# Thermal Barrier Coatings in burner rig experiment analyzed through LAser
Shock for DAmage Monitoring (LASDAM) method

Lara Mahfouz  Vincent Maurel  Vincent Guipont  Basile Marchand  Rami El
Hourany  Florent Coudon  Daniel E. Mack  Robert Vaßen  MINES Paris - PSL , MAT
- Centre des Matériaux, CNRS UMR 7633, BP 87 91003 Evry, France  Safran Tech,
Etablissement Paris Saclay, rue des Jeunes Bois-Châteaufort, 78772 Magny-les-
Hameaux, France  Forschungszentrum Jülich GmbH, Institute of Energy and
Climate Research (IEK-1), Jülich, Germany

###### Abstract

This study investigates failure mechanisms in a typical thermal barrier
coating (TBC) system comprising an EB-PVD columnar top coat, an aluminide bond
coat, and a Ni-based single crystal superalloy substrate, simulating gas
turbine operating conditions using a burner rig. TBC degradation, initiated by
interfacial defects from the LASAT method, was studied during thermal gradient
cycling under fast and slow cooling. In-situ optical and infrared imaging,
along with ex-situ SEM cross-sectional analysis, monitored failure mechanisms.
The Laser Shock for Damage Monitoring (LASDAM) technique provided insights
into gradient and cooling rate impacts on columnar TBC damage. Results showed
significant effects of cooling rate on delamination and localized failure at
blister sites, with LASDAM revealing significant overheating at damage sites.
Analysis included full-field temperature and damage assessment, emphasizing
blister-driven delamination under severe thermal gradients. Discussion focused
on elastic stored energy effects, noting that fast cooling induced transient
conditions where reversed temperature gradients increased damage, limiting TBC
lifespan.

###### keywords:

TBC , thermal gradient cycling , LASAT , blister driven delamination

††journal: JECS

##  1 Introduction

A standard Thermal Barrier Coating (TBC) system comprises an underlying bond
coat (BC), typically (Ni,Pt)Al, and an outer ceramic top coat (TC). In many
industrial applications, yttria-stabilized zirconia (YSZ) is the preferred
material for the top coat [1, 2]. When in service at elevated temperatures, a
thermally grown oxide (TGO), primarily alumina, naturally forms at the TC/BC
interface. The interfacial stress arising from the thermal expansion
coefficient (CTE) mismatch and the thickening of TGO during high-temperature
exposure are widely acknowledged as failure mechanisms for EB-PVD thermal
barrier coatings [1]. Previous studies have demonstrated that thermal cycling,
coupled with growth strain from oxidation, induces rumpling and micro-cracking
at the TC/BC interface [3, 4, 5]. It has been reported that the typical
failure pattern for TBCs involves interfacial crack initiation at the TC/BC
interface, progressing to macroscopic delamination. This is followed by the
formation of a blister through the buckling of the ceramic top coat layer,
ultimately leading to spallation failure of the TC [6, 1, 7].

These findings underscore the necessity for robust experiments to investigate
blister-driven delamination thoroughly [8]. In line with this perspective, the
LAser Shock for DAmage Monitoring (LASDAM) technique introduces new avenues
for analyzing the driving forces behind delamination from an initial blister
subjected to thermal cycling in a furnace [9]. LASDAM involves inducing an
artificial decohesion at the weak interface of the Thermal Barrier Coating
(TBC) and monitoring its evolution during cycling. The former is achieved
through precise control of the decohesion size and location using the LAser
Shock Adhesion Testing (LASAT) technique [10]. The latter entails combining
in-situ high-resolution CCD images, infrared thermography (IRT), and the
measurement of 3D blister morphology. When applying the LASDAM method with
thermal cycling in a furnace, it was observed that a significant initial
increase in blister buckling precedes further delamination [11]. The driving
forces for buckling were also identified as driving forces for this blister-
driven delamination, a major intrinsic failure mode in TBCs [9].

The aforementioned failure mechanisms are primarily derived from assessments
done with a uniform temperature field. However, real service conditions
involve non-homogeneous temperature fields, where thermal gradients lead to
different stress distributions in the TBC, resulting in varied damage
behaviors compared to uniform temperature fields [12, 13]. Burner rig cycling
enhances the representativeness of thermal cycling testing. In contrast to
traditional furnace cycling, burner rig testing enables the introduction of a
thermal gradient throughout the thermal barrier system, along with a surface
temperature gradient. This closely simulates service conditions, where
combustion gases, cooling channels, and the low thermal conductivity of the
ceramic induce a significant gradient of about
$100\text{\,}\mathrm{\SIUnitSymbolCelsius}$ through the coating thickness [14,
15].

Most burner rig tests for YSZ TBCs have been performed on air plasma spray
(APS) top coats and show as a notable result that the temperature gradient
associated with the low conductivity of APS leads to partial spallation of the
top coat [16]. However, the oxidation prior to the burner rig test shows a
transition from cracking within the TC layer to interfacial cracking,
depending on the burner rig cycling conditions and YSZ thickness, promoting
interfacial decohesion [17]. In addition, at elevated temperatures within a
temperature gradient, some authors have reported sintering of the surface of
the top layer, increasing both its conductivity and local Young’s modulus
[18]. It is important to note that some studies dealing with APS TC assume
that the presence of interfacial cracks does not modify the temperature
gradient across the TC [19], while others have demonstrated that a small
temperature difference between crack faces results in a significant change in
the mechanical state of the TBC [20, 13].

For EB-PVD YSZ top coats, damage is initiated by interfacial decohesion under
both homogeneous temperature and burner rig conditions [21]. This same study
indicates that thermal gradient cycling requires a greater number of cycles to
induce failure compared to uniform temperature cycling, given a similar
Thermal Growth Oxide (TGO) or interfacial temperature. The lower this
temperature, the more pronounced the difference in experimental lifetime.
However, these findings are influenced by substantial variations in cooling
rates that could impact subsequent lifetime, coupled with edge effects driving
Thermal Barrier Coating (TC) failure [12].

In addition to through-thickness temperature gradients, the surface
temperature field on the turbine blade exhibits significant variability,
creating surface thermal gradients that have been largely overlooked in the
existing literature. Numerical studies, for instance, reveal temperature
fluctuations exceeding 200 K over a distance of less than 1 mm, particularly
near cooling holes [22]. Therefore, enhancing our comprehension of thermal
barrier system behavior necessitates considering the effects of both through-
thickness and in-plane thermal gradients.

