Heinrich J.G., Aldinger F. (Eds.) Ceramic Materials and Components for Engines

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Acknowledgments

BET measurements were carried out by Dr. L.-M.

Berger, Fraunhofer Institute of Ceramic Technolo-

gies and Sintered Materials, Dresden.

Careful

sam-

ple preparation by J. Brien, C. Kroder, H. Mangers

and H. Schurmann

highly appreciated.

References

R.B. Dinwiddle, S.C. Beecher, W.D. Porter,

B.A. Nagaraj: The effect of thermal aging on

the thermal conductivity of plasma sprayed and

EB-PVD thermal barrier coatings, ASME

pa-

per 96-GT-282 (1996)

F. Sziics: Thermomechanische Analyse und

Modellierung plasmagespritzter und EB-PVD

aufgedampfter Wi4rmedWmschichtsysteme

fiir

Gasturbinen, Fortschr.-Ber.

VDI

Reihe

Nr.

18. Dusseldorf.

VDI

Verlag 1998

J.A. Haynes, M.K. Ferber, W.D. Porter, E.D.

Rigney, Oxid. Metals

(1999) 31-76

H.E. Eaton, R.C. Novak,

Surf.

Coat. Technol.

L.M. Berger: Powder and compact porous mate-

rials characterization by adsorption, Roc. 1992

Powder Metallurgy World Congress, Vo1.8,

J.M. Capus and R.M.

German

eds.,

(1987) 227-236

MPIF/APMl

1992,235-247

Brunauer, P.H. Emmett, E. Teller, J. Amer.

Chem. SOC.

(1938)93

K. Fritscher, M. Schmucker, C. Leyens, U.

Schulz, Mater.

Sci.

Forum

251-254

(1997)

U. Schulz, M. Schmucker, Mater.

Sci.

Eng.

L. Lelait: "Etude microstructurale fine de rev&

tements ceramiques de type baniere ther-

mique", Ph. D. thesis, Universite

Orsay,

France 1991

965-970

A276

(2OOO)l-8

10. R.M.

German,

Powder metallurgy science 2nd

edition,

MPIF

Princeton, New Jersey 1994,

11.

Smelzer, P.K. Talty, J. Amer. Cer. SOC.

12.

Knopp, N. Htife, W. Weppner, P. Kountou-

ros, H. Schubert, Science and Technology of

Zirconia

Technomic Publication Comp.,

Lancaster, Pennsylvania 1993, p. 567-575

13. U. Schulz, K. Fritscher, C. Leyens: Two-source

jumping beam evaporation for advanced

EB-PVD

TBC

systems, ICMCTF 2000, San Di-

ego, accepted for publication

p.250-252

(1975) 124-130

522

CRACK PROPAGATION IN A THERMAL BARRIER COATING SYSTEM

Blandin*, R.W. Steinbrech,

Singheiser

Forschungzentrum Julich, Institut fur Werkstoffe und Verfahren der Energietechnik

52425

Julich, Germany

ABSTRACT

Yttria stabilized zirconia (YSZ) coatings with a cor-

rosion resistant MCrAlY bond coat provide an effective

thermal protection for advanced gas turbine components

of Ni-based superalloys. The crack propagation trough

the top YSZ layer and the intermediate metallic bond

coat of a thermal barrier coating system has been stud-

ied. Three-layer specimens

IN 792 substrate

NiCo-

CrAlY bond coat

air plasma sprayed YSZ thermal

barrier coating (TBC)

were mounted on a four-point

bending device with the ceramic layer on the tensile

surface. Isothermal experiments were conducted be-

tween room temperature and 950°C. During deforma-

tion, the load/deflection curve was monitored and the

propagation of cracks was observed in situ.

At room temperature, cracks first appeared

the

zirconia TBC at about

0.2

strain. Under further de-

formation, the cracks grew through the whole ceramic

layer, finally entering and causing fracture of the metal-

lic bond coat. Residual stresses

ceramic and bond

coats were determined by comparing strength values

from literature with the measured bending stress at

which a crack appeared in each layer. At high tempera-

ture, cracks stopped at the interface between ceramic

TBC and metallic NiCoCrAlY layer. This behavior

reflects the high ductility of the bond coat material. The

change from brittle fracture

non-fracture of the bond

coat was utilized to estimate the ductile-to-brittle transi-

tion temperature (DBTT) of the material. The

stress/strain considerations used in the present approach

and the residual stresses in the thermal barrier system

are discussed.

