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Doctoral Thesis

Fast and robust joining process for aerospace components bylocal heating of paste adhesives

Author(s): Sánchez Cebrián, Alberto

Publication Date: 2014

Permanent Link: https://doi.org/10.3929/ethz-a-010164453

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DISS.-No. ETH 21842

FAST AND ROBUST

JOINING PROCESS

FOR AEROSPACE COMPONENTS

BY LOCAL HEATING

OF PASTE ADHESIVES

ALBERTO SÁNCHEZ CEBRIÁN

2014

DISS.-No. ETH 21842

FAST AND ROBUST JOINING PROCESS

FOR AEROSPACE COMPONENTS BY

LOCAL HEATING OF PASTE ADHESIVES

A thesis submitted to attain the degree of

DOCTOR OF SCIENCES OF ETH ZÜRICH

(Dr. sc. ETH Zürich)

presented by


Dipl. Industrial Engineer Universitat Politècnica de Catalunya (2009)

born 20.07.1985

citizen of Spain

accepted on the recommendation of

Prof. Dr. Paolo Ermanni, examiner Dr. Markus Zogg, co-examiner

Prof. Dr. Heinz Voggenreiter, co-examiner

2014

i

Allá donde estés, esto va por ti.

iii

Summary (English)

Since the 70s, industry is doing research in new adhesive sys-

tems for joining composites for structural applications. Adhe-

sive bonding permits having lighter structures and thus more

efficient solutions than today’s aerospace state of the art, rivet-

ing. Additionally, the use of adhesives does not require drill-

ing the components, which is especially beneficial for joining

composite materials.

Typically, the curing of paste adhesives is carried out in forced

convection ovens, applying isothermal heating cycles recom-

mended by the supplier. The variation of the curing tempera-

ture during the process is today not considered due to its high

energy consumption. To avoid this limitation, the use of in-

duction heating to heat locally the bondline is considered in

this doctoral thesis. This technique allows heating the assem-

bly faster and thus saving energy resources and time. In this

context, the acceleration of the curing process of a paste adhe-

sive by increasing the heating temperature during the process

is investigated.

However, the increment of the curing temperature leads to

void formation and thus to the thermal degradation of the

paste adhesive. This phenomenon is caused by the evapora-

tion of moisture, the degradation of chemical components and

the expansion of air trapped during the mixing process. The

increment of void content affects the mechanical performance

of the bonded system. To deal with this limitation, a novel

modeling of the void formation during the curing process con-

sidering the degree of cure is developed in this doctoral thesis.

The fundamental idea of this approach is that the evaporation

of the paste adhesive that leads to void formation decreases

iv

with curing progression. Therefore, a novel thermal degrada-

tion model for the evaporation of paste adhesives is defined as

a function of temperature, time and the degree of cure. Once

the mass evaporation is modeled, the relation between evapo-

rated mass and void formation is experimentally established,

completing the void prediction model for paste adhesives.

This model is then used to find the optimal strategy to acceler-

ate the curing process of paste adhesives without increasing

the void formation. This approach is applied to two different

paste adhesive systems in order to ensure the repeatability of

the research findings. The results show a reduction of more

than 85 % of the processing time compared to the reference

curing cycle without affecting the mechanical performance of

the joint. Finally, the methodology investigated is applied to a

real aircraft component, defining the guidelines to ensure the

robustness of the process under real component geometries.

v

Zussammenfassung (Deutsch)

Seit den siebziger Jahren forscht die Industrie an neuen

Klebstoffsystemen zur Verbindung von Faserverbunden für

Strukturanwendungen. Kleben erlaubt leichtere Strukturen

und damit effizientere Lösungen als der heutige Standard in

der Luft- und Raumfahrt, das Nieten. Zudem erfordert die

Verwendung von Klebstoff kein Bohren der Komponenten,

was besonders vorteilhaft für das Fügen von

Faserverbundwerkstoffen ist. Typischerweise wird das

Aushärten von Pastenklebstoffen in Öfen mit erzwungener

Konvektion durchgeführt, wobei vom Hersteller empfohlene

Heizzyklen zur Anwendung kommen. Die Änderung der

Aushärtetemperatur während des Prozesses wird heutzutage

wegen ihres hohen Energieverbrauchs nicht berücksichtigt.

Um diese Beschränkung zu überwinden, wird in dieser

Doktorarbeit die Anwendung induktiven Heizens betrachtet,

um die Klebenaht lokal zu erwärmen. Diese Technik erlaubt

schnelleres Aufheizen und damit Ersparnisse an Ressourcen

und Zeit. In diesem Kontext wird die Beschleunigung des

Aushärteprozesses eines Pastenklebstoffs durch Erhöhung der

Heiztemperaturen während des Prozesses untersucht.

Die Steigerung der Aushärtetemperatur führt jedoch zu

Porenbildung und damit zur thermischen Degradation des

Klebers. Dieses Phänomen wird durch die Verdunstung von

Feuchtigkeit, die Degradation chemischer Komponenten und

die Ausdehnung während des Mischvorgangs

eingeschlossener Luft verursacht. Die erhöhte

Porenkonzentration beeinträchtigt die mechanischen

Eigenschaften der Klebeverbindung. Um mit dieser

Beschränkung umzugehen, wird in dieser Dissertation eine

neuartige Modellierung der Porenbildung während des

vi

Aushärteprozesses unter Berücksichtigung des

Aushärtegrades entwickelt.

Die grundlegende Idee dieses Vorgehens setzt daran an, dass

die zur Porenbildung führende Verdampfung des Klebstoffes

mit fortschreitender Aushärtung abnimmt. Deshalb wird ein

neuartiges Degradationsmodell für die Verdunstung von

Pastenklebstoffen in Funktion von Temperatur, Zeit und

Aushärtegrad formuliert. Im Anschluss an die Modellierung

der Verdampfung wird eine Beziehung zwischen verdampfter

Masse und Porenbildung experimentell hergestellt, welche das

Modell vervollständigt. Dieses Modell wird anschliessend

dazu verwendet, die optimale Strategie zur Beschleunigung

des Aushärtevorgangs von Pastenklebstoffen ohne erhöhte

Porenbildung zu finden. Dieser Ansatz wird auf zwei

verschiedene Klebstoffsysteme angewandt, um die

Wiederholbarkeit der Forschungserkenntnisse sicherzustellen.

Die Ergebnisse zeigen eine Reduktion der Verarbeitungszeit

um mehr als 85 % im Vergleich zum Referenzzyklus, ohne

Beeinträchtigung der mechanischen Eigenschaften der

Verbindung. Schliesslich wird die untersuchte Methodik auf

eine reale Flugzeugkomponente angewandt, wobei Richtlinien

aufgestellt werden, welche die Robustheit des Verfahrens bei

realen Komponentengeometrien gewährleisten sollen.

vii

Resumen (Español)

Desde los años 70, la industria investiga nuevos sistemas de

pastas adhesivas para unir estructuras de materiales compues-

tos. Su uso permite que las uniones sean más ligeras, consi-

guiendo así soluciones más eficientes que la técnica de refe-

rencia hoy en día en la industria aeroespacial, ribetes. Además

el uso de adhesivos tiene la ventaja de que los componentes a

unir no son agujereados, un hecho especialmente beneficioso

en la unión de materiales compuestos.

Normalmente, el proceso de curado de pastas adhesivas tiene

lugar en hornos de convección forzada, mediante la aplicación

de ciclos de calentamiento isotérmicos recomendados por el

fabricante. Hoy en día, la variación de temperaturas durante el

proceso no está considerada debido a su alto consumo energé-

tico. Para evitar esta limitación, en esta tesis doctoral se consi-

dera el uso de calentamiento por inducción. Esta técnica per-

mite calentar la unión rápidamente con un bajo coste de ener-

gía. Bajo este concepto, esta tesis investiga la aceleración del

proceso de curado de pastas adhesivas incrementando la tem-

peratura durante el proceso.

De todas maneras, el incremento de la temperatura de curado

está limitado por la degradación térmica de la pasta adhesiva

debido a la formación de poros cuando se aplican altas tempe-

raturas. Dichos poros se forman debido a la evaporación de

humedad, degradación de los componentes químicos y a la

expansión del aire atrapado durante la mezcla del adhesivo.

Esto provoca un descenso de las propiedades mecánicas de la

unión. Para solucionar este problema, esta tesis desarrolla un

nuevo modelo para formación de poros considerando el grado

de curado durante el proceso de calentamiento.

viii

La idea fundamental de este modelo es que la cantidad de de-

gradación térmica producida por la evaporación de la pasta

adhesiva disminuye durante proceso de curado. Por este mo-

tivo, esta tesis define un nuevo modelo de degradación en

función de temperatura, tiempo y grado de curado. Una vez el

nuevo modelo está definido, la relación entre masa evaporada

y generación de poros se establece experimentalmente, com-

pletando así un modelo para la predicción de porosidad en

pastas adhesivas.

Dicho modelo se utiliza para definir la estrategia óptima de

calentamiento de pastas adhesivas sin incrementar la porosi-

dad. Esta técnica se aplica a dos pastas adhesivas diferentes

para asegurar su respetabilidad. Los resultados muestran una

reducción del 85 % del tiempo de curado respecto al ciclo de

curado recomendado por el fabricante sin afectar las propie-

dades mecánicas de las uniones. Finalmente, la metodología

investigada es aplicada a un componente real de un avión,

definiendo las directrices para poder así asegurar la solidez

del proceso para componentes reales.

ix

Resum (Català)

Des dels anys 70, la indústria fa recerca sobre nous sistemes de

pastes adhesives per unir estructures fetes de materials com-

posts. El seu ús permet que les unions siguin més lleugeres,

aconseguint solucions més eficients que la tècnica de referèn-

cia avui en dia a la indústria aeroespacial, els rivets. Amés, l’ús

d’adhesius te l’avantatge de que els components a unir no son

foradats, un fet especialment beneficiós en la unió de materials

compostos.