This study aims to elucidate the local effects of through-thickness and in-
plane temperature gradients on a typical Thermal Barrier Coating (TBC) system
with a columnar EB-PVD top coat. For the first time, the LASDAM method is
introduced to the burner rig setup and plays a central role in the analysis.
The paper begins with a detailed description of the burner rig configuration
and the LASDAM test conditions. The burner rig setup allows for the adjustment
of the cooling rate. The analysis of local temperature, microstructure
evolution and damage rate for two specified cooling conditions is then
discussed. The results are compared to homogeneous temperature furnace cycling
tests, and the subsequent discussion attempts to elucidate the extent to which
local temperature and temperature gradient influence interfacial damage rates.

##  2 Experimental methods

###  2.1 Specimen

The studied TBC system is yttria (Y2O3) partially stabilized zirconia (ZrO2),
YPSZ, ceramic EB-PVD top coat and (Ni,Pt)Al bond coat processed in the vapor
phase with high temperature low activity (APVS) treatment, deposited on a
first generation Ni base single crystal superalloy AM1. Disk-shaped specimens,
typical for burner rig tests, are employed in this study [23, 24]. The
superalloy substrate has a thickness of 3 mm and a diameter of 23 mm. In the
as-processed state, the bond coat is approximately 50 µm thick, and the top
coat is 145 ± 5 µm thick. To reduce stress levels and prevent failure at sharp
edges, a 1.5 mm radius of curvature is machined at the outer edge of the
samples. Additionally, a 1.2 mm hole is radially drilled in the substrate at
mid-thickness up to the center of the sample, housing a thermocouple for
measuring substrate temperature during burner rig cycling. A notch on the
sample rim facilitates mounting in the sample holder of the burner rig test
facility. Refer to Figure 1 for a detailed view of the geometry. This study
focuses on testing a single batch of coating.

Figure 1: Samples geometry (dimensions are given in mm)

###  2.2 Designing Interfacial Decohesion Using LAser Shock Adhesion Testing
(LASAT)

Interfacial cracks are processed within the coated samples by a laser shock
using the LASAT method on a Saga Thales facility. This involves employing a
Nd-YAG laser pulse with a wavelength
$\lambda$=$532\text{\,}\mathrm{n}\mathrm{m}$ and a maximum energy of 2 J,
lasting 5.2 ns. The LASAT process is preceded by 100 thermal cycles from 100
to $1100\text{\,}\mathrm{\SIUnitSymbolCelsius}$, each with a 50-minute dwell
time at high temperature in a cyclic oxidation furnace at Safran [9]. This
pre-aging, consistent with [9], aims to initiate localized decohesion at the
TC/TGO interface after the laser shock.

The laser shock diameter on the substrate surface is adjusted using a
converging lens, and confinement is achieved by covering the substrate side
with transparent adhesive tape. The laser power density ranges from 0.1 to 10
GW/cm², with laser diameters typically between 1 and
$4\text{\,}\mathrm{m}\mathrm{m}$; in this study, a 3 mm laser diameter was
consistently used.

To measure the initial size of the debonded area post-LASAT processing, the
sample is placed on a heating plate at 60 °C, and the surface temperature is
measured using infrared thermography (IRT) (see Figure 2(a)). The IRT camera,
an INFRATEC ImageIR® 8300 hp with zoom capability, provides an in-plane
spatial resolution of approximately $42\text{\,}\mathrm{\SIUnitSymbolMicro
m}$.

A preliminary assessment of the adhesion level of the TBC was conducted on a
reference sample (ref.) by creating a ”LASAT curve,” illustrating the
evolution of the measured crack diameter in relation to the laser power
density. The laser power density was systematically increased by gradually
raising the laser energy for a constant laser spot diameter (3 mm in this
study). The reported LASAT diameter represents the equivalent diameter of the
debonded area measured by IRT, as depicted in Figure 2(b). The LASAT curve
(ref. in 2(b)) provides insights into the debonding laser power density
threshold, serving as a reference to determine the laser energy level for
inducing defects in the test samples.

Moreover, optical observations allow for the examination of whether
delamination leads to buckling of the TC and if the resulting blister exhibits
cracking or experiences local/total spallation. Notably, for higher laser
energies, a characteristic LASAT transition from a ”perfect blister” (without
cracking or spallation) to a cracked/spalled state has been identified for a
given TBC. This transition can be discerned by plotting a LASAT curve [25].
The pointed arrow in Figure 2(b) emphasizes this transition, indicating the
onset of spallation, with full symbols denoting perfect blisters and empty
ones signifying local or total spallation. The SEM cross section of a blister
from the reference sample used to plot the ”LASAT curve” is shown in figure
2(c). It shows significant oxidation, oxide thickness locally reaching
$2\text{\,}\mathrm{\SIUnitSymbolMicro m}$, and phase transformation from
$\beta$ to $\gamma^{\prime}$ phase, along with decohesion occurring at the
interface between the oxide layer and the top coat layer, consistently with
previous results [11]. The localization of the decohesion induced by LASAT
confirms that the pre-aging weakens the oxide/top coat interface properties,
consistent with the state-of-the-art for such TBC systems.

(a) IRT after LASAT (sample A)

(b) LASAT curve

(c) SEM cross-section (Ref. sample)

Figure 2: LASAT method: (a) Infrared thermography (IRT) employed for assessing
the debonded diameter after laser shock (b) Equivalent debonded diameters
derived from IRT, with the arrow marking the initial occurrence of local
spallation (full symbols denote perfect blisters, while empty symbols
represent blisters with local or total spallation) (c) SEM cross section of a
blister from the reference sample after LASAT test

.

We chose to introduce three defects on each burner rig test specimen,
strategically avoiding critical areas that might interfere with the LASAT
test, specifically, above the drilled hole and at the laser-engraved specimen
reference. The application points for laser shocks were deliberately spaced
apart to prevent potential interaction or premature coalescence. Laser shocks
were applied at a target laser energy of 900 mJ, corresponding to a laser
power density of approximately 2.2 GW/cm2. The decohesion areas achieved on
one of the samples (A) are shown in Figure 2(a). Equivalent diameters for all
blisters introduced on test samples are plotted on the reference LASAT curve,
as shown in Figure 2(b). While there is some variability in the resulting
equivalent diameter among the samples, it can be considered negligible within
the same sample. This variability is attributed to potential variations in TBC
adhesion, even though the test samples are selected from a single batch. It is
noteworthy that the LASAT technique has demonstrated high sensitivity to
initial adhesion [26, 27].