INTRODUCTION

Thermal barrier coatings (TBCs) are increasingly

applied to gas turbine components to improve service

life and efficiency

[l].

The combination of a plasma

sprayed yttria stabilized zirconia

(YSZ)

top coating and

an oxidation resistant MCrAlY bond coat (M

Ni, Co)

is one of the most commonly used TBC systems

[2-31.

The YSZ coating protects the turbine materials from hot

spots and reduces the average temperature of the coated

component, e.g. a

300

pm TBC reduces the temperature

by 80°C

[3].

The primary failure mode of

TBCs

spallation,

typically occurring after extended cyclic thermal expo-

sure during cooling down from operation temperature

[4].

As one of several failure mechanisms, spallation

results from cracks which grow along the interface be-

tween top and bond coats when the ceramic layer is

under compressive stress.

The determination of residual stresses in the YSZ

top coat resulting from spraying process and thermal

expansion misfit between the bonded layers

[5]

is es-

sential to estimate the critical failure stresses of TBCs.

In recent years, much work focussed on the characteri-

zation of residual stresses in TBC systems, using the

established sin2'€'- and hole drilling methods, the layer

removal technique

[6]

the changes in curvature of

TBC specimens

[7-81.

In the present study, tensile strain is applied to the

three-layer composite in bending experiments. Using a

four-point bending device, the specimens are strained

between room temperature and 950°C. The cracking in

ceramic TBC and metallic BC is observed in situ with a

high temperature telescope system. The work focuses on

the initiation and propagation of cracks in the top coat

and their behavior at the interface with the bond coat.

The residual stresses in top and bond coat are de-

duced from first appearance of cracking and strength

values reported in the literature. The experimentally

determined residual stresses are compared to a theoreti-

cal stress distribution in TBCs, calculated with a linear

thermoelastic approach. Finally the behavior of the

NiCoCrAlY material is discussed with respect to tem-

perature and ductility.

THEORETICAL ASPECTS

During bending, the stress at a position

of the

composite material is given by:

where

is the residual stress,

the thermal stress

(when the experiment is performed at high temperature)

and

the bending stress.

Fig.

shows a three-layer specimen strip and the

adopted geometric notations. The surface of the YSZ

top coating is chosen as origin

(y=O).

O(Y)

(y)

O'f

(Y)

(1)

Fig.

Geometry

a three-layer

TBC

composite

523

In this approach, materials are considered to be iso-

tropic and homogeneous with a pure elastic response to

external forces. As long as the three materials behave

elastically, the three-layer system is representative of

the TBC system

(k=3).

Due to thermal expansion mis-

match, the TBC experiences tensile stresses at elevated

temperatures, causing cracking of the TBC. Once the

top coat is cracked, only the two metallic layers have to

be taken into account

(k=2).

If cracking also occurs in

the MCrAlY bond coat, the bending behavior of the

TBC specimen is dictated by the substrate only.

THERMAL STRESS

We assume that the substrate layer is thick enough to

maintain the composite specimen straight during all

manufacturing steps and during temperature exposure

without mechanical loading. In this case, the thermal

stress at a position

can be calculated as a function

of the thermoelastic and geometric parameters of each

layer

[8]:

E)dj

(2)

j=l

where

is the biaxial elastic modulus of each layer

(i=l

...

with

k=l

...

3),

the thermal expansion

coefficient and

the thickness.

and

define the

temperature before and after heating.

BENDING

STRESS

at a posi-

tion

proportional to the distance from the neutral

axis and inversely proportional to the radius of curva-

ture at the neutral axis

Following bending theory, the stress

with

the elastic modulus of the layer

(i=

. .

with

.3)

and

the position of the neutral axis

[9].

From the equilibrium equation of bending stresses

i=l

the position of the neutral axis can be determined as a

function of the elastic and geometric parameters:

AEidi[2gdj

+di\

i=I

j=1

(5)

Eidi

i=l

The radius of curvature at the neutral axis is directly

related to the deflection of the beam

(Fig.

2):

where

is the span length of the four-point bending

device.

Using this geometric relation, equation

becomes:

Fig.

Deflection and curvature

a bending specimen

EXPERIMENTAL

mm-thick plates of the Ni-based superalloy IN

792

were coated

(IWV1,

Forschungszentrum Jiilich) with

industrial NiCoCrAlY variant by vacuum plasma

spraying. After a two-step heat treatment

h at

112OOC

and

h at 850°C), a ceramic coating

(ZrOz

7-8

Yz03) was air plasma sprayed onto the substrate

bond

coat composite. The thickness of metallic bond coat and

ceramic top coat was

100

and

340

respectively.