Normalment, el procés de curat de pastes adhesives té lloc a

horns de convecció forçada, mitjançant l’aplicació de cicles

d’escalfament isotèrmics recomanats pel fabricant. Avui en

dia, la variació de les temperatures durant el procés no està

considerada degut a l’elevat consum energètic. Per evitar

aquesta limitació, en aquesta tesis doctoral es considera l’ús de

l’escalfament per inducció. Aquesta tècnica permet escalfar la

unió ràpid, estalviant temps i energia. Sota aquest concepte,

aquesta tesis fa recerca sobre l’acceleració del procés de curat

de pastes adhesives incrementant la temperatura del procés.

Aquest increment, però, està limitat per la degradació tèrmica

de la pasta adhesiva degut a la formació de porositat quan

s’apliquen altes temperatures. Aquests, es formen degut a

l’evaporació d’humitat, degradació del components químics i

a l’expansió de l’aire atrapat durant la barreja de l’adhesiu.

L’increment de porositat provoca una caiguda de les propie-

tats mecàniques de la unió. Per a solucionar-ho, aquesta tesis

desenvolupa un nou model per a la formació de porositat con-

siderant el grau de curat durant el procés d’escalfament.

La idea principal d’aquest model és la quantitat de degradació

tèrmica produïda per l’evaporació de la pasta adhesiva dismi-

x

nueix durant el procés de curat. Per aquesta raó, aquesta tesis

defineix un nou model de degradació en funció de la tempera-

tura, el temps i el grau de curat. Un cop el nou model està de-

finit, s’estableix experimentalment la relació entre la massa

evaporada i la formació de porositat, completant un model per

a la predicció de porositat de les pastes adhesives.

Aquesta model s’empra per a recercar l’estratègia òptima

d’escalfament de pastes adhesives sense incrementar la seva

porositat. Aquesta tècnica s’aplica a dos sistemes adhesius

diferent per a poder assegurar la seva repetibilitat. Els resul-

tats mostren una reducció de més del 85 % del temps de curat

comparat amb el cicle de referència recomanat pel fabricant

sense afectar les propietat mecàniques de les unions. Final-

ment, la metodologia recercada s’aplica a un component real

d’un avió, definint les directrius per a poder assegurar així la

solidesa del procés en components reals.

xi

Acknowledgements

The Ph.D. studies summarized in this volume have required

giving my best every single day for the last years. This has

been a huge challenge, exhausting and discouraging at many

moments, but also exciting and very gratifying.

Unfortunately, this project has sometimes required more

knowledge and motivation than what I had. At this point, I

can only express my most honest and sincere gratitude to all

the people that have helped me in many different ways, not

only with their wisdom but also with their support and opti-

mism, helping me to keep on working restless for so much

time. From the deepest of my heart, thanks a lot:

Prof. Dr. Paolo Ermanni for giving me the chance to work in a

fascinating field and complete my doctoral studies under your

supervision.

Dr. Markus Zogg for your endless support since the first day I

arrived to Switzerland and for trusting always on me.

Prof. Dr. Heinz Voggenreiter for accepting the co-supervision

of this doctoral study.

Dr. Florian Klunker for your thorough supervision and for

helping me to improve so much the quality of my research.

Dr. Stefan Cron for your technical advice during my research

and for being always ready to help me.

Beat Bucher, Ulrich Fischer, Melanie Libsig, Maria Benavente,

Dr. Sophie Cazuc, Prof. Dr. Veronique Michaud, Maxime

Roux and Yver Papier for your support and technical advice in

the Clean Sky project.

xii

Dr. Gerald Kress for giving me the change of supervising your

lecture of Mechanics of Composites during the last 3 years.

Thomas Heinrich for your help and technical advice in the

laboratories during these years.

Hans Peter Eigenmann for helping me so much in the lab and

being always so nice with me.

Dr. Stephan Busato for your help and advice with the thermal

analysis equipment.

Dr. Andrea Bergamini and Dr. Andres Arrieta for being al-

ways willing to help me.

Désirée Hess and Anke Kleint for the administrative support.

To all the students who made their semester thesis under my

supervision. Thank you all for the huge effort that you made

and for choosing me for supervising your studies: Stefano

Lucchini, Daniel Aeschlimann, Richard Hutter, Fabian Stuck-

er, Mario Bissig, Patrick Moser, Björn Luginbühl, Ralf Basler,

Simon Bühler, Christoph Walch, Lukas Mayrhofer and

Alejandro Lorenzo.

To the former Ph.D. students who helped me since my first

day at the IMES-ST. Your work was a huge motivation and

inspiration for me: Dr. Yi Liu, Dr. Grégoire Lepoittevin, Dr.

Joanna Chi-Hing Wong, Dr. Florian Bachmann and Dr. Ben-

jamin Schläpfer.

To my office mates, who always made me feel like at home:

Davi Melo Montenegro, Mattia Serra, Bryan Louis, Francesco

Previtali, Max Fickel, Sebastian Kollert and Emian Furger.

To all my other friends at the IMES-ST. It has been an honor to

work with people as wise as you. Giulio Molinari, probably

the most intelligent person I have ever known, thanks for the

help with the optimization. Tommaso Delpero, thanks for be-

xiii

ing a great support always that I needed you. Luigi di Lillo for

willing always to help me with me experiments. Simon Steiner

thanks for helping me to understand all the chemical reactions

that I was dealing with. Jesus Maldonado thanks for your help

with the cure kinetics modeling. Claudio di Fratta and Mario

Danzi thanks for sharing the responsibilities of Clean Sky with

me. Wolfram Raither thanks for the help to translate the ab-

stract to German. Also thanks Manfred Quack, Vitaly

Dmitriev and Oliver Häfner for your support.

Very special thanks to my friends that always trusted on me.

Despite the distance, I always felt that you were encouraging

me. Thanks Agustin, Alex, Marc, Albert, Ainhoa, Miquel, La-

ra, Eli, Ingrid and Xexi. Thanks to David for your support

when I felt alone during all these years in Zürich. Also thanks

to Jorge, Carlos, Nikita, Cristian and of course thanks Pons for

being so great. Thanks to my friends in Zürich, who actively

encouraged me to keep on with my work day after day: Luís,

Elena, Ximo, Pepo, David, Mireia, Alejandro, Franco, Sandra

and Bellinda. Thanks to Jano, Antonio, Migueles, Nerea, Ama,

Diego, Marcos, María, Mati, Joaquin and Eva for always wor-

rying about me despite time and distance. Thanks to my fami-

ly who supported me endless in both good and bad moments.

Very especial thanks to Tijs, Paul, Samuel, Leo and Rafael.

Your work is a great motivation and a reference to me.

Last but not least, thanks to my mum María del Pilar, to my

brother Héctor, to my godmother Elvireta and to my granny

Clotilde; this work would have been impossible without you.

Also very special thanks to Berta for being always on my side;

supporting me in all the tough moments and making me be-

lieve that this was possible even when my faith was gone.

And finally thanks to you, dear reader, for spending your time

on reading this work.

xiv

Publications

The contents of this dissertation have led to the following

journal publications and conference contributions, which are

listed below by chronological order.

Peer reviewed Journal Publication

Sánchez Cebrián, A., Zogg, M., and Ermanni P., Methodology

for optimization of the curing cycle of paste adhesives. International

Journal of Adhesion and Adhesives, 2013. 40: p. 112-119.

Sánchez Cebrián, A., Klunker, F., Zogg M., Simulation of the

Cure of Paste Adhesives by Induction Heating, first published on

May 17, 2013 as doi:10.1177/0021998313487933.

Sánchez Cebrián, A., Basler, R., Klunker, F., Zogg M., Accelera-

tion of the curing process of a paste adhesive for aerospace applica-

tions considering cure dependent void formations, International

Journal of Adhesion and Adhesives, 2014. 48: p. 51-58.

Sánchez Cebrián, A., Klunker, F., Zogg M., Modeling of void

formation during the curing process of paste adhesives, Journal of

Adhesion Science and Technology, 2014.28 (7): p. 731-747.

Conference contributions

Sánchez Cebrián, A., Lucchini, S., Zogg, M., and Ermanni, P.,

“Effect of different surface treatments on mechanical proper-

ties of CFRP bonded joints”, in 32nd SAMPE Europe Interna-

tional Technical Conference & Forum. 2011: Paris, France.

Sánchez Cebrián, A., Zogg, M., and Ermanni, P., “Indicators

for optimizing cure temperature of paste adhesives”, in 18th

International Conference on Composite Materials .Composites and

Reinforced Plastics. 2011: Jeju, KO.

http://e-citations.ethbib.ethz.ch/view/pub:59177


xv

Sánchez Cebrián, A., Moser, P., Zogg, M., and Ermanni, P.,

“Paste adhesive modification for induction curing” in SAMPE

2012. 2012: Baltimore, MD, USA.

Sánchez Cebrián, A., Basler, R., Zogg, M., and Ermanni, P.,

“Multistep heating to optimize curing process of a paste

adhesive”, in 15th European Conference on Composite Materials.

2012: Venice, IT.