Table 1: Details of specimen tested, $\Phi$ is the applied shock energy, D0 (mm) is the initial debonded diameter, $\dot{T}$ is the average cooling rate, and Nd and Ns correspond respectively to the number of cycles for debonding and spallation when observed in situ. For the reference specimen used for LASAT curve, only ranges are indicated. sample | Spot |  $\Phi$ (GW/cm2) |  D0 (mm) |  $\dot{T}$ (.s-1) |  Nd |  Ns
---|---|---|---|---|---|---
Ref. | 9 | 1-3.3 | 0-4 | - | - | -
A | 1 | 2.20 | 2.60 | 30 | - | -
2 | 2.18 | 2.99 | 30 | 450 | 500
3 | 2.19 | 3.16 | 30 | 250 | 300
B | 1 | 2.23 | 2.90 | 30 | - | -
2 | 2.21 | 3.02 | 30 | 490 | 500
3 | 2.19 | 3.17 | 30 | 290 | 300
C | 1 | 2.23 | 3.61 | 9 | 90 | 130
2 | 2.21 | 3.53 | 9 | - | 300
3 | 2.19 | 3.84 | 9 | 10 | 90

###  2.3 Burner Rig setup

The burner rig test facility utilized in this study is installed at IEK1
Julich. The set-up is detailed in Figure 3. The burner operates on a mixture
of natural gas and O2. During the heating phase, the burner flame is
positioned over the ceramic surface, while the back side of the sample is
cooled by compressed air to establish the desired temperature gradient across
the thermal barrier coating. The temperature on the cold side of the sample is
monitored using a thermocouple inserted into the substrate, while the
temperature on the gas side of the sample is measured with a long-wave KT15.99
IIP (Heitronics®) pyrometer (9.6-11.5 $\mu$m). The recorded temperature value
represents an average over a spot size of 10 mm in diameter focused on the
center of the sample.

Figure 3: Buner rig test set-up.

In this study, the temperature was set to
$1200\text{\,}\mathrm{\SIUnitSymbolCelsius}$ on the ceramic surface and
$1050\text{\,}\mathrm{\SIUnitSymbolCelsius}$ in the substrate. The coating’s
emissivity ($-eps-converted-to.pdfilon$) was assumed to be constant at 1.
While this value is considered to underestimate the surface temperature by
approximately 3%, realistic emissivity values for YSZ ceramic in the
pyrometer’s wavelength range are around 0.98. However, emissions from the
flame within the same spectral range partially counteract the impact of an
overestimated emissivity [28]. Furthermore, it’s important to note that the
emissivity at the blisters might vary. This variability introduces uncertainty
into the measurements, which is later addressed and considered in the
analysis.

Two cooling rates were examined. Fast cooling was achieved by moving the
burner away immediately after the high-temperature dwell phase, while slow
cooling was attained by maintaining the flame over the ceramic surface during
cooling and reducing the gas flow rate. Details of the tested cycles will be
provided in Section 3.1.

A FLIR A655sc long-wave infrared (LWIR) infrared thermal camera with a
resolution of 640 x 480 pixels and a spectral range of 7.5 - 14.0 $\mu$m
provides in situ acquisition of the temperature field distribution on the
surface of the ceramic, with a lateral resolution of about 63 $\mu$m.
Simultaneously, an optical camera with a CMOS sensor (1400x1024 pixels)
provides continuous high-resolution in situ observation of the coating
surface. Both cameras are housed in cooled chambers during testing.

Intermittent pauses in cycling were implemented to use a CyberTECHNOLOGIES
profilometer for measuring the 3D evolution of blisters. The profilometer has
a lateral resolution of 3.5 $\mu$m and a vertical resolution of 35 nm. Cycling
tests were stopped upon the observation of macroscopic restricted blister
cracking, and samples were handled carefully to prevent additional damage to
the top coat for subsequent analyses.

This study conducts a detailed examination of three distinct samples. Since
three LASAT blisters are processed in each, a total of nine different blisters
are tested using the proposed LASDAM method. The next section focuses on the
thermal loading throughout burner rig cycling. Damage evolution in presence of
the processed blister is later analyzed.

##  3 Thermal loading experienced by blisters in burner rig cycling

###  3.1 Prescribed thermal cycles

The burner rig cycles with slow and fast cooling rates are denoted as BR-slow
and BR-fast, respectively. Figure 4 illustrates the temperature evolution over
one cycle for both conditions. The surface temperature is obtained from the
pyrometer measurement, while the substrate temperature is obtained from the
thermocouple at mid-thickness in the substrate (see Figure 4(a)). Similar
heating rates are observed in both conditions. The BR-fast cycle exhibits an
average cooling rate of
$30\text{\,}\mathrm{\SIUnitSymbolCelsius}\mathrm{/}\mathrm{s}$ from the
maximum substrate temperature to 200 , with a notably faster initial cooling
rate. On the other hand, the BR-slow cycle has an average slope of 9 /s. The
through-thickness temperature gradient is assessed as the temperature
difference between the surface temperature measured with the pyrometer and the
substrate temperature measured with the thermocouple at mid-thickness (see
Figure 4(b)). Notably, the fast cooling rate cycle shows a reversal in the
sign of temperature difference during the first few seconds of cooling, while
the slow cooling rate cycle exhibits a slight increase in temperature
difference during the initial seconds of cooling.

(a) Temperature cycles at the ceramic surface ($T_{\text{surf}}$), represented
by full lines, and in the substrate ($T_{\text{sub}}$), represented by dashed
lines.

(b) Through thickness temperature gradient

Figure 4: Temperature profiles recorded with a pyrometer on the ceramic
surface and a thermocouple in the substrate during both fast and slow cooling
rate cycles.

###  3.2 Dwell at maximum temperature

During the dwell time at maximum temperature, infrared thermography (IRT)
captures the temperature distribution on the surface of the top coat, as shown
in figures 5(a) and (b). These observations are made after a few stabilization
cycles for both samples A and C, tested with BR-slow and BR-fast,
respectively. Throughout the sample, a central hot spot can be seen, due to
the direct exposure of the surface of the top layer to the flame, with an off-
center maximum temperature observed in the upper left corner of the infrared
images. Additionally, local maximum temperatures are observed at blister
locations.

(a) BR-slow (sample A)

(b) BR-fast (sample C)

(c) BR-slow (sample A)

(d) BR-fast (sample C)

Figure 5:  Temperature distribution captured by infrared thermography on the
top coat surface during the dwell time at maximum temperature.

Temperature profiles are generated along the black lines intersecting the
blisters centers in Figures 5(a) and 5(b) for BR-slow and BR-fast conditions,
respectively. The flame-associated temperature gradient is approximately 10 K,
measured from 20 to 80 pixels on the temperature profile intersecting blister
A-2 in Figure 5(c). Across blister A-2, the gradient is around 100 K, measured
from 80 to 140 pixels (blue curve in Figure 5(c)). In particular, a
significant overheating of about 60 K is observed at the blister sites,
extending from the centers to the edges of the blisters, as shown in Figures
5(c) and 5(d).