Fig.

3 shows a cross-section micrograph

a plasma

sprayed TBC system. The morphology of the ceramic

layer is typical of plasma sprayed materials, consisting

of splats formed during the deposition of the molten

particles. Fig.

also reveals a high concentration of

defects (characteristic pores and microcracks). Image

analysis indicated for this material a porosity of about

15%.

Substrale

Fig.

Cross-section micrograph

a plasma sprayed

TBC

system

Some relevant thermoelastic data of the three mate-

rials

are

given in Tables

and

as a function of tem-

perature

[lo-131.

524

9.1

1.8~10'~ T

1.6~10-'~ T2

1.7~10-~' T3

Top coat

12.4

3.7~10" T

2.7~10-~ T2

4.8~10-~ T3

9.8

1.3~10-~ T

1.9x10-' T2

1.2~

O-*

Bond coat

Substrate

[

101

Table

Temperature dependence

the thermal expansion coefficient

for the three materials.

Thermal expansion coefficient measured with free standing materials

Temperature RT 600°C 950°C

Topcoat

[ll]

Bondcoat* 122 109 99

Elastic modulus

(GPa)

Substrate

136 114 90

Top coat

[

121 35

4-point tensile

strength (MPa)

strength

Bond coat

[

131 980

(MPa)

700

Table

Temperature dependent elastic properties.

Elastic modulus measured in bending on thick materials

Specimen strips (47

4 mm2) were prepared from

the composite material. The thickness of the substrate

was mechanically reduced to produce a specimen thick-

ness of 3 mm. Finally one side face was polished using

a final 1 pm diamond paste.

Isothermal four-point bending tests were performed

with the three-layer specimen strips from room tem-

perature to 950°C. The bending tests were controlled by

the deflection of the specimen. A total deflection of 1

mm was applied in 30 minutes. For high temperature

tests, the specimens were mounted in the four-point

device under a

N pre-load and heated up with a rate of

8"C/min. The bending tests were started after a two-

hour holding time at the chosen temperature.

All bending strips were positioned in the bending

device

that tensile stress developed in the top coat

(Fig. 4). Since the bending device was mounted in a

furnace with a silica window, in situ observation of the

polished side face could be carried out by using a high

temperature telescope system.

Specimen

Fig.

Four-point bending device

Load

and deflection

were recorded and used to

calculate the equivalent modulus of elasticity in bending

E of the composite material using

[

141:

(8)

43L2

-4a2)

4bd3

with L

the span length,

the distance

from the support to the load applicator, b=4

the

specimen width and

mm the specimen thickness.

RESULTS

LOAD/DEFLECTION CURVES

Fig.

shows the applied loads as a function of de-

flection at room temperature (RT),

600

and 950°C.

At RT and 600"C, the loaddeflection curves can be

approximated by straight lines indicating that the com-

posite material behaves elastically during the entire

experiment. At 950"C, the elastic behavior is limited to

deflections below 300 pm.

1000

800

600

400

200

400

600

800

Deflection

(pm)

Fig.

Loaddeflection experimental curves

The equivalent elastic modulus of the composite is

proportional to the slope of the straight part of the

loaddeflection curve (Equation

8).

The obtained moduli

are presented in Table

as a function of temperature

and compared

theoretical values

(9)

Eldl

E2d2

E3d3

ETh

Temperature RT 600°C 950°C

1.180 0.909

0.585

Slope of the linear part

(Npm-')

120 93 60

125 106

Experimental equivalent

modulus (GPa)

Theoretical equivalent

modulus (GPa)

Table

Comparison between experimental and theoretical equivalent

elastic modulus in bending of the three-layer specimen

With increasing temperature, the elastic modulus de-

creases (Table 2). At RT and 6OO0C, both experimental

and theoretical moduli are similar. At 950"C, the theo-

retical modulus is about 30% higher than the measured

one, probably resulting from the larger error in slope

determination.

525

IN SITU OBSERVATIONS

Fig. 6,

and

illustrate the crack propagation in the

TBC system at room temperature.

Cracks were first observed at deflections between

250

and

300

pm (Fig.

6).

Bending cracks do not start at

the ceramic surface but seem to initiate Dreferablv at

penetrate the BC, e.g., pores located near the interface

of TBC and BC play the role of crack stoppers. Cracks

enter into the pore and do not cross the interface (Fig.

b). Additional strains are necessary to induce further

propagation.