Sánchez Cebrián, Zogg, M., and Ermanni, P., “Beschleunigung

der Aushärtung von pastösen Epoxiharzklebstoffen durch

variotherme Prozessführung mittles Induktion“, in FTK-

Tagung 2013 Flertigunstechnologie Kleben: Klebtechnik trifft

automobilen Leichtbau. 2013: Stuttgart, DE.

Sánchez Cebrián, Zogg, M., and Ermanni, P., “Fast and Robust

Joining Process for Aerospace Components by Local Heating

of Paste Adhesives“, in SETEC 13, SAMPE Technical Confer-

ence, 2013: Wuppertal, DE.

Colloquiums

A. Sánchez Cebrián, Fast and Robust Joining Process for Aero-

space Components by local Heating of Paste Adhesives, Fertigung-

stechnischen Kolloquium: Faserverstärkte Kunststoffe erö-

ffnen neue Möglichkeiten, November 28th, 2013, ETH Zürich,

Switzerland.

Student theses supervised

10-020- Stefano Lucchini, Research in surface treatment for cold

bonding, June 2010.

10-021- Daniel Aeschlimann, Clean Sky: Fast curing research for

bonding systems, June 2010.

10-030- Fabian Stucker, Optimization of the structural thermal

isolation for the batteries of the SunCar, June 2011.


xvi

11-031- Richard Hutter, Investigation of waterjet technologies for

surface treatment of CFRP laminates for structural bonded joints in

aerospace applications, June 2011.

11-032- Mario Bissig, Study of the robustness of an induction heat-

ing stand and influence of the convection coefficient, June 2011.

11-033- Patrick Moser, Paste adhesive optimization for the bonding

process of CFRP components with induction heating, June 2011.

12-003- Ralf Basler, Acceleration of curing process of paste adhesive

by multistep induction heating, January 2012.

12-040- Lukas Mayrhofer, Bonding an aileron for the Dornier 228

by induction heating, December 2012.

12-041- Christoph Walch, Effect of orientation of CFRP laminates

in induction heating, December 2012.

12.042- Simon Bühler, Innovative methods to optimize induction

heating in aerospace composite joints, December 2012.

xvii

Contents

Summary (English) ....................................................................... iii

Zussammenfassung (Deutsch) ..................................................... v

Resumen (Español) ...................................................................... vii

Resum (Català) .............................................................................. ix

Acknowledgements ...................................................................... xi

Publications .................................................................................. xiv

Contents ....................................................................................... xvii

List of Symbols & Abbreviations .............................................. xxi

Chapter 1 Introduction ............................................................... 3

1.1. Overview ............................................................................... 3

1.2. Clean Sky JTI ......................................................................... 7

1.3. Motivation ............................................................................. 8

1.4. Research needs ...................................................................... 8

1.5. Research objective .............................................................. 10

1.6. Approach ............................................................................. 10

1.7. Structure of the thesis ........................................................ 11

Chapter 2 State of the Art ......................................................... 13

2.1. Heating strategies in bonding technologies ................... 13

2.2. Preliminary investigations ................................................ 20

2.3. Discussion ........................................................................... 24

Chapter 3 Modeling of the Cure Kinetics ............................... 29

xviii

3.1. Introduction ........................................................................ 29

3.2. Experimental ....................................................................... 33

3.3. Cure kinetics modeling of the LME 10049-4 / LMB 6687-

2 ................................................................................................... 34

3.4. Validation of the model ..................................................... 48

3.5. Cure kinetics modeling of the LME 10625 / LME 1062651

3.6. Conclusions ......................................................................... 53

Chapter 4 Modeling of the Degradation Kinetics ................. 55

4.1. Introduction ........................................................................ 55

4.2. Mathematical model .......................................................... 57

4.3. Experimental ....................................................................... 58

4.4. Mass evaporation modeling ............................................. 60

4.5. Relation between degradation and void formation ....... 66

4.6. Degradation of the LME 10625 / LME 10626 .................. 70

4.7. Conclusions ......................................................................... 74

Chapter 5 Simulation of the curing process of paste

adhesives by induction heating ................................................... 77

5.1. Introduction ........................................................................ 77

5.2. Theoretical aspects ............................................................. 80

5.3. Modeling of the tool ........................................................... 83

5.4. Validation and discussion of the tool .............................. 85

5.5. Impact of parameters on the curing process................... 94

5.6. Conclusions ....................................................................... 101

Chapter 6 Isothermal heating process .................................. 105

6.1. Introduction ...................................................................... 105

6.2. Experimental ..................................................................... 106

xix

6.3. Physical properties of the paste adhesive ..................... 107

6.4. Mechanical properties of the paste adhesive ................ 115

6.5. Isothermal acceleration of the process ........................... 121

6.6. Conclusions ....................................................................... 125

Chapter 7 Cure dependent heating process ......................... 127

7.1. Introduction ...................................................................... 127

7.2. Approach ........................................................................... 128

7.3. Impact of the processing parameters ............................. 132

7.4. Conclusions ....................................................................... 142

Chapter 8 Process optimization ............................................. 145

8.1. Introduction ...................................................................... 145

8.2. Optimization of the LME 10049-4 / LMB 6687-2 .......... 147

8.3. Optimization of the LME 10625 / LME 10626 ............... 150

8.4. Conclusions ....................................................................... 152

Chapter 9 Design and assembly of the demonstrator ........ 155

9.1. Introduction ...................................................................... 155

9.2. Procedure for bonding complex parts ........................... 156

9.3. Description of the demonstrator .................................... 158

9.4. Assembly process ............................................................. 160

9.5. Evaluation of the aileron ................................................. 168

9.6. Conclusions ....................................................................... 172

Chapter 10 Analysis of the demonstrator .......................... 173

10.1. Introduction .................................................................... 173

10.2. Temperature distribution analysis with natural

convection. ............................................................................... 175

xx

10.3. Temperature distribution in complex shapes heated by

induction................................................................................... 182

10.4. Modification of the optimal curing cycle .................... 187

10.5. Assembly and analysis of the second demonstrator . 192

10.6. Conclusions ..................................................................... 196

Chapter 11 Concluding remarks ......................................... 199

11.1. Conclusions ..................................................................... 199

11.2. Outlook ............................................................................ 205

Appendix A Mechanical testing of composite samples ...... 211

A.1: Adherents and adhesives ............................................... 212

A.2: Production of bonded samples ...................................... 214

A.3: Single lap shear testing ................................................... 216

A.4: Fracture modes ................................................................ 217

A.5: Preparation of samples for mechanical testing ........... 221

Appendix B Induction setup ................................................. 229

B.1: Geometry and materials .................................................. 229

B.2: Calculation of the convection coefficient ...................... 235

B.3: Adaptation of the simulation tool for natural convection

conditions ................................................................................. 237

B.4: Production of samples by induction heating ............... 242

B.5: Bonding of non-electrical conductive materials by

induction................................................................................... 245

Bibliography................................................................................. 257

List of Tables ................................................................................ 269

List of Figures .............................................................................. 273

About the Author ........................................................................ 279

xxi

List of Symbols & Abbreviations

Abbreviations for Units

A Amperes

° Degrees

°C Degrees Celsius

K Degrees Kelvin

F Farad

g Grams

H Henry

Hz Hertz

h Hours

J Joule

kg Kilo

L Liters

MPa Mega Pascal

m Meters

µm Micrometers

mm Millimeters

min Minutes

mol Mole

N Newton

Ω Ohm

Pa Pascal

wt% Percent by weight

rpm Rotations per minute

s Seconds

S Siemens

T Tesla

V Volts

xxii

W Watt

Definition of Symbols

[J/g] Activation energy

[K] Ambient temperature

S [KVa] Apparent power

p [m] Coil perimeter

[A/m] Current density

α [%] Degree of cure

[kg/m3] Density of the adhesive

ρ [kg/m3] Density of the material

µ [Pa·s] Dynamic viscosity of the fluid

[-] Efficiency of the induction coil

σ [S/m] Electrical conductivity of the material

[V] Electrical voltage of induction

∆H [J/Kg] Energy generated in the curing reaction

[W/m2] Forced convection coefficient

ω [Hz] Frequency of the alternating field

[K] Glass transition temperature

[W/m] Heat generated by the magnetic fields module

[W/m] Heat generated in the chemical reaction

[W/m] Heat transferred by convection

[W/m] Heat transferred by radiation

[A] Induction Current

l [m] Length

B [T] Magnetic-B field

φ [Wb] Magnetic flux

Hf [T] Magnetic-H field

A [V·s/m] Magnetic vector potential

[-] Number of boundaries

Nu [-] Nusselt number

n [-] Order of the chemical reaction

P [W] Power

xxiii

Pr [-] Prandtl number

A [1/s] Pre-exponential factor of the cure kinetics

[-] Reaction constant

[-] Reaction constant of the chemical controlled

region

[-] Reaction constant of the diffusion controlled

part

[-] Relative permeability

[-] Relative permittivity

R [Ω] Resistance

Re [-] Reynolds number

[J/Kg·K] Specific heat capacity

υ [m/s] Speed of the fluid

[W/m2·K4] Stefan’s- Boltzmann constant

ε [-] Surface emissivity

T [K] Temperature

K [W/m·K] Thermal conductivity

R [J/K·mol] Universal gas constant

[F/m] Vacuum permittivity

[N/A2] Vacuum permeability

Abbreviations

CO2 Carbon Dioxide

CFRP Carbon Fiber Reinforced Polymers

D-MAVT Department of Mechanical and Process Engi-

neering, ETH Zürich

DSC Differential Scanning Calorimetry

DCB Double cantilever Beam

EU European Union

GFRP Glass Fiber Reinforced Polymers

IMES Institute for Mechanical Systems, ETH Zürich

ITD Integrated Technology Demonstrator

JTI Joint Technology Initiative

OoA Out of Autoclave

xxiv

SLS Single Lap Shear

SoA State of the Art

TGA Thermo Gravimetric Analysis

ETH Zürich Swiss Federal Institute of Technology Zürich

1

Part I

Heating concepts

3

Chapter 1

Introduction

1.1. Overview

The joining of structural elements is an important step in the

production of modern aircrafts. The reason is that the integrity

of the structure must always be guaranteed in a scenario with

temperature variations between -50 °C and 125 °C where high

loads are applied. In this context, several approaches to join

components have been investigated during the last decades.