Regardless of the maximum temperature recorded at the centers of the blisters,
similar local temperatures are experienced at the edges of the blisters and
their associated crack tips for both BR-fast and BR-slow conditions,
approximately ranging from 1215 to
$1220\text{\,}\mathrm{\SIUnitSymbolCelsius}$ at blister location #3. At
blister location #2, corresponding to a lower temperature, both the local
temperature and temperature gradient exhibit remarkable similarity for both
conditions. These consistent observations affirm the robustness of the chosen
experimental setup.

The observed overheating linked to each blister is likely a result of a layer
of air beneath the blister, limiting the conduction of heat to the substrate.
The maximum temperature observed in the center of a blister corresponds to the
location of the maximum air gap thickness, while the gradual temperature
decrease toward the edges of a blister corresponds to a continuous reduction
in the air gap thickness, which raises questions about the influence of the
blister’s local height on overheating.

###  3.3 Temperature evolution of blisters during cycling

Stable heating conditions are observed during cycling; nonetheless, there is a
notable evolution in blisters temperature. In the case of BR-slow cycling, the
maximum temperature at each of the three blisters exhibits a two-stage
evolution. Initially, a continuous temperature increase is observed (e.g., up
to N=280 for A-3 and N=470 for A-2, followed by a sudden temperature surge
over a few cycles, as illustrated in Figure 6(a). For A-1, only the first
stage of evolution is evident, characterized by a slight temperature increase
during cycling.

Similar two-step evolutions are observed for the BR-fast cycling condition.
However, in the case of C-3, a drastic temperature increase occurs in fewer
than 10 cycles.

Assuming that local overheating results from the presence of the air gap
induced by blister buckling, the evolution of maximum temperatures indicates
that the BR-fast cycling condition is more detrimental than the BR-slow
cycling condition. Blisters C-1 and C-3 are exposed to higher local
temperatures than C-2 and experience a more rapid instability, a topic to be
discussed in the 4.1 section.

(a) BR-slow (A)

(b) BR-fast (C)

Figure 6: Evolution of local maximum temperatures at different blister
locations over cycling.

###  3.4 A focus on cooling condition

While the temperature range for IRT calibration applies to temperatures above
$400\text{\,}\mathrm{\SIUnitSymbolCelsius}$, our focus is on cooling down to
$400\text{\,}\mathrm{\SIUnitSymbolCelsius}$. Under the BR-slow cycling
condition, achieved by decreasing the burner gas flow rate, Figure 7(a)
illustrates that within less than 15 seconds, blister temperatures gradually
align with the temperature measured at the adherent reference point. The
temperature decreases from $1530\text{\,}\mathrm{\SIUnitSymbolCelsius}$ to
$800\text{\,}\mathrm{\SIUnitSymbolCelsius}$, considering the maximum
temperature measured at A-2 after 490 cycles.

Conversely, for the BR-fast cycling condition, a similar temperature decrease
is observed at cycle 10 in less than 1.5 seconds, as shown in Figure 7(b).
Notably, the blister temperature drops below the temperature at the adherent
reference point. This occurrence may be attributed to the sustained insulation
effect of the air gap during cooling. The increased cooling rate, achieved by
swiftly moving the burner away from the specimen can result in increased heat
loss from the blister through direct radiation, surpassing the heat loss
through conduction to the substrate in the adherent region. In the BR-slow
condition, this effect is mitigated as the burner continues to heat the sample
during the cooling phase, albeit at a reduced gas flow rate.

(a) BR-slow (A) at N=500

(b) BR-fast (C) at N = 90

Figure 7: Temperature profiles during cooling captured by infrared imaging at
the blisters centers and a 10 mm circular reference area in the sample
centers.

##  4 Damage analysis

###  4.1 Evolution of blister morphology

In-situ imaging with the CMOS camera allows for a detailed analysis of the
blister surface evolution throughout cycling. For the BR-slow condition
(sample A), blister A-3, exposed to the maximum temperature, exhibits
significant blistering after 200 cycles and develops cracks after 300 cycles
(see Figure 8). In contrast, blister A-2, located in the lowest temperature
region, displays substantial blistering after 450 cycles, and milder cracking
becomes evident towards the end of the 500-cycle test. Blister A-1 remains
more stable than the other two blisters. Cycling is stopped immediately upon
the observation of cracking at blister A-2. Whether the damage observed at
blister A-1 is related to initial lower adhesion at this location or to a
higher in-plane temperature gradient requires further clarification by a more
detailed analysis of the in-plane temperature gradient, possibly with a higher
number of induced blisters within the sample area.

Figure 8: In situ CMOS analysis of samples subjected to BR-slow cooling rate
(Sample A)

For the BR-slow condition, both tests conducted on samples A and B were
stopped after 500 cycles, revealing a strikingly similar trend. The only
exception was B-3, exhibiting spallation rather than severe cracking as
observed for A-3 (Figure 9). Clearly, the evolution of blisters is influenced
by their specific location and associated maximum temperature. Notably,
blister damage, whether in the form of cracking or spallation, is most
significant at the location corresponding to blister #3’s maximum temperature.
However, buckling and delamination are more pronounced at location #2 compared
to location #1 on both samples A and B. This suggests that earlier cracking at
site #3 is associated with a maximum surface gradient (Figure 6(a)).

(a) Sample A

(b) Sample B

Figure 9: In situ CMOS analysis of samples tested with the BR-slow cooling
rate (samples A and B).

In the BR-fast condition (sample C), blister C-3, exposed to the highest
temperature, exhibits significant blistering after just 10 cycles, followed by
cracking at 90 cycles and eventual spallation at 320 cycles (see Figure 10).
Meanwhile, the other two blisters display less blistering and no spallation at
N=320. In particular, C-1, which is at a higher temperature, exhibits more
pronounced blistering than C-2. In addition, C-1 exhibits surface roughness,
which may be associated with localized cracking. The evolution of the blisters
follows the observed temperature increase over the cycles, as shown in Figure
6(b). This figure highlights that the cooling rate determines the damage rate,
with increased damage correlating with the maximum overheating of the blister.

Figure 10: In situ CMOS captures at different number of cycles of the ceramic
surface of a sample tested with BR-fast cooling rate (sample C).

###  4.2 Microstructure evolution

#### BR-slow

After testing, a higher-resolution ex situ Infrared Thermography (IRT) control
of specimen B is conducted, as depicted in Figure 11(a). The results affirm
that the locations of damage align consistently with the maximum local
temperatures at the blisters, and no other macroscopic defects are
discernible. It is noteworthy that spallation on blister B-2, though not
observed in situ, is evidently induced by a desktop spallation effect, as can
be concluded from the comparison between Figures 11(a) and 9(b). A top-view
scanning electron microscope (SEM) image in Figure 11(c) reveals pronounced
sintering of the top coat, highlighting the internal pores’ aspect, and the
absence of the typical feather-like structure.