Fig.

Crack initiation at two locations

the

YSZ

top coat at

for

deflections between

250

and

300

Fig.

shows the crack pattern at two locations of the

top coat at the end of the bending experiment

(f==950

pm).

Fig. 9: In situ observation at 600°C: crack a) initiation at 185

deflection and

propagation at

500

deflection in the top coat

Fig. 7: Crack pattern at two locations

the TBC system a1

for

950

deflection

Fig.

10:

Crack pattern at two locations

the TBC system at 600°C

for

a 970 pm deflection

Once the interface is reached, cracks grow into the

brittle bond coat following a straight trajectory within

only a few micrometers of deflection (Fig. 7-a). They

stop finally in the tougher substrate material, dissipating

the crack driving force into the plastic zone in front of

the crack tip (Fig.

8).

However, cracks not always

In the ceramic top coat, the first crack is observed at

a deflection of about

185

pm (Fig. 9-a). As room tem-

perature behavior, the cracks initiate preferably at big-

ger defects near the top coat surface. Once initiated, the

526

cracks propagate again towards the interface between

top coat and bond coat. At 500 pm deflection, they

cover about two thirds of the total ceramic thickness

(Fig. 9-b) and reach the interface with the NiCoCrAlY

bond coat at 700 pm deflection. However, none of the

cracks enters the bond coat layer, even when additional

bending strain is applied. The crack stop is also inde-

pendent from the presence of pores near the interface

(Fig.

lo).

Fig. 10-a and b show two cracks in the ce-

ramic top coat for a 970 pm deflection.

At 950°C testing temperature, cracks are also ob-

served in the ceramic layer, without any mechanical

loading. Again pores near the surface are the origin

(Fig.

11).

The formation of cracks probably results from

the tensile thermal mismatch stress, which develops in

the ceramic layer once the initial compressive residual

stress is relaxed. With additional bending strain, the

thermal cracks of the YSZ-layer grow again towards the

interface with the bond coat and arrest there. Even

higher strains do not induce crack propagation. During

cooling, the bond coat contracts faster than the top coat

so that compressive stresses develop in the ceramic and

the crack opening diminishes (Fig. 12).

DISCUSSION

The bending stress as a function of deflection in a

multilayer specimen is given by equation

At room

temperature, major cracks are first observed in the ce-

ramic top coat at deflections between 250 and 300 pm.

Below 250

pm,

the YSZ layer is intact and bending

strain and stress can

be calculated by using a three-layer

model

3). The obtained strain and stress distribu-

tions are shown in Fig. 13. Cracks first appear in the

YSZ top coating at a bending strain and stress of about

EB=0.2

and oB=90 MPa, respectively. Since the

tensile strength of the porous TBC material is about

(TC

35 MPa (room temperature and as sprayed condi-

tions, Table 2,

[

12]), we can deduce that the ceramic top

coat is under a relatively low compressive residual

stress

~O=oc

-og

MPa.

0.20

0.00

-0.20

250

Bond coat

-T

Substrate

YS?

Fig. 11:

situ observation of the ceramic top coat cracks before

bending test a) at

and b) at 950°C

-250

Fig.

13:

a) Bending strain and b) bending stress in a three layer

speci-

men for a 250 pm deflection

Fig. 12: Crack in the

YSZ

top coating a) at 950°C for

nun

deflection

and b) after cooling to

Deflections for crack initiation in the ceramic top

coat and propagation through the bond coat

are

resumed

Table 4.

Temperature RT 600°C 950°C

Crack in the top coat

250 pm

Thermal

185pm

cracks

Crack in the bond coat 950 pm Cracks don't enter

Table

Deflection at observation

cracks in top coat and bond coat

The bending cracks reach the interface between top

and bond coat at deflections of about 950

pm.

At this

stage of the bending experiment, the ceramic coating is

cracked over its entire length (segmentation) and does

not influence further the behavior of the composite

material. A two-layer model has been adopted to calcu-

late the strain and stress

2). At 950 pm deflection,

the bond coat faces bending strains of about 0.6

and

tensile stresses between

10 and

765

MPa. These values

are much lower than the strength of MCrAlY materials

oc=980 MPa (Table 2, [13]), indicating that the bond

coat is under a tensile residual stress of about

00~215

MPa.

Independently, equation 2 can be used to estimate

the theoretical residual stress distribution in three-layer

specimen strips. Considering the thermal and mechani-

cal history of specimen preparation, the following steps

have to be considered:

Heat treatment of bond coat

substrate composite

24h at 850°C. We assume that the composite is

stress free.