These joining technologies mainly consider permanent solu-

tions that cannot be disassembled e.g. riveting, bonding or

welding.

Typically, the joining of aerospace components in commercial

airplanes has been an operation carried out by riveting [1]. The

main advantage of this technique compared to other joining

techniques, such as bonding or welding, is the possibility to

have a multiple load paths. As a consequence, if some rivets

fail, other can take the load without failure at the joint. For this

reason, this technique is still widely used today also for join-

ing structural aerospace components, as shown in Figure 1.1,

where an image of a fuselage section of the Boing 787 is

shown.

CHAPTER 1: INTRODUCTION

4

Figure 1.1: Image of the carbon fiber reinforced polymer fuselage section of

the Boeing 787 [2].

However, riveting has some drawbacks: 1) the introduction of

load peaks at the region close to the rivets, 2) the relative high

price for the high number of rivets required, and 3) the added

weight to the structure. For these reasons, the aerospace in-

dustry is researching into alternative joining technologies. One

of these joining technologies is adhesive bonding.

Bonding has some advantages compared to riveting. As thou-

sands of rivets are not required it is potentially cheaper, and

adds lower weight to the structure. Additionally, this joining

method introduces a continuous connection with a uniform

load at the joint, forming a rigid, sealed and isolated union

with a less number of components [3, 4]. This is shown in as

shown in Figure 1.2, where a comparison the load case in riv-

eting and bonding is shown.

http://upload.wikimedia.org/wikipedia/commons/2/2b/787fuselage.jpg

1.1. OVERVIEW

5

Figure 1.2: Adhesive bonding introduces a uniform load.

The main drawback to adhesive bonding is that the aerospace

requirement of multiple load paths cannot be fulfilled with

classical bondline designs. As a result, damaged bondlines

tend to peel off. Another drawbacks of this technology is that

the quality if the bondlines cannot be controlled by nonde-

structive techniques today, making them difficult to qualify.

Additionally, it requires a number of process operations that

can affect the quality of the joint; e.g. the surface treatment of

the adherents or the curing process of the adhesives. These

operations increase the energy consumption and processing

time [5, 6].

Adhesive systems can be divided with respect to its applica-

tion, having 3 main groups [7]:

Structural adhesives

Non-structural or fixing adhesives

Sealing

For aerospace applications, the bonding systems used are typ-

ically structural adhesives; systems that can join highly loaded

components for long periods of time without failure [8]. Most

commonly used materials are epoxies, cyanoacrylates, acrylics,

urethanes, silicones and thermoplastic hot melts [9]. Of these,


6

thermoset epoxies are the most commonly used type of adhe-

sive systems for aerospace applications. They are found in

single component or bi-component systems, where the hard-

ener is applied in a mixing process [8].

Structural adhesives can be found in film form or as a

paste [9]. Today, the state of the art adhesives for bonding

metals and Carbon Fiber Reinforced Polymers (CFRP) struc-

tures are epoxy based films due to the higher mechanical per-

formance and processing ease [10]. One typical application of

film adhesives is the bonding of the honeycomb core to the

facings in composite sandwich structure [11]. These systems

are typically one component thus making its application rela-

tively simple. They use a support carrier to control the mini-

mum thickness between the bonding partners what guaran-

tees a good and reproducible quality of the bonded joint [12-

15]. The drawback of this technology is related to the required

accuracy of the bondline thickness, leading high requirements

in terms of geometrical accuracy of the components to be

bonded. Additionally, these single component adhesive films

typically require refrigerated storage and a curing process that

is in most of the applications carried out in an autoclave, in-

creasing the energy consumption and the costs.

In contrast, paste adhesives can be applied in regions with

small geometry variations in the bondline thickness. The min-

imum bondline thickness is guaranteed by the use of spacers

e.g. glass beads. The adherents are accurately fixed during the

curing stage as some force is needed to keep the components

together, making the tooling typically more complex than for

film adhesives. These systems are typically used as liquid

shims in riveted structures and also in repair operations. Paste

adhesives are typically cheaper than film adhesives and do not

require a refrigerated storage. Additionally, they do not re-

quire autoclave pressure in the curing process, as it would

1.2. CLEAN SKY JTI

7

squeeze the adhesive. As a consequence, the curing process

can be applied locally at the bondline, reducing the process

energy consumption.

1.2. Clean Sky JTI

This research project is conducted in the frame of the Clean

Sky JTI. The objective of this unique public-private partner-

ship is to speed up technological breakthrough developments

and shorten the time to market for new solutions tested on full

scale aircrafts demonstrators.

The Clean Sky project is divided in one technology evaluator

and six integrated technology demonstrators (ITD):

Green regional aircraft

SMART Fixed wing Aircraft

Green rotorcraft

Sustainable and Green engines

System for Green Operation

Eco-Design

This doctoral study is part of the Eco-Design ITD. The main

objective of this ITD is to reduce the product environmental

impact considering the competitiveness of the aeronautical

industry. Its goals are set by the ACARE (Advisory Council

for Aviation Research and Innovation in Europe):

“To make substantial progress in reducing the environmental impact

of the manufacture, maintenance and disposal of aircraft and related

products.”[16]

The research in Eco-Design ITD is focused on design, produc-

tion, maintenance, and withdrawal and recycling of compo-

nents. The research activities include the development of new

material and process technologies and its demonstration in

real aircraft components. In this context, the work presented in


8

this dissertation, focuses a novel processing methodology for

paste adhesives.

1.3. Motivation

Aerospace suppliers are aiming to improve the efficiency and

reduce costs for the next generation of aircrafts. As a part of

this, adhesive suppliers are doing research into toughened

epoxy paste adhesive formulations with high mechanical

properties and low curing temperatures.

These systems aim for mechanical performance and pro-

cessing robustness comparable to film adhesive systems, yet

having the advantage that they can be easily stored and do not

require autoclave curing. Additionally, the low curing tem-

perature reduces the power consumption, making the process

more sustainable. However, the major drawback of decreasing

the curing temperature is that the paste adhesive systems re-

quire long curing processes.

In this context, the reduction of the curing time of paste adhe-

sive is desirable to deal with the increasing delivery demand

in the aerospace industry [10]. Reducing the adhesive curing

time could increase the production rate of aircrafts and reduce

the repair time, meaning benefits in production costs and en-

ergy consumption. For this reason, this research project inves-

tigates the possibility to accelerate the curing process of paste

adhesives used in aerospace composite structures.

1.4. Research needs

In order to reduce the processing time of paste adhesives,

higher curing temperatures could be used [17-19]. However,

this is not typically considered for two reasons:

The first reason is the heating methods that are typically used

today; thermal heating in forced convection ovens have a high

1.4. RESEARCH NEEDS

9

thermal inertia. As a result, they require long times and a large

amount of heat energy to increase the temperature, limiting

the temperature profiles that can be applied.

The second reason, is that the increment of the curing temper-

ature is limited by void formation, which increases when

higher temperatures are applied [20]. Voids are formed due to

the evaporation of moisture, the degradation of some chemi-

cals, and the expansion of air trapped during the mixing pro-

cess. The presence of voids leads to a decrease of the mechani-

cal performance of the bonded system [21-23]. A micrograph

of a paste adhesive with voids is shown in Figure 1.3, showing

how voids reduce the effective bonding area.

Figure 1.3: Adhesive sample with high void formation.

In this context, the investigation needs to define a strategy to

increase the curing temperature that does not affect the me-

chanical performance by production of voids. Last but not

1000 µm


10

least, this strategy must be validated with different paste ad-

hesive systems and in different geometries to ensure its re-

peatability and robustness.

1.5. Research objective

The main goal of this dissertation is to accelerate the curing

process of paste adhesives used in the bonding of aerospace

composite structures. The approach must ensure the robust-

ness of the process with respect to the different processing

parameters.

As a result, the optimal conditions to join the structural com-

ponents and the limitations of the proposed approach must be

defined.

1.6. Approach

The research contained in this dissertation defines an original

methodology aiming to answer to the following question:

How can the curing temperature of paste adhesives be increased to

accelerate the curing process, while ensuring good bondline quality

and a robust bonding process?

To answer this question, the curing cycle of an epoxy paste

adhesive is thoroughly analyzed, establishing the relation be-

tween the temperature applied, the resulting void formation

in the paste adhesive and the mechanical performance of the

adhesive bonded joint. This study is used as a baseline to set

the hypotheses of the research and to define the limits of the

process acceleration. It is based on the idea that voids are

formed at the beginning of the curing process, due to the low

viscosity of the paste adhesive. Considering the cure progres-

sion of the paste adhesive, e.g. when the gelling stage is

achieved, the temperature of the curing process can be in-

creased as the paste adhesive is more resistant to void for-

1.7. STRUCTURE OF THE THESIS

11

mation. As a consequence, a novel methodology to accelerate

the process is defined.