Figure 11: Ex situ analysis of a sample tested with BR-slow cooling rate
(sample B): (a) IRT observation of the surface of the specimen, (b) SEM
surface observation of location B-2 orange dashed line corresponds to initial
location of debonding processed by LASAT, orange arrows indicate further
decohesion under burner rig cycling, (c) SEM surface observation of location
B-3 and (d) SEM cross-section corresponding to the black dot lines in (a) and
(b) at the location B-2.

The global surface view shown in Figure 11(b), corresponding to blister B-2,
shows the initial debonding site processed by LASAT, highlighted by the dashed
line. The maximum observed rumpling is confined within a circular region,
clearly showing the onset of debonding. Beyond this region, one can observe
the area where the interfacial crack has grown from this initial debonding
site in the direction of the arrows. In this area of further debonding, few
local undulations are observed, consistent with established findings that
rumpling manifests exclusively at points of local decohesion [9]. Similarly,
cross-sectional analysis of blister B-2, as shown in Figure 11(d), confirms
that the peak of rumpling occurs beneath the blister and is significantly
reduced in regions where the topcoat adheres to the substrate. In addition,
the adherent portion of the topcoat layer shows no signs of sintering,
consistent with the lower surface temperature measured in this particular
area, see Figure 5(c).

#### BR-fast

Examination conducted ex situ verifies that, under the BR-fast condition,
blisters C-1 and C-2 remain unspalled. However, infrared thermography (IRT)
reveals the presence of small cracks on blister C-1, as illustrated in Figure
12(a). Meanwhile, spallation is complete for blister C-3, with no residual
ceramic observed on the TGO surface. Additionally, sintering impacts the
columnar structure of the top coat in the vicinity of blister C-3, as shown in
Figure 12(c). Notably, a feather-like structure persists at the coolest point,
adjacent to adherent area C-4, where internal porosity is visible, as shown in
Figure 12(b). This porous structure is less prominent at other locations
exposed to higher temperatures, illustrated in Figure 12(b) for blister C-2.

Figure 12: Ex situ analysis of a sample tested with BR-fast cooling rate
(sample C): (a) IRT observation of the surface of the specimen, (b) SEM
surface observation of location B-2 orange dashed line corresponds to initial
location of debonding processed by LASAT, orange arrows indicate further
decohesion under burner rig cycling, (c) SEM surface observation of location
B-3 and (d) SEM cross-section corresponding to the black dot lines in (a) and
(b) at the location B-2.

The cross-sectional analysis along the line connecting blisters C-1 and C-2 in
Figure 12(a) reveals a noteworthy correlation between local high temperature
during cycling and the sintering process, as depicted in Figure 12(d). At the
center of blister C-1, a highly dense microstructure is evident, indicative of
localized extremely high temperatures. Importantly, the sintering is more
pronounced in the lower part of the top coat (TC) throughout its thickness,
consistent with local compressive stresses induced by the bending of the
ceramic layer, but also corresponding to an initial denser part of the ceramic
layer.

Remarkably, at the location of the crack tip, and consistent with the lower
local temperature, no discernible evidence of sintering is observed, as
illustrated in Figure 12(d). This observation underscores that, despite the
elevated local temperature resulting from the blister-induced air gap, the
crack propagates in an area characterized by lower and more uniform
temperatures within the specimen’s region of interest. In essence, the hot
points induced by the air gap solely influence the temperature of the
blisters, and blister driven delamination is analyzed at the targeted
temperature.

More precise quantitative results are derived through cross-sectional image
analysis. The average porosity is computed across a moving rectangular area,
with a thickness matching that of the top layer, along the radial direction
starting from the crack tip. The lateral dimensions of the rectangle are set
to $21\text{\,}\mathrm{\SIUnitSymbolMicro m}$. To enhance the clarity of the
obtained measurements, a moving average method is applied over 10 rectangles.

It is evident that for blisters C-1 and B-3, the sintering is more pronounced
in the center (resulting in lower porosity) and gradually decreases toward the
crack tip (accompanied by an increase in porosity). In the case of sample C,
where less time is spent at high temperatures compared to sample B, the
porosity is expected to be inherently higher, as shown in Figure 13(a). This
is true for C-1 and C-2 compared to B-3. However, the measured porosity for
B-2 is higher, possibly due to the slightly off-center cross section caused by
the partial spallation of the blister, as shown in Figure 11(d). It is
noteworthy that the local temperature for B-2 may be lower than for C-2.
Finally, it is important to note that given the inherent variability observed
in the EB-PVD deposition process, the initial porosity may vary from one
sample to another.

(a) Porosity variation as a function of the distance from blister, denoted as
r

(b) Oxide thickness hox

(c) $\gamma$’ phase fraction fs

Figure 13: Microstructure analysis considering (a) variation of porosity as a
function of the distance from the center of the blister, (b) oxide thickness
and (c) $\gamma$’ phase fraction evolutions for burner rig (BR/current study)
and furnace cycling (S-FCT/data from [29])

In addition to sintering of the top coat, SEM cross sections allow assessment
of both local oxide thickness and phase transformation in the bond coat layer,
from $\beta$-NiAl to $\gamma$’-Ni3Al, Figures 13(b) and (c), respectively.
Both were measured using SEM observations and segmentation through standard
thresholding techniques using ImageJ/Fiji software [30], for regions
approximately $1.4\text{\,}\mathrm{m}\mathrm{m}$ wide in adherent top coat
areas. The $\gamma^{\prime}$ phase fraction corresponds to the surface
fraction of the outer layer, defined as the area between the oxide layer and
the embedded grit alumina particles, see Figure 2(c).

Oxide thickness measurements were conducted for both BR-slow and BR-fast
conditions on samples B (N=500, equivalent time at high temperature: 37.5
hours) and C (N=320, equivalent time at high temperature: 24 hours). These
measurements were taken at two locations: the adherent part of the TC near the
point used for pyrometer surface control and at the crack tip positions at B-2
for BR-slow and C-1 for BR-fast. Notably, B-2 exhibits the highest oxide
thickness, possibly due to final overheating, while C-1 displays the lowest
oxide thickness, likely attributed to local oxide spallation.

However, upon comparing oxide thickness below the blister and in areas where
the TC is adherent, no significant difference in oxide thickness or phase
transformation is observed (this corresponds to measurements at the same time
but with different values). This observation suggests that, despite the
elevated surface temperature caused by the blister, the variation in
interfacial temperature can be considered small enough not to substantially
impact the oxidation and phase transformation phenomena. A comparison with
previous results obtained for S-FCT cycling will be discussed later.