527

Cooling down of the two-layer specimen to RT and

heating up to 300°C for the plasma spraying of the

ceramic TBC.

Spraying of the ceramic top coat: the molten YSZ

particles are rapidly cooled down. The high tensile

stresses, which should result in the TBC

are

consid-

ered to relax through the formation of microcracks.

Thus the tensile strength of the YSZ-TBC provides

an upper boundary value for the remaining residual

TBC stress.

Cooling down of the three-layer system to RT.

As shown in Table 4, the residual stresses in the

YSZ

top coat and the NiCoCrAlY bond coat, which are

calculated on these assumptions, are quite similar to the

values obtained from the bending experiments.

flection of the specimen reaches 1 mm, which corre-

sponds to bending stresses at the interface with the top

coat of 720 and 650 MPa at 600°C and 950°C respec-

tively. In any case, the resulting stresses of 780 and 610

MPa should be sufficient to induce cracking in the me-

tallic bond coat. However, the cracks stop at the inter-

face between top coat and bond coat (Fig.

and 12),

which proves the high ductility of this metallic layer.

If we assume that the BC only fractures in a brittle

mode below the ductile-to-brittle transition temperature

(DBTT), then 600°C seems to be above this temperature

and it is unlikely that the calculated stresses can be

maintained. The conducted experiments lead to the

conclude that the DBTT of the NiCoCrAlY layer is

below 600°C.

Top coat Bond coat

Bending experiments

-55

MPa 215 MPa

Thermoelastic calculation -63 MPa 200 MPa

Table

Residual stresses in top and bond coat

In the case of the high temperature tests, thermal

stresses have to be added to the residual stresses to

obtain the stress distribution in the TBC system before

bending. Fig. 14 shows the respective stress distribu-

tions at RT, 600 and 950°C.

Bond coat

Substrate

-4

200

-0-

600°C

-0-

950°C

Fig.

14:

Stress

situation in

TBC

specimen at RT.

600

and

950°C

With increasing temperature, the compressive resid-

ual stress in the top coat relaxes and tensile stress de-

velop. Thus the ceramic layer expands at a lower rate

than the two metallic materials (Table

1).

At 950"C, the

calculated thermal stress

(T~~~~~=~O

MPa is already

significantly above the bending strength of the YSZ

layer and cracks are formed thermally before any me-

chanical loading (Fig. 11).

Following the same thermoelastic considerations,

nominal bond coat stresses can be derived. In the bond

coat, the tensile residual stress relaxes

(60

MPa at

600°C) and compressive stress develops

(-40

MPa at

950°C). At the end of the bending experiment, the de-

CONCLUSION

Using in situ observations of cracking in a TBC

system during bending, the critical stresses of ceramic

top coat and metallic bond coat were determined. Re-

sidual stresses of about

-55

MPa and 215 MPa in top

and bond coat were deduced respectively. These values

are in a good agreement with residual stresses predicted

from linear elastic calculations. With increasing tem-

perature, tensile stresses develop in the YSZ top coat as

a result of thermal expansion mismatch. They are suffi-

cient to induce thermoelastic cracking at 950°C.

At high temperature, the cracks do not penetrate the

metallic bond coat, reflecting the high ductility of the

NiCoCrAlY material. The change from brittle fracture

at RT to non-fracture at 600 and 95OOC indicates that

the ductile-to-brittle transition temperature (DBTT) of

the material is below 600°C. Complementary bending

tests at temperature between RT and 600°C would allow

a more precise evaluation.

ACKNOWLEDGEMENT

The authors are grateful to Ms. Funke, IWV1, For-

schungszentrum Jiilich, for providing the coated speci-

mens.

REFERENCES

W.J. Brindley, Thermal Barrier Coatings of the

Future, J. of Thermal Spray Technol., 6[1], (1997)

3-4.

Stecura, Optimization of the NiCrA1Y/Zr02-

Yz03 Thermal Barrier System, Adv. Ceram. Mat.,

P. Fauchais, A. Vardelle and M. Vardelle, Recent

Developments in Plasma Sprayed Thermal Barrier

Coatings, Workshop on Thermal Barrier Coatings,

85" Meeting of AGARD, Structures and Materials

Panel, 13-17 October 1997, Aalborg, Denmark.

R.V. Hillery, B.H. Pilsner, R.L. McKnight,

T.S.

Cook and M.S. Hartle, Thermal Barrier Coating

11, (1986) 68-76.