To validate this hypothesis, a heating method that can vary

the temperature fast and independently of the component size

is needed. As a result, the flexibility of different heating meth-

ods that could be used as an alternative to thermal heating is

investigated. Additionally, a simulation tool is modeled to

investigate the robustness of the process by analyzing the im-

pact of the process parameters to the bondline quality. Finally,

the optimal curing process obtained as a result of the pro-

posed simulation approach is experimentally validated at both

a coupon level and in a real aerospace component level to as-

sess the applicability of the research conducted.

1.7. Structure of the thesis

This doctoral thesis has 5 parts and 11 chapters, dividing the

whole research into the diverse topics considered.

In Part I (Chapters 1 and 2), the research topic is presented,

stating with the motivation, the goal and the approach of this

investigation. Additionally, the state of the art for the heating

of paste adhesives is summarized.

In Part II, the modeling and simulation of the different phe-

nomena occurring during the curing process are described.

This Part makes up Chapters 3 to 5, detailing the investiga-

tions carried out to model the cure kinetics, the degradation

kinetics and the overall simulation process of the heating pro-

cess respectively.

In Part III, the methodology for the acceleration of the curing

process of paste adhesives is defined. First, the effects of the

temperature in an isothermal curing process are investigated

in Chapter 6. Then, the effect of the degree of cure in the cur-


12

ing reaction process is investigated in Chapter 7. Finally, the

curing process is optimized in Chapter 8.

In Part IV, the methodology investigated is applied to a real

aerospace component, defined in Chapter 9. Then, this com-

ponent is evaluated in Chapter 10, defining the impact of a

complex geometry to the optimization of the curing process

previously defined.

To conclude, the main results of this research and the outlook

for possible future investigations are summarized in Part V.

13

Chapter 2

State of the Art

In this chapter, the state of the art for different heating strate-

gies used today for heating carbon fiber reinforced polymers

are compared, showing the advantages and drawbacks of each

of them.

2.1. Heating strategies in bonding technolo-

gies

Adhesive bonding of composites for aircraft application is to-

day well documented in literature [15]. One of the drawbacks

of this technique is that it requires a number of process steps

e.g. curing, surface treatment of the adherents or correct posi-

tioning of the bonding partners. The curing process is a crucial

step, because it has a direct impact on the mechanical strength

of the joint. This process is usually initiated by heat. Heating

procedures are relying on three physical principles:

Conduction

Convection

Radiation

Typically in industry, the most common techniques to cure

structural adhesives are thermal heating by forced convection

oven or autoclave curing. In the case of paste adhesives, as

pressure cannot be used because it squeezes the adhesive,

force convection ovens are today state of the art. However,

this research aims to investigate other heating strategies that

CHAPTER 2: STATE OF THE ART

14

could accelerate the process considering the environmental

impact. For this reason, the use of different heating principles

is discussed.

2.1.1. Conduction

Conduction principle transfers the energy by the vibration of

the molecules, from those more energetic to those with a lower

energy level [24]. This method is typically used by heat press-

es, heating directly the adherents, as shown in Figure 2.1. In

this scheme, the heating method is represented by red blocks,

transferring the heat at the contact areas to the adherents.

Figure 2.1: Hot press heats the adherents by conduction.

The main advantage of this method is that a very accurate

temperature control is possible. The main disadvantage of this

process is that the number of bonded assemblies and the ge-

ometry of them are limited to the geometry of the plates. This

system is difficult to move and cannot be applied directly to a

big part to be bonded. Another disadvantage is that this sys-

tem requires a long time to heat up the rather heavy plates.

2.1.2. Convection

The transfer of energy by convection is composed by two sim-

ultaneous mechanisms, a molecular diffusion and a global

movement, which can be improved by the increase of the fluid

flow [25]. This heat transfer mechanism is typically used by

Hot press

Adherents

Adhesive

2.1. HEATING STRATEGIES IN BONDING TECHNOLOGIES

15

forced convection ovens and autoclaves. By this method, the

air inside the oven is initially heated up. Then, the air transfers

the heat to the parts to be bonded. The temperature measure-

ment is carried out by thermocouples placed inside the oven,

allowing certain temperature control. A scheme of this heating

method is shown in Figure 2.2, representing a warm oven as a

red rectangle, heating the parts at the free surface areas.

Figure 2.2: Oven heats all the components

The main advantage of this method is that it can control the

temperature accurately and allows heating several parts at the

same time. Another advantage is that this method allows heat-

ing all kind of materials because the heat exchange is carried

out through the air. As a consequence, the heat exchange de-

pends only on the heat capacity of the material to be heated.

The main disadvantage of this heating method is the high

power consumption compared to other methods resulting in a

low efficiency [26]. The reason is that this method requires

heating up all the air inside the oven, the inner shell of the ov-

en and the complete component and assembly rig, requiring

time and energy. Another disadvantage of this method is the

slow flexibility on the temperature cycle applied to the part to

be heated, requiring longer time than other methods to in-

crease the temperature.

Adherents

Adhesive

Forced convection oven


16

2.1.3. Radiation

In heating strategies based on radiation such as Electron

Beams (EB), induction, lasers and microwaves, the energy

transfer is occurring through electromagnetic waves as a result

of changes in the electronic configurations of atoms and mole-

cules [27]. The transferred energy depends on the amplitude

of the electromagnetic waves to heat directly the adhesive or

the bonding partners. Typical frequencies for heating cover X-

rays [28], ϒ-rays, ultraviolet (UV) [29] and infrared (IR). Lasers

can also be used for heating applications. Performance of UV

light and EB systems is considered to be insufficient for aero-

space applications [30]. Microwaves are very suitable for heat-

ing polymers but its application on composites is still in the

research phase [31].

Induction heating is based in the exposition of an electrical or

magnetic conductive material, called susceptor, to an alternat-

ing electromagnetic field [32]. The heat is generated in the sus-

ceptor by two mechanisms, induction of Eddy currents and

Hysteresis, depending on the susceptors properties.

Induction heating by Eddy currents

If the material is electrically conductive, then it is heated by

the effect of Eddy currents, following the induction law,

shown in Eq. 2.1. It states that the induced voltage [V] de-

pends on the change of the magnetic flux [Wb] with

time [33].

2.1

The Magnetic flux through a conductive loop can be written as

in Eq. 2.2, and depends on the vacuum permeability,

[V·s/A·m], the area enclosed by the loop, [m2], and the mag-

netic field


17

2.2

The induced current [A] will depend on the resistance of

the conductive material, and can be written like Eq. 2.3

by combining Eqs. 2.1 and 2.2.

2.3

This will lead in a heat generation, [W], according to Joule’s

law, shown in Eq. 2.4.

2.4

Finally, the power dissipated in the susceptor P [W] by induc-

tion currents can be calculated as shown in Eq. 2.5.

2.5

Induction heating by Hysteresis

If the material is ferromagnetic, hysteresis effect appears [34].

When an alternating magnetic H-field is applied to a ferro-

magnetic material, it is initially magnetized until it achieves

the saturation point S, shown in Figure 2.3. Once magnetized,

it follows the curve S-R-C-S’-S. This means that if no magnetic

field is applied, a remaining magnetic field Br [T], least. The

area surrounded by the hysteresis loop is proportional to the

power generated by hysteresis associated to processes losses.


18

Figure 2.3: Hysteresis loop of a ferromagnetic susceptor.

The heating power can be calculated with the following rela-

tion:

∫

2.6

where [W] is the power generated by the hysteresis effect,

V [m3] is the volume of the work piece, f [Hz] is the frequency

and [Vs/Am] is the magnetic permeability, being the product

of the relative permeability [-] and the vacuum permeabil-

ity [Vs/Am].

In this context, the power generated can be controlled by:

Frequency of the alternating magnetic field

Permeability of the ferromagnetic material

The magnetic H-field applied.

Total susceptors volume.

Skin effect

B [T]

O [A/m]-

-

C

R

Field removal

Initial magnetization


19

This effect occurs in the heating process of an electrical con-

ductive material. It is characterized by the skin depth, also

known as penetration depth. It represents the maximum depth

that the material absorbs the 87 % of the energy induced [35].

If the conductor is thinner than this value, the heating by in-

duction is considered to be homogeneous in all the thickness.

Otherwise, a temperature gradient is created inside of the

work piece. For alternating current, the skin depth can be cal-

culated with the following relation:

√

√ 2.7

Where [m] is the skin depth and [S/m] is the electrical

conductivity of the material to be heated.

Induction heating of paste adhesives

Induction heating is today widely used in industry, principal-

ly for forging or brazing of electrical conductive materials [36,

37]. However, its applicability depends on the electromagnetic

properties of the material to be heated.

In the case of heating composites with induction, the good

electrical conductivity of carbon fibers, together with the high

fiber volume content, makes CFRP a good option to be used as

adherent in a bonding process [26, 38, 39]. By the induction of

Eddy currents, CFRP plates can be locally heated in the

bondline area, transferring the heat to the paste adhesive by

conduction, as shown in the scheme in Figure 2.4 [40, 41].


20

Figure 2.4: Induction heats the CFRP area close to the coil.

Advantages of this technology are the low power consump-

tion, the possibility to rapidly change the energy input to the

adherent and the local application of the heat generated by

induced currents [40]. Other advantages are that this heating

system permits an accurate control of the temperature applied

and it is also a portable system, allowing bonding complex

shapes and big assemblies.