In summary, the BR-fast condition induces spallation earlier than the BR-slow
condition. The selected setup for the burner rig generates a non-uniform
temperature field, and consequently the location of the blister with respect
to the burner flame plays a crucial role in the damage induced. Furthermore,
the presence of an air gap between the ceramic layer and the TGO surface
promotes an increase in thermal resistance, consequently elevating the blister
temperature and advancing sintering in the top coat.

###  4.3 Analysis of damage rate

As outlined earlier, the LASDAM technique enables the monitoring of both in
situ interfacial debonding and ex situ blistering [9]. The evolution of
decohesion diameter as a function of the number of cycles for BR-slow testing
(A and B samples) is presented in Figures 14(a) and 14(b). The progression of
blisters in locations 2 and 3 exhibits remarkable similarity for both
specimens, particularly concerning the number of cycles leading to the
initiation of debonding, which occurs around 260 cycles for both specimens:
The disparity between samples A and B is minimal, with a difference of no more
than 10 cycles for location 3, as illustrated in Figure 14. Additionally, at
N=500, no debonding is observed for blister #1 in both samples. The number of
debonding and spallation cycles for each blister is detailed in Table 1.

(a) Debonded diameter sample A

(b) Debonded diameter sample B

(c) Blisters height sample A

(d) Maximum temperature sample A

Figure 14:  Evolution of blisters under slow coooling rate: debonded diameter
for (a) sample A as a function of N and (b) sample B as a function of N, (c)
blister height for sample A as function of N and (d) maximum temperature at
blisters from sample A as a function of blister height. (full symbols denote
perfect blisters, while empty symbols represent blisters with local or total
spallation)

For sample A, profilometer measurements were conducted every 100 cycles,
necessitating testing interruptions. Given the analogous debonding behavior
observed in both samples, these interruptions are presumed to have a minimal
impact on the overall damage rate. The evolution of blister height over the
number of cycles for sample A is shown in Figure 14(c). The results indicate
that, while the equivalent delaminated diameter initially remains relatively
constant, the blister height undergoes a significant increase. This trend
aligns with observations made during furnace thermal cycling, where blister
evolution is initially characterized by the activation of a buckling mechanism
until a point is reached where further growth of interfacial debonding is
observed [11]. The height evolution presented in Figure 14(c) also illustrates
a threshold, approximately $50\text{\,}\mathrm{\SIUnitSymbolMicro m}$, beyond
which spallation occurs. Additionally, Figure 14(d) demonstrates a consistent
increase in maximum temperature, plotted previously in Figure 6, with
increasing blister height, supporting the importance of analyzing local
overheating in understanding blister evolution.

As the crack tip temperatures associated with each blister location remain
similar during the dwell at high temperature for both BR-slow and BR-fast
conditions, the two loading conditions can be considered comparable in terms
of interfacial oxidation. To eliminate the influence of initial blister
dimensions, the relative debonding and relative blistering are used for
comparison. These are described, respectively, as:

$\frac{D(N)-D(0)}{D(0)}$ (1)

and

$\frac{\delta(N)-\delta(0)}{\delta(0)},$ (2)

where $D(N)$ and $D(0)$ represent the current and initial debonded diameters,
and $\delta(N)$ and $\delta(0)$ represent the current and initial blister
height respectively.

Significantly accelerated macroscopic evolution kinetics are evident for both
delamination and blistering in BR fast cycling compared to BR slow cycling, as
shown in Figures 15(a) and (b), respectively. To clarify the role of thermal
gradient on damage rate, S-FCT data were added to these plots, combining data
from four different samples. For three of these samples, three individual
blisters are examined, while one other sample contains only one blister. To
emphasize the blister condition, full symbols indicate blisters with no
evidence of topcoat damage, while empty symbols indicate local cracking and/or
spalling of the topcoat at the blister location. In particular, local cracking
or spallation manifests much earlier at the fast cooling rate, with the first
local crack observed after only 10 thermal cycles and the second after 90
cycles, see Figure 15 and Table 1 for values. In contrast, these events occur
at later stages, specifically after 300 and 500 cycles, respectively, for the
slow cooling rate.

(a) relative diameter

(b) relative height

Figure 15:  (a) Evolution of the relative debonded diameter with respect to
the number of cycles; (b) Evolution of the relative blister height with
respect to the number of cycles for both fast and slow cooling rates (full
symbols denote perfect blisters, while empty symbols represent blisters with
local or total spallation).

##  5 Discussion

###  5.1 Microstructure evolution for burner rig and FCT

A previous study on the same system has provided data on microstructure
evolution, particularly focusing on oxide thickness, denoted as hox, and phase
transformation of the outer bond coat layer from $\beta$-(Ni,Pt)Al to
$\gamma$’-Ni3Al phase [31, 29]. This was analyzed through the surface fraction
(fs) of phase $\gamma$’-Ni3Al. The selected FCT cycle for this study, as
described in [9], involves linear heating (and cooling) within 5 minutes, from
100 to $1100\text{\,}\mathrm{\SIUnitSymbolCelsius}$, with 5 minutes dwell at
$1100\text{\,}\mathrm{\SIUnitSymbolCelsius}$.

Considering a similar cumulative time spent at maximum temperature,
$t_{\textrm{HT}}$, the results of the BR experiment are very similar to those
of the FCT regarding oxide thickness, with the variation being relatively
small compared to the inherent uncertainty in their quantification. However,
the difference for phase transformation is more significant, suggesting a
local lower temperature in the bond coat for burner rig compared to FCT.
However, this point should be mitigated by the fact that this measurement is
highly sensitive to the width of the analyzed area, e.g. as shown in previous
results dealing with the same TBC system [32]. This suggests that the observed
increase in blister evolution, both in height and debonding, is mainly
controlled by a change in mechanical loading rather than a change in
interfacial oxidation. The only significant difference in the microstructure
evolution when comparing BR to FCT is the modification of the sintering at the
center of the blister. This suggests, a priori, only a mechanical effect.

A robust analysis of the temperature gradient is needed to more accurately
assess the temperature reached at the crack tip during the BR cycle, by the
way, the microstructure was observed only on the remaining adherent areas to
avoid artifacts induced by local oxide spallation. This will be addressed in a
future work.

###  5.2 Influence of temperature, temperature gradient and cooling rate on
damage

In a previous study on the TBC system investigated in this paper, LASDAM
results were obtained under homogeneous temperature conditions throughout the
system, using short Furnace Cycling Tests (FCT) with dwell times of 5 minutes
[9]. The same experimental methodology was then used: the interfacial defects
were induced after an initial aging of 100 long FCTs (L-FCT) with a dwell time
of 50 minutes at high temperature. Figures 15(a) and 15(b) illustrate that
both the damage rate associated with relative variations in debonded diameter
and the stress buildup associated with relative variations in blister height
are higher for burner rig (BR) cycling compared to S-FCT.