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Life Prediction Model Development, NASA Con-

tractor Report

180807, (1988).

M.K. Hobbs and H. Reiter, Residual Stress in ZrOz-

8%Y203

Plasma Sprayed Thermal Barrier Coatings,

Thermal Spray: Advances in Coatings Technology,

ASM International,

(1988) 285-290.

P. Bengtsson and C. Persson, Modelled and meas-

ured residual stresses in plasma sprayed thermal

barrier coatings, Surf. Coat. Technol.,

92, (1997)

S.C. Gill and T.W. Clyne, Investigation of residual

stress generation during thermal spraying by con-

tinuous curvature measurement, Thin Solid Films,

G. Blandin, R.W. Steinbrech and L. Singheiser,

Response of Thermal Barrier Composites to Tem-

perature Exposure, EUROMAT'99,

27.-30.9.1999,

Munich, Germany.

78-86.

250, (1994) 172-180.

G.H. Ryder, Strength of materials

(1970).

COST

501,

Advanced Blading for Gas Turbines,

3rd

Annual Report

(

1992).

D. Basu, G. Blandin,

Singheiser and R.W. Stein-

brech, Elastic Behavior

Thermal Coatings: A

Comparison of Mechanical and Thermoelastic

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W. Mannsmann, Keramische Warmedammschicht-

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mechanischer, thermischer und thermomechanis-

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M.G. Hebsur and R.V. Miner, High Temperature

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529

This Page Intentionally Left Blank

POROSITY GRADED SILICON CARBIDE EVAPORATOR TUBES

FOR GASTURBINES WITH PREMIX BURNERS

M. Droschel",

Oberacker", M.

Hoffmann", W. Schaller**, D. Munz**

(*)

Institut fur Keramik im Maschinenbau, Universitat Karlsruhe

(TH)

D-76131 Karlsruhe, Germany

(**)

Institut fur Materialforschung, Forschungszentrum Karlsruhe

D-76021 Karlsruhe, Germany

ABSTRACT

Porous silicon carbide (Sic) ceramics are promising

materials for liquid fuel evaporator tubes in gas turbine

combustors. For evaporator tubes with four different

design variations (two-layer concept and three continu-

ous porosity gradients) finite element method calcula-

tions were made to determinate the thermal stresses due

to the temperature gradient during steady state and

transient operation. By comparing the local stress dis-

tribution with the local strength it could be shown that a

tailored porosity gradient is necessary to meet the local

stress/ strength requirements. For tubes with such con-

tinuous porosity gradients a new processing route based

pressure filtration has to be developed. The forma-

tion of one and more dimensional wax concentration

gradients in the filter cake (corresponding to porosity

and pore size gradients after sintering) is adjusted by

controlling the composition of a mixture of Sic- and

Sic-wax slurries and the filtration pressure. The design

concept of a laboratory casting apparatus which uses

this new concept and an experimental/ numerical

method for the derivation of the process parameters is

also described in this paper.

INTRODUCTION

A concept for a gas turbine combustor with premix

burner is shown in Fig.

The evaporation of the liquid

fuel and the burning zone are decoupled

[

11. The fuel is

sprayed

the outer surface of the Sic-evaporator tube.

Due to the flowing air from the compressor, a thin film

of fuel is formed

the outer surface of the evaporator,

which vaporizes completely. The homogeneous mix of

air and fuel vapor enters the reaction zone inside the

tube where the combustion takes place. It is known, that

a porous evaporator surface allows much higher evapo-

ration rates compared to a dense material. This could

result in advantages with respect to the design of the

combustion system, if its reliability is maintained de-

spite of the porosity.

However, the evaporizer tube has to be gas tight

the inner side and it should be porous at the outer sur-

face, where the evaporation takes place [2]. Between

this porous outer surface and the dense inner surface a

porosity gradient should be introduced.

POROSITY GRADIENT DESIGN

LATIONS

BY FINITE ELEMENT CALCU-

GEOMETRICAL BOUNDARY CONDITIONS

The evaporator is an axial symmetric component

with an inner radius of ri

22.5 mm. The inner surface

consists of a dense Sic layer with 2.0 mm thickness

which is followed by a porous Sic-layer, serving as an

evaporator (Fig. 2).

radius

nozzle

Fig.

Concept ofa gas turbine combustor with premix burner

[I].

porosity

"/o

Fig.

Design ofthe evaporator tube.

For the porosity gradient different design variants

were studied. According to a previous work [2] four

design variants of the evaporator tube fulfill the de-