The main disadvantage of this technique is it cannot be initial-

ly applied to directly heat non-conductive materials, as the

paste adhesive.

2.2. Preliminary investigations

After defining the three heat transfer methods existing, hot

press, forced convection oven and induction heating are com-

pared. The evaluation is performed according to the following

criteria.

Mechanical properties of the joint.

Heating rate generated.

Temperature distribution in the bondline.

Energy consumption.

Controllability of the process.

Adherents (Susceptor)

Adhesive

Induction heating coil

2.2. PRELIMINARY INVESTIGATIONS

21

In order to analyze the effect on the mechanical properties,

single lap shear samples (SLS) are bonded with the different

heating methods and then they are mechanically tested. De-

tails of the preparation of samples, bonding process are given

in the Appendix A.

To analyze the heating rate, the temperature generated and

the controllability of the process, a qualitative test is carried

out comparing the different techniques. This test consists on

heating a CFRP plate (200 mm x 100 mm x 1 mm) from room

temperature (20 °C) to 100 °C with full power, comparing the

performance of the different heating methods. The tempera-

ture is measured by a Thermocouple type J, excepting for in-

duction heating. In this case an IR pyrometer (Impac IN 510-N,

LumaSense Technologies, Santa Clara, CA) is used, as com-

mented in Appendix B.1. Additionally, the energy consump-

tion is calculated considering the heating stages and the cool-

ing systems used.

2.2.1.1. Effect on the mechanical properties

The samples are cured at about 100 °C for 60 minutes with the

processing parameters summarized in Table 2.1. The results of

the SLS test are summarized in Table 2.2 and Figure 2.5.


22

Table 2.1: Parameters used to cure SLS samples.

Heating

method Supplier Model Input

Oven Herareus UT 6120 100 °C

Hot press Fontijne Grotnes TP 400 100 °C

Induction Ambrell EasyHeat Distance coil-

plate= 6 mm

I = 44A freq =

262’000 Hz Table 2.2: Results of the SLS test.

Heating method

Bondline

thickness

[mm]

Shear

strength

[MPa]

Fracture

mode

Oven 0.35 ± 0.05 23.6 ± 2.4 Adherent

Hot press 0.35 ± 0.04 24.5 ± 2.9 Adherent

Induction 0.42 ± 0.01 24.4 ± 1.5 Adherent

Figure 2.5: SLS results.

0

5

10

15

20

25

30

Oven Hot press Induction

Sh

ear

Str

eng

th [

MP

a]

2.2. PRELIMINARY INVESTIGATIONS

23

Result show that all the heating methods tested cure the paste

adhesive without causing a decrease on the mechanical per-

formance.

2.2.1.2. Heating rate, temperature distribution, controllability

of the process and energy consumption

The time to heat up a CFRP plate (200 mm x 100 mm x 1 mm)

is measured for the different heating methods at full power.

Results are summarized in Table 2.3.

Table 2.3: Time necessary to heat up a CFRP plate.

Heating

method

S

[kVA]

Time

[min]

Heating

rate

[°C/min]

Power con-

sumption

[kWh]

Oven 2.2 16.7 4.79 0.55

Hot press 10 6.9 11.59 1.04

Induction 3.6 0.15 533 7.0·10-3

The heating rates differ within the methods tested in some

orders of magnitude. Induction heating can generate big heat-

ing rates in the CFRP meaning a more flexible and controllable

process than oven and hot press.

Regarding the temperature distribution, oven and hot press

can heat the entire CFRP surface by different physical princi-

ples: convection and conduction respectively. Induction is lim-

ited to the coil area, but a proper design can cover all the

bondline area. With oven and hot press, the heating is con-

ducted from the surface of the assembly. This leads to a higher

temperature gradient than with induction heating, which

heats from the interface between adhesive and adherent, lead-

ing to a smother distribution.


24

Finally, the power consumption of induction is lower than

oven and hot press. However, the use of oven curing allows

heating more than one component at the same time.

2.3. Discussion

After analyzing the performance of the three heating methods

investigated, their further use is discussed.

Table 2.4: Summary of the performance of the heating methods investigated.

Property Oven Hot

press Induction

Mechanical properties of the

joint ++ ++ ++

Heating rate generated - - ++

Temperature distribution in the

bondline ++ + +

Energy consumption - - ++

Controllability of the process - -- +

Oven heating, despite being today’s state of the art, shows

certain limitations compared to other heating methods consid-

ered in this dissertation e.g. high power consumption. Anoth-

er drawback is the low heating rates that can achieve, limiting

the flexibility of the process. The same limitations can be

found with the hot press, which additionally require parallel

samples or special mold design to be heated.

Finally, induction heating shows low power consumption and

a flexibility to generate high heating rates. The flexibility on

the temperature application makes induction a heating meth-

od with a big potential for research. The main drawback of

this method is that it is initially limited to conductive adher-

2.3. DISCUSSION

25

ents and/or adhesives. However, this concept is not of interest

in the case of bonding CFRP structures, as the heat is conduct-

ed from the adherents to the paste adhesive. For this reason,

the results of a study analyzing potential alternatives for

bonding non electrical conductive adherents are summarized

in Appendix B.5.

For these reasons, this heating method is selected for further

research in order to fulfill the goals and challenges of this dis-

sertation.


26

27

Part II

Modeling & Simulation

29

Chapter 3

Modeling of the Cure Kinetics

In this chapter, the modeling of the cure kinetics of a paste

adhesive by DSC analysis is detailed. This study defines the

relationship between temperature, time and degree of cure of

the paste adhesives under study.

A summary of the results presented in this chapter have been

published as a part of an Article in the International Journal of

Adhesion and Adhesives (see Sánchez Cebrián, A., Basler, R.,

Klunker, F., Zogg M., Acceleration of the curing process of a paste

adhesive for aerospace applications considering cure dependent void

formations, International Journal of Adhesion and Adhesives,

2014. 48: p. 51-58.)

3.1. Introduction

The modeling of the cure kinetics permits obtaining the rela-

tion between temperature, time and the degree of cure of a

chemical reaction. In the context of this research, the modeling

of the cure kinetics of the paste adhesives under study is nec-

essary to optimize the process and to evaluate samples. The

reason is that this model gives the information about the cur-

ing progression of the paste adhesive during the heating stage.

This information is used thorough this dissertation to calculate

the degree of cure of a sample for any curing cycle applied at

any point of the process.

One of the most widely used methods for the analysis of the

cure kinetics is the thermal analysis by DSC [42]. This tech-

CHAPTER 3: MODELING OF THE CURE KINETICS

30

nique permits to measure the exothermal energy during the

curing process for a defined thermal program, which can be

isothermal or dynamic. By this energy released, the degree of

cure can be measured following Eq. 3.1. This value varies from

0 in non-cured samples to 1 for samples completely cured.

∫

3.1

The degree of cure can be calculated as the relation of released

heat for a certain period of time respecting the overall heat

enthalpy of the chemical reaction.

The overall heat enthalpy is calculated by direct integration of

the heat flow. One option is using isothermal measurements at

temperatures high enough to complete the curing process.

Nevertheless, this procedure is not recommended by many

authors [43] because a part of the heat may not be detected by

the calorimeter. This can occur when the temperature is too

low and the released heat is lower than the sensitivity of the

equipment or if part of the energy is released during the stabi-

lization process of the equipment. Another reason is that if the

heating rate is too fast, the curing process may not be complet-

ed in the range of temperatures selected, leading to a wrong

measurement. For these reasons, typically a dynamic curing is

used to measure the overall heat enthalpy. One example of a

DSC measurement is shown in Figure 3.1. The chemical reac-

tion is limited between and , selected by the change of the

slope of the measurement curve.

3.1. INTRODUCTION

31

Figure 3.1: DSC curve with a heating rate of 10 °C/min.

Since the 70s, the polymerization kinetics of thermosets has

been mater of intense research. As a consequence, there are

today several models defining a relationship between temper-

ature, time and the degree of cure for thermoset resins as the

paste adhesives [44-46]. Different approaches have been pro-

posed, distinguishing principally between the phenomenolog-

ical kinetic and the mechanistic models [44, 45].

The difference is that whereas the first one is based on empiri-

cal results, the second accounts the balance of chemical species

involved in the reaction thus forming mathematical relations.

Mechanistic models can result to be very complex in the case

of the curing of paste adhesives. For this reason in this disser-

tation, the phenomenological kinetic modeling by DSC is con-

sidered [44, 45].

The phenomenological kinetic model is described by an ordi-

nary differential equation, Eq. 3.2, defining the conversion rate

as a function of degree of cure and temperature.

-2

-1

0

1

2

3

4

5

0 50 100 150 200 250 300

En

do

ther

mic

hea

t fl

ow

[m

W]

Temperature [°C]

T1 = 28 °C

T2 = 211 °C

Overall heat = 260.97 [J/g]


32

3.2

There exist several models defining this chemical reaction [46].

The most important are:

n-th order reaction:

3.3

Autocatalytic reaction:

3.4

n-th order + autocatalytic reaction (“Kamal and So-

rour” model [43].)

3.5

Parameter k is the overall rate constant having an Arrhenius

form as shown in Eq. 3.6:

(

) 3.6

These models differ mainly on the maximum curing conver-

sion rate point, different for each thermoset. Autocatalytic

chemical reactions the maximum conversion rate is at about 30

% of the chemical reaction. For the n-th order reactions the

maximum conversion rate is at the beginning of the chemical

reaction.