Another aspect of the analysis is to evaluate the height-diameter
relationship, which serves as an indicator of the stress state within the
blister. In this context, for a given diameter, a higher blister height
corresponds to increased stresses within the blister [33, 9]. In the as-
received (AR) condition, corresponding to the LASAT defect introduced after
100 preliminary L-FCTs, indicated by the black symbols in Figure 16, the
maximum height is approximately $40\text{\,}\mathrm{\SIUnitSymbolMicro m}$ for
a diameter of $4\text{\,}\mathrm{m}\mathrm{m}$. No noticeable difference in
height compared to the AR state is observed after 160 cycles of BR fast
cycling. However, for a higher number of cycles, a trend becomes apparent
indicating increased stress in the system for both BR-fast and BR-slow
conditions. The blister height obtained under BR conditions approaches that
achieved after 2015 S-FCT cycles, as indicated by the green triangle symbols
in Figure 16. This further demonstrates that BR testing is more detrimental
than S-FCT, as the stress build-up is higher in the former than in the latter,
in consideration of the lower number of cycles.

Figure 16: Relationship between blister height and debonded diameter for Short
Furnace Cycling Test (S-FCT) fatigue (results from [9]) and Burner Rig (BR)
experiment (current study, with maximum height measured before spallation)

.

On the other hand, considering the similar oxidation and phase transformation
rates in BR and FCT conditions, it could be assumed that the average bond coat
temperature at the crack tip was similar for both test conditions. This point
is of high interest because it has been shown in the LASDAM experiment that
the BC plasticity plays a key role in determining the applied torque to the
blister [9]: the higher the plasticity in the BC, the higher the blistering
and damage rate. Thus, the local behavior of the BC might be comparable.

###  5.3 Impact of temperature, temperature gradient and cooling rate on
energy

The previous analysis is entirely focused on experimental outcomes. It is
worth exploring whether state of the art approaches of mechanical loading
analysis would yield consistent conclusions. For sake of simplicity, we
initially describe the system’s energy without considering the blister but
incorporating the experimentally measured temperature gradient across adherent
coating.

The first selected model is proposed by Levi et al. in Ref. [12]. The authors
evaluate the elastic energy, $U$, stored in the top layer, assuming plane-
strain conditions. The interest of the model is its sensitivity to the
temperature gradient across the interface and the temperature drop on cooling,
for both the top coat and the substrate. The model follows:

$U(t)=\frac{E(t)h(1+\nu)}{2(1-\nu)}\Big{\\{}(\Delta\alpha\Delta
T_{sub})^{2}-(\Delta\alpha\Delta T_{sub})(\alpha\Delta
T_{sur/sub})+\frac{1}{3}(\alpha\Delta T_{sur/sub})^{2}\Big{\\}}$ (3)

where E, $\nu$, and $\alpha$ are the Young’s modulus, Poisson’s coefficient,
and CTE of the TC layer, respectively. The thickness of the TC is denoted $h$.
The CTE mismatch between the substrate and the TC is denoted $\Delta\alpha$.
The temperature gradient $\Delta T_{sur/sub}$ is defined as the difference
between the temperature of the surface and that of the substrate, and the
cooling $\Delta T_{sub}$ as the substrate temperature drop from the maximum
temperature during dwell time. The model assumes an equi-biaxial stress state
function of the analyzed point in the ceramic thickness $y$ :

$\sigma(y)=\frac{E}{1-\nu}\big{(}-\Delta\alpha\Delta T_{sub}+(y/h)\alpha\Delta
T_{sur/sub}\big{)}$ (4)

Using this model, the temperature evolution during cooling from experimental
data is used to evaluate BR-slow and BR-fast conditions, as shown in Figure 7.
To provide a benchmark, S-FCT cycle is examined, assuming a linear cooling
rate of $3.3\text{\,}\mathrm{K}\mathrm{/}\mathrm{s}$ and no gradient (i.e.,
$\Delta T_{sur/sub}=0$ in equation 3). The energy is normalized based on its
value obtained for BR at the minimum temperature shown in Figure 17(a) to
facilitate comparative analysis.

The normalized energy release rate curves suggest a faster reaching of the
maximum energy level of the system at higher cooling rates. The asymptotic
values of the energy release rate at the minimum temperature are identical for
both BR-slow and BR-fast conditions, but a significant transient difference is
observed: BR-slow shows a monotonic energy increase, while BR-fast shows a
peak in the first few seconds of cooling. As demonstrated in [12], the cooling
of the coating relative to the substrate causes tensile stresses in the upper
layer, resulting in peak energy release rates which may exceed those observed
in the cooled state. This tensile stress phenomenon was further validated by
finite element modeling of thermal gradient cycling in [9]. Surprisingly,
S-FCT achieves a higher steady-state energy than the burner rig conditions,
contradicting the experimental results. The model proposed by Levi et al. [12]
explains that the energy release rate in coatings subjected to thermal cycling
is influenced by two main factors: the differential contraction arising from
the mismatch in thermal expansion coefficients between the substrate and the
coating, and the self-constriction of the coating due to cooling at its
surface. In scenarios where the cooling rate is slow for thermal gradient
cycling conditions, the gradient’s effect can help alleviate the strains and
stresses experienced by the coating, potentially lowering the energy release
rate. This inconsistency with experimental results suggests a potential
sensitivity of damage to cooling rates, implying that faster energy input may
result in greater damage.

The curves in Figure 17(a) were contrasted with those generated using Vaßen et
al.’s model from [34]. The latter estimates the energy release rate stored in
the coating by integrating the in-plane stresses across the thickness of the
spalling coating and incorporating these values into the energy release rate
of a straight interface crack. This model follows:

$U(t)=\tau\frac{1-\nu^{2}}{2(1-\nu)^{2}}\int_{y}^{h}E(t,x)\Big{\\{}\Delta
T_{sur/sub}\alpha+\Delta T_{sub}\Delta\alpha\Big{\\}}^{2}dx$ (5)

where $\tau$ is a parameter fitted to match the observed time to spallation,
other parameters being consistent with the notations used for Levi’s model in
equation 3.

The resulting normalized energy release rate curves are shown in Figure 17(b).

Comparing Levi’s model with Vaßen et al.’s [34], the latter presents a lower
stabilized energy release rate for furnace cycling, aligning better with
experimental results. However, it fails to identify an energy peak for BR-fast
at the start of cooling.

(a) Using model from [12]

.

(b) Using model from [34]

.

