temperature measurements at various positions in the specimen
during the induction heating. For this study a temperature variation of 50ºC over the
specimen gauge length was considered acceptable, although a lower temperature
Appendices
214
variation is achievable.
Procedure for the thermal expansion tests was based on ASTM E831 [19] and for
the tensile tests on ASTM E21 [15] and E8M [16]. Thermal expansion tests were
undertaken with no applied load and the thermal strain was measured once
temperature was at steady state. During heating the specimens for tensile testing
thermal expansion was allowed by using load control set at zero load. Load control
was also used to apply a small tensile load at a constant load rate for the
determination of the modulus of elasticity. Displacement control was used for the
tensile failure tests to apply a large displacement to the specimen at a constant strain
rate. All measurements and test data were recorded by the computer for post-
processing.
3.3 Validation Study
To verify the test rig the linear coefficient of thermal expansion and tensile modulus
of elasticity were determined as functions of temperature for carbon steel 1020 (Fig.
6). Tests were undertaken at temperatures up to 850°C and compared with literature
data [17,20]. Room temperature of 25°C was used as reference for the thermal
expansion coefficient calculations and the elastic tensile tests were performed at a
constant load rate of 5 MPa s-1. The experimental values obtained are within ±4%
and ±5% of the literature data for the expansion coefficient and elastic modulus,
respectively. At temperatures above 850°C the limitations of contact strain
measurement were experienced with the contact forces from the extensometer
affecting the small test specimens.
Appendices
215
10
11
12
13
14
15
0 150 300 450 600 750 900
according to Smithells
Temperature (°C)
Line
ar C
oeffi
cien
t of T
herm
al E
xpan
sion
(�m
/m °C
-1)
experimental resultswith error bars of ±4%
(a)
60
110
160
210
260
0 150 300 450 600 750 900
according to Ashby and Waterman
Temperature (°C)
Mod
ulus
of E
last
icity
(GPa
)
experimental resultswith error bars of ±5%
(b)
Appendices
216
Fig. 6—Carbon steel 1020 experimental results for the temperature
dependence of (a) the linear coefficient of thermal expansion compared to
literature data according to Smithells [17]; and (b) the tensile modulus of
elasticity compared to literature data according to Ashby and Waterman [20]
4 Selected Results
Tensile test were undertaken for aluminium alloy 7000 T4 to determine the
stress-strain curves at temperatures of 260°C and 480°C (Fig. 7). A constant strain
rate of 10-3s-1 was used throughout all tests. The lower temperature specimen
underwent a moderately ductile failure with regions of both strain hardening and
necking before the final sudden fracture. An elastic modulus of 63 GPa, 0.2% yield
strength of 180 MPa, ultimate strength of 233 MPa and a fracture strain of 27% were
determined. The higher temperature specimen experienced a highly ductile failure
with no sudden fracture. An elastic modulus of 37 GPa, 0.2% yield strength of 68
MPa, ultimate strength of 73 MPa and a fracture strain of 44% were determined.
Appendices
217
0
50
100
150
200
250
Engineering Strain (m/m)
Engi
neer
ing
Stre
ss (M
Pa)
480 °C
260°C
0.0 0.1 0.2 0.3 0.4 0.5
Fig. 7—Experimental results for the tensile stress-strain curves obtained at
temperatures of 260°C and 480°C for aluminum alloy 7000 T4
5 High Temperature Constitutive Models
The results of thermal expansion coefficient (CTE), modulus of elasticity and
stress-strain behaviour as functions of temperature were used to develop high
temperature constitutive models, in a form that can be implemented in finite element
analysis (FEA). At a given temperature, the constitutive equation has the form
( ) ( ) mTET εσ '= (1)
where σ is the stress, ε is the strain (calculated from the temperature dependent
CTE values), E’ is the modulus of elasticity, m is the strain exponent respectively
and T is the temperature. The constitutive models developed in this study were
Appendices
218
utilised in the thermal-structural analysis of the HyCause Scramjet engine for flight
test [21].
6 Conclusion
Various test methods and equipment for high temperature thermo-mechanical
testing have been reviewed. Based on these findings; the low cost high temperature
test rig designed and built for these investigations was then described. An induction
heater provides fast heating and cooling rates, simple operation and access to the test
specimen for temperature and strain measurement. Specially designed high
temperature specimen grips with water-cooling allow for testing over long periods of
time. Thermocouples measure temperature and provide a feedback signal for PID
control of the induction heater. Strain is measured by a mechanical extensometer
with ceramic extension rods. A microcontroller and laptop computer allow for
complete control and monitoring of all equipment and test parameters.
Calibration of the thermocouples made temperature measurement of the test
specimen possible to within ±4%. A validation study of the linear coefficient of
thermal expansion and tensile Young’s modulus of carbon steel 1020 verified the test
rig. Results for the thermal expansion coefficient and elastic modulus were obtained
within ±4% and ±5% of literature data, respectively. Tensile tests to failure of the
aerospace material aluminium alloy 7000 T4 further verified the reliability of the test
set-up.
Limitations were found with the use of contact strain measurement due to the
small size of the test specimens. At the higher temperatures contact forces from the
mechanical extensometer were sufficient to buckle the test specimens. With further
expense improvements could be made with the test rig, such as the use of non-
contact temperature and strain measurement. The capability could also be extended
to measure Poisson’s ratio.
The test rig developed for these investigations allowed for accurate and reliable
experimental results. Test equipment was chosen based on the ability to provide
Appendices
219
accurate results as well as low cost and simplicity in design and operation. The
ability for the test rig to be used for a range of thermo-mechanical tests was also an
important factor. The measured thermo-mechanical properties have enabled the
development of high temperature constitutive models for structural design
calculations and the thermal-structural analysis of a hypersonic vehicle for flight test.
Acknowledgements
This work was completed with the financial support of the Defence Science and
Technology Organisation (DSTO) of Australia.
Appendices
220
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Appendices
222
Appendix C
Porosity Measurements
Figures were taken from ASTM E2109-01 standard.
Figure 1: 0.5 % Porosity
Figure 2: 1.0 % Porosity
a1172507Text Box NOTE: Appendix C is included in the print copy of the thesis held in the University of Adelaide Library.
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TITLE: SLURRY SPRAYED THERMAL BARRIER COATINGS FOR AEROSPACE APPLICATIONSAPPENDICESAppendix AAppendix BAppendix C
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