Using these three models as a base, there are many extended

models using functions instead of constant values as reaction

parameters [47, 48]. One of the most common extensions, con-

sidered in this dissertation, is the one considering a division of

the curing reaction in “chemical controlled” part and “diffu-

3.2. EXPERIMENTAL

33

sion controlled” part. This extension models the vitrification of

the thermoset, having the chemical controlled part until vitri-

fication occurs and then having a diffusion controlled part

[49]. This extension can be applied to any of the three main

models in order to improve the accuracy of the prediction.

In order to determine the parameters to model the chemical

reaction of the resin under study, experimental data is ob-

tained by DSC. This data is analyzed and compared to existing

models, selecting the one that fits best. Then, this data is used

by the fitting algorithm called lsqcurvefit (Matlab®) together

with the model selected, calculating the reaction parameters.

Finally, the model is validated by comparing predicted con-

version values with samples cured combining isothermal and

dynamic curing cycles.

3.2. Experimental

The equipment used in this investigation is the DSC 1 (Perkin

Elmer, USA). The samples analyzed are previously weighted

(AD-6 Microbalance, Perkin Elmer, USA) having always a

weight between 5 and 10 mg. In order to consider the effect of

instabilities of the equipment e.g. irregular flow of the gas

used for the cooling, the measurements are corrected with

baselines. The thermal program described in this dissertation,

if not indicated the contrary, has four steps as follows:

Initial step of 1 minute at 30 °C to stabilize the meas-

urement.

Heating at 50 °C /min to targeted isothermal tempera-

ture.

Isothermal heating for 60 minutes.

Final post cure, heating up until 300 °C at 10 °C/min.


34

3.3. Cure kinetics modeling of the LME

10049-4 / LMB 6687-2

3.3.1. Overall heat enthalpy measurement

The overall heat enthalpy of the paste adhesive LME 10049-4 /

LMB 6687-2 is used as reference to calculate the degree of cure

of all the validated samples.

In order to select a correct heating rate, a preliminary study is

carried out. Previous research has proved that the value of the

heating rate must be in the range between 2 °C/min and

15 °C/min to avoid similar problems as commented for iso-

thermal measurements [50]. In this context, 8 samples are

heated between 0 °C and 300 °C considering different heating

rates between 5 and 15 °C/min.

Results of all the samples are summarized in Table 3.1.

Table 3.1: Results of the overall heat enthalpy measurement.

Heating

rate

[°C/min]

[J/g] Average

Value

262.0 ± 6.9

1 2 3 4 5

5 259.3 261.0 269.0 258.5 273.0

10 261.2 252.3 - - -

15 231.2 - - - - -

The decrease of the overall measured enthalpy in the samples

cured with 15 °C/min is most likely due to that the chemical

reaction is not completed. All the other samples measured

show a similar value. For this reason, the validation of all the

samples measured in this dissertation is carried out with a

heating rate of 10 °C/min in order to save time, using a refer-

3.3. CURE KINETICS MODELING OF THE LME 10049-4 /

LMB 6687-2

35

ence overall heat of 262 J/g. This value is the average of the 7

values obtained by heating at 5 °C/min and 10 °C/min.

3.3.2. Model selection

After the overall heat is determined, the model of the cure ki-

netics is selected. As mentioned in Section 3.1, the models con-

sidered in this dissertation differ on the point of maximum

heating rate. In order to measure this point, different samples

are cured isothermally at different temperatures. Samples are

cured twice for each temperature in order to ensure repeatabil-

ity. Results of the degree of cure are conversion rate are shown

in Figure 3.2 and Figure 3.3 respectively.

Figure 3.2: DSC measurements.

Deg

ree

of

cure

[-]

Time [min]

0 20 40 60 80 100

1.0

0.9

0.8

0.7

0.6

0.5

0.4

0.3

0.2

0.1

0.0

70 °C80 ° C90 ° C100 ° C110 ° C


36

Figure 3.3: Conversion rate of the different measurements.

Figure 3.3 shows that the maximum conversion rate is ob-

served at the beginning of the curing process. This means that

the curing reaction behaves as an n-th order.

3.3.3. Modeling of the cure kinetics

The logarithmic form of the n-th order equation (Eq. 3.3) can

be rewritten as in Eq. 3.7.

(

) 3.7

As shown in Eq. 3.7, the model has a linear form, .

The reaction order n [-] is determined by calculating the slope

of the curve for each isothermal measurement. The results are

summarized in Table 3.2.

Time [min]

6

5

4

3

2

1

00 10 20 30 40 50 60

70 °C80 ° C90 ° C100 ° C110 ° C

Co

nv

ersi

on

rate

[-]

x 10-3


LMB 6687-2

37

Table 3.2: n-th order model parameters by lineal regression.

Temperature[K] n [-] ln (k) [-]

343 1.57 -7.26

353 1.71 -6.80

363 1.42 -6.19

373 1.68 -5.72

383 1.78 -5.24

As experienced by other authors [48], the parameter n is not a

constant value. For this reason an average value of n = 1.66 is

used, as proposed in [51]. Figure 3.4 shows the ln (k) as a func-

tion of the inverse isothermal temperature T.

Figure 3.4: Measurement and fitting of ln (k).

The linear behavior of the curve in Figure 3.4 is explained by

the logarithmic formulation of Eq. 3.6:

l[-

]

Temperature-1 [1/K] x 10-3

-5.0

-5.5

-6.0

-6.5

-7.0

-7.52.6 2.65 2.7 2.75 2.8 2.85 2.9 2.95 3

MeasurementFitting


38

3.8

The pre-exponential factor A [1/s] is calculated by the intersec-

tion with the y-axis. The activation energy E [J/mol], is the

slope of the line. Results of the parameters are summarized in

Table 3.3 and experimental and predicted values are com-

pared in Figure 3.5.

Table 3.3: Parameters of the n-th order model.

Parameter Value Unit

A 233’403 [1/s]

E 56’087 [J/mol]

n 1.66 [-]

Figure 3.5: Comparison of n-th order and experimental data.

The n-th order model gives higher predicted conversion val-

ues than measured. This can be explained by the phenomena

Deg

ree

of

cure

[-]

Time [min]

1.0

0.9

0.8

0.7

0.6

0.5

0.4

0.3

0.2

0.1

0.00 20 40 60 80 100

Experimental 70 °CFitting 70 °CExperimental 80 ° CFitting 80 ° CExperimental 90 ° CFitting 90 ° CExperimental 100 ° CFitting 100 ° CExperimental 110 ° CFitting 110 ° C


LMB 6687-2

39

of diffusion that occurs after a system approaches vitrification.

The effect of vitrification in the cure kinetics is that the reac-

tion rate decreases two or three orders of magnitude, explain-

ing the overestimation of the n-th order model. In order to

consider this effect, the values obtained are used as initial val-

ues for a fitting algorithm with the experimental data obtained

by DSC. This means that the new values fit more accurately to

the chemical controlled region at the beginning of the chemical

reaction, before vitrification occurs. Then, once this region is

correctly modeled, further parameters must be considered to

complete the modeling of the whole chemical reaction.

As a result, the parameters in Table 4.3 are calculated. The pa-

rameters are shown in Table 3.4 and Figure 3.6.

Table 3.4: n-th order model parameters in the chemical controlled part.


A 112’280 [1/s]

E 53’795 [J/mol]

n 1.63 [-]


40

Figure 3.6: n-th order model in the chemical controlled part.

In this case, a better fitting is obtained at the beginning of the

chemical reaction. This part of the chemical reaction is named

chemical controlled region, as can be observed in Figure 3.7.

Deg

ree

of

cure

[-]

Time [min]

1.0

0.9

0.8

0.7

0.6

0.5

0.4

0.3

0.2

0.1

0.00 20 40 60 80 100



LMB 6687-2

41

Figure 3.7: Chemical and diffusion controlled parts.

Once the parameters of the chemical controlled part are calcu-

lated, the model is extended to consider vitrification phenom-

ena [47]. The modified n-th order model considers the overall

reaction rate constant k divided in a chemical controlled [-]

and a diffusion controlled part [-], as shown in Eq. 3.9.

3.9

The chemical controlled part is defined by the parameters cal-

culated in Table 3.4. The diffusion controlled part is described

as function of the temperature T [K], the conversion [-] and

the glass transition temperature, following the William-

Landel-Ferry (WLF) equation, shown in Eq. 3.10 [47, 52, 53].

Deg

ree

of

cure

[-]

Time [min]

1.0

0.9

0.8

0.7

0.6

0.5

0.4

0.3

0.2

0.1

0.00 20 40 60 80 100 120 140 160 180

70 °C80 ° C90 ° C100 ° C110 ° C


42

(

)

|

| 3.10

To calculate all the parameters, at first [K] and

[K]

are experimentally determined and then the parameters [K]

and [1/s] are calculated. The parameter [K] is the

WLF universal rate constant, equal to 51.6 K [47, 52, 54].

To determine [K] and

[K], the relationship between

the glass transition temperature and the corresponding con-

version rate [-] is defined using the DiBenedetto equation,

shown in Eq. 3.11 [55, 56]

3.11

In a DSC-signal, the glass transition can be found as a shift in

the curve into the endothermic direction. The reason is that the

specific heat capacity Cp is increasing at the glass transition

temperature. [57]. The experimental data is fitted into the

DiBenedetto equation, as shown in Figure 3.8.

Figure 3.8: Conversion rate vs. Tg in the measurements.