Figure 17: Normalized elastic energy per unit area in the coating using
experimental measurements of thermal condition on cooling

The thermal loading analyzed in section 3.2 shows that much higher
temperatures were reached at the blister sites, with these local temperatures
increasing during cycling as a result of blistering. Considering the
temperature profiles extracted at the maximum temperatures at the blister
locations from infrared thermography measurements during cooling to
$400\text{\,}\mathrm{\SIUnitSymbolCelsius}$, normalized energy curves based on
Levi’s model, eq. 3, are plotted in Figure 18 for Sample A tested with the BR
slow cycle, at cycle 250 and cycle 500. This approach is a rough approximation
of the mechanical state as Levi et al.’s model [12], subsequently used,
ignores the change in stress state in the presence of a blister. The aim is to
evaluate the trend in terms of the effect of overheating on the mechanical
state. This analysis is subsequently only qualitative.

Observing the cooling process at cycle 250, it is noteworthy that the elastic
energy release rate calculated at the adherent top coat area exceeds that at
the blister sites, Figure 18(a). As explained earlier, according to Levi et
al.’s model, a mild thermal gradient throughout the top coat can be beneficial
and potentially lower the energy release rate. This is evident in the case of
blisters A-1, A-2 and A-3 compared to the adherent top coat at N=250. However,
as the gradient increases for A-2 and A-3, it amplifies the energy release
rate compared to the adherent top coat. In fact, at cycle 500, the elastic
energy release rate has escalated for blisters A-2 and A-3, with A-2 showing
higher energy values, Figure 18(b). Figure 19 takes into account comparable
maximum temperatures at the blister sites for BR-slow and BR-fast, and cooling
down to 400°C, which occurs within approximately 20 seconds under fast cooling
rates, and around 70 seconds under slow cooling rates. The normalized energy
curves show exceptionally high energy values for blister C-3, which peaks at a
normalized value of about 5.

Thus, limiting the analysis to the adherent temperature alone fails to provide
strong insights. However, considering the actual thermal loading at blister
locations reveals higher steady-state energy during both peak and minimum
temperatures in burner rig cycling compared to furnace cycling.

(a) BR-slow (A) at N = 250

(b) BR-slow (A) at N = 500

Figure 18: Normalized elastic energy per unit area in the coating, based on
Levi’s model eq. 3 [12]

Figure 19: At blisters for BR-fast and BR-slow during cooling down to
$400\text{\,}\mathrm{\SIUnitSymbolCelsius}$, based on Levi’s model eq. 3 [12]

.

In conclusion, both models provide valuable insights into both asymptotic and
transient aspects. The analysis further emphasizes that the cooling rate
transient significantly influences the mechanical state in the TBC.

Local overheating of the blister dictates the steady-state behavior,
highlighting the strong influence of the blister on the mechanical loading of
the TBC. These results motivate numerical solutions to evaluate a more
realistic mechanical state, taking into account both the temperature gradient
and the mechanical evolution of the system, as well as the modification of the
blister dimensions during thermal cycling. The strong correlation between
local maximum temperature and blister height is a key factor in the local TBC
stress as shown in Figure 14(d) and should be considered in further studies.
The coupling between oxidation rate, growth strain and BC creep has been
established as a crucial parameter controlling the mechanical state of the TBC
considering the homogeneous temperature in the FCT [9]. This drastic
temperature gradient should therefore play a key role in the local loading,
i.e. the stress intensity factor or other driving force controlling the damage
rate.

Last but not least, the measurement of temperature in FCT specimens is not as
precise compared to the in-depth analysis of BR experiments proposed in this
study: in FCT specimens, thermocouples do not take into account either the
temperature gradient in the systems or the thermal inertia of the specimen,
which could affect the above analysis. Furthermore, a more accurate
temperature measurement should be considered for each experimental condition
in order to conclude on the transient aspects in the temperature evolution and
the subsequent impact on the TBC damage, this possibility being also offered
by the LASDAM method.

##  6 Conclusions

This study demonstrates the efficiency of the LASDAM method for testing
complex configurations, accounting for thermal gradients, and notably,
avoiding edge effects. In essence, damage localization is initiated by the
presence of the initial blister, enabling robust monitoring of the damage rate
at that specific location. The harsh environment induced by the burner rig
condition has been thoroughly characterized through surface analysis of
temperature. The in-depth characterization of the damage rate highlights the
substantial impact of temperature and, more crucially, cooling rate on damage
evolution. The through-thickness temperature gradient is more challenging to
analyze, but a comparison of thermocouple temperature measurements within the
volume of the substrate and surface temperature measurements indicates that
the inversion of the temperature gradient during cooling for a higher cooling
rate must drive the higher damage rate compared to a lower cooling rate.

For the first time, it is revealed that the top coat layer experiences very
high local temperatures due to the blister and the associated air gap between
TC and oxide. This scenario mirrors real component applications where
delamination precedes final spallation. This local temperature increase is
observed both through in situ temperature field measurement with IRT and the
sintering of the TC: proximity to the center of the blister corresponds to
higher temperatures and increased sintering.

Comparison with furnace thermal cycling is difficult due to uncertainties in
local interface temperature. However, the cooling rate in burner rig testing
is typically higher than in FCT testing. As demonstrated for APS coatings,
both the higher local temperature and the higher cooling rate have a
synergistic effect on the increased damage rate observed under burner rig
conditions compared to FCT conditions. In summary, the following findings were
thoroughly addressed in this study:

  * 1.

The introduction of pre-defect using the LASDAM method allows in-situ
monitoring of temperature and damage evolution;

  * 2.

A fast cooling rate has been shown to increase the damage rate compared to a
slow cooling rate;

  * 3.

Local overheating is a direct function of the air gap between the blister and
the metallic bond layer. SEM analysis shows that the overheating is high
enough to promote sintering of the TC layer at the top of the blister, while
the adherent TC remains unaffected;

  * 4.

Sintering, oxide thickness and phase transformation tend to justify that the
temperature at the crack tip associated with the artificial blister is
consistent with the adherent areas of the TC, thus limiting the artifact
associated with overheating;

  * 5.

However, the mechanical state of the blister is significantly altered by this
overheating;

  * 6.

Models that consider the elastic energy stored in the TC layer underscore the
critical role of fast cooling rates, particularly evident in the Levi’s model,
where thermal transients result in a sudden increase of elastic energy at high
temperature;

  * 7.

Comparisons with standard furnace cycling tests are limited due to data
constraints.

The significant influence of the thermal transient has been clearly
demonstrated using the LASDAM technique and should be further strengthened in
the FCT analysis. These aspects should be addressed in future publications
through a combination of finite element analysis and correlation with
experimental temperature measurements. In addition, the LASDAM method has the
potential for multiple shock analysis, allowing further exploration of in-
plane gradients, which should be developed in future work.

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