LMB 6687-2

43

The parameters obtained from the fitting are summarized in

Table 3.5. These parameters are used just as reference and are

calculated in a later fitting. The reason is that the measure-

ments of the glass transition temperature are carried out by

the DSC software, instead with a temperature modulated DSC

(TM-DSC), which would give more accurate results [43].

Table 3.5: Fitted parameters at the DiBenedetto equation.


0.2348 [-]

259 [K]

398 [K]

In order to solve the WLF equation, is still unknown.

This parameter is calculated following Eq. 3.11 but with

instead of , being the conversion rate in the change

between chemical controlled and diffusion controlled part.

is calculated using data from the measurements and the

calculated parameters shown in Table 4.3. The values of

are shown in Table 3.6 and Figure 3.7.

Table 3.6: αonset at different temperatures.

Temperature[K] [-]

343 0.64

353 0.72

363 0.78

373 0.85

383 0.90

To show the dependency of with the temperature,

Eq. 3.12 is used [58].


44

(

) 3.12

is a fitting parameter and is the hypothetical temperature

below which there is no chemical reaction. The plot of

and the resulting parameters are shown in Figure 3.9 and Ta-

ble 3.7.

Figure 3.9: Fitting of αonset

Table 3.7: Fitted parameters of αonset.


273 [K]

856 [K]

Afterwards, the parameter is rewritten considering the

WLF and the DiBenedetto equations. Logarithms are also ap-

plied to the resulting equation, having the final form shown in

Eq. 3.13.

Temperature [K]

[-]

1.0

0.95

0.9

0.85

0.8

0.75

0.7

0.65

0.60

MeasurementFitting

340 350 360 370 380


LMB 6687-2

45

|

|

3.13

With this equation having linear form, it is possible to calcu-

late and , as shown in Table 3.8.

Table 3.8: Values of C1 and KDonset.

Temperature[K] [-]

[1/s]

Average value

[K]

[1/s] 343 9.3 7.510-4

3.6 0.0084

353 2.7 0.0045

363 3.2 0.0060

373 1.6 0.0110

383 1.3 0.0199

As the adhesive does not show a clear separation between

chemical and diffusion controlled part, the average values of

both parameters, and , are used. A summary of all

the parameters calculated is shown in Table 3.9.

Table 3.9: Summary of all the fitted parameters.

Parameter A E n

Value 112’280 53’795 1.63 0.2348 259 398

Unit 1/s J/mol - - K K

Parameter

Value 856 273 3.6 51.6 0.0084

Unit K K - K 1/s

With these parameters, experimental and predicted values are

compared, as shown in Figure 3.10.


46

Figure 3.10: n-th order model fitting at the chemical region.

The n-th order model considering chemical and diffusion con-

trolled regions adjusts more to the experimental data. All

these parameters are then used in by the curve fitting algo-

rithm (lsqcurvefit) as starting values. Then, the fitted parame-

ters are calculated and summarized in Table 3.10.

Deg

ree

of

cure

[-]

Time [min]

1.0

0.9

0.8

0.7

0.6

0.5

0.4

0.3

0.2

0.1

0.0

0 20 40 60 80 100



LMB 6687-2

47

Table 3.10: Summary of all the fitted parameters.

Parameter A E n

Value 112’280 53’795 1.63 0.3746 252 399


Parameter

Value 889 265 9.1 51.6 0.1677·10-3

Unit K K - K 1/s

The comparison between experimental and predicted degree

of cure is shown in Figure 3.11.

Figure 3.11: Fitting of the final model.

These values are obtained with the only limitation of the pa-

rameter , defined by literature. Finally, all process

Deg

ree

of

cure

[-]

Time [min]

1

0.9

0.8

0.7

0.6

0.5

0.4

0.3

0.2

0.1

0

0 20 40 60 80 100



48

for the modeling of the cure kinetics is summarized in Figure

3.12.

Figure 3.12: Scheme for the modeling of the cure kinetics.

3.4. Validation of the model

For the validation of the model, a multiple isothermal curing

profile is considered, shown in Figure 4.13 and Figure 3.14.

The procedure to produce and analyze the samples with the

profiles shown in Table 3.11 is explained in detail in Appendix

B.4. Results are shown in Figure 3.15.

Experimental data obtained byDSC

Cure kinetics model

Fitting of n-th order model

Fitting only in chemical controlled region

Extension of the chemical and diffusion controlled model

Fitting of obtained model parameters

n-th orderAutocatalitic

n-th order + autocatalitic

3.4. VALIDATION OF THE MODEL

49

Figure 3.13: The temperature is increased from T1 to T2.

Figure 3.14: The process will have a faster curing process.

Deg

ree

of

cure

[-]

1

0.9

0.8

0.7

0.6

0.5

0.4

0.3

0.2

0.1

0

0 10 20 30 40 50 60Time [min]

T1

T2

t1

Deg

ree

of

cure

[-]

1

0.9

0.8

0.7

0.6

0.5

0.4

0.3

0.2

0.1

00 10 20 30 40 50 60

Time [min]

80 ° C100 ° C120 ° C140 ° C

T1

T2

t1


50

Table 3.11: Samples for cure kinetics model validation.

Heating

method Temperature profile

Oven 15 min at 40 °C 0.11 0.13 Oven 50 min at 40 °C 0.30 0.30 Oven 60 min at 40 °C 0.34 0.37 Oven 100 min at 40 °C 0.47 0.45 Oven 40 min at 50 °C 0.37 0.37 Oven 30 min at 60 °C 0.44 0.43 Oven 70 min at 60 °C 0.65 0.64 DSC 23 °C to 140 °C at 10 °C/min 0.78 0.8

DSC 23 °C to 140 °C at 25 °C/min 0.57 0.61

DSC 23 °C to 60 °C at 20 °C/min, hold 10

min, 60 °C to 80 °C at 20 °C/min and

hold 10 min.

0.53 0.57

Induction 23 °C to 60 °C at 33 °C/min, hold 10

min, 60 °C to 80 °C at 19 °C/min, hold

10 min.

0.53 0.57

Induction 23 °C to 125 °C at 68 °C/min and hold

2.5 min. 0.67 0.69

Induction 23 °C to 80 °C at 22 °C/min, hold 5

min, 80 °C to 100 °C at 14 °C/min,

hold 5 min.

0.66 0.71

Induction 23 °C to 140 °C at 25 °C/min 0.77 0.80

Induction 15 min at 100 °C, 100 °C to 140 °C at

25 °C/min and hold 5 min. 0.95 0.96

Induction 23 °C to 100 °C at 5 °C/min, 100 °C to

160 °C at 5 °C/min 0.96 0.98

Induction 15 min at 100 °C, 100 °C to 160 °C at

25 °C/min and hold 5 min 0.98 0.99

Induction 12.5 min at 90 °C, 90 °C to 160 °C at

25 °C/min and hold 15 min 1 0.99

3.5. CURE KINETICS MODELING OF THE LME 10625 / LME

10626

51

Figure 3.15: Validation of the cure kinetics model.

The model shows a good accuracy on the prediction of the

degree of cure in the whole range of values tested, showing a

small error especially for a degree of cure higher than 90 %,

where the conversion rate is slower and the accuracy needed

is higher.

3.5. Cure kinetics modeling of the LME

10625 / LME 10626

The modeling of the cure kinetics is also carried out for the

paste adhesive system LME 10625 / LME 10626 from Hunts-

man Advanced Materials considering the approach described

in Figure 3.12. The rational for this investigation is that the

acceleration of the curing process of paste adhesives investi-

gated is carried out for two different paste adhesive systems to

ensure the repeatability of the research. In this context, the

cure kinetics of both systems is individually modeled.

Deg

ree

of

cure

mea

sure

d[-

]

1

0.9

0.8

0.7

0.6

0.5

0.4

0.3

0.2

0.1

0

Degree of cure modeled [-]0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

Induction HeatingForced convection ovenDSC-equipment


52

To determine the overall reaction enthalpy, 7 samples are

heated between 0 °C and 250 °C considering different heating

rates between 5 and 15 °C/min. The results are summarized in

Table 3.12.

Table 3.12: Results of the overall heat enthalpy.

Heating

rate

[°C/min]

[J/g]

1 2 3 4 5 Average

Value

322.75 ± 4.4

-

5 324.8

10 325.5 326.6 318.1 316.9 327.7

15 319.6

To determine the fitting parameters, 8 samples are analyzed

by DSC curing them at 4 different temperatures from 80 °C to

140 °C for 1 hour. The results are summarized in Table 3.13.

Table 3.13: Cure kinetics parameters of the LME 10625 / LME 10626.

Parameter A E n

Value 451’711 56’862 0.80 15.32 0.0196 7’163


Parameter

Value 244 149 6.42 51.6 0.958·10-5

Unit K K - K 1/s

Finally, a comparison between experimental results by DSC

and the model with the fitted parameters is carried out to val-

idate the model, shown in Figure 3.16.

3.6. CONCLUSIONS

53

Figure 3.16: Model validation for the cure kinetics model.

The model shows a good accuracy, having a relative error

lower than 1 % for all the temperatures measured.

3.6. Conclusions

In this chapter, the cure kinetics of the paste adhesive LME

10049-4 / LMB 6687-2 is modeled and validated. Well-known

modeling techniques are used as a reference to select a proper

model and obtain accurate results. An n-th order model is

used as base kinetic model, extending it to consider the vitrifi-

cation process. The model is validated with a dual step heat-

ing process, getting an accurate prediction of the degree of

cure. Finally, the same approach is applied to the paste adhe-

sive LME 10625 / LME 10626, defining as well an accurate

model to predict the cure kinetics.

0,0

0,1

0,2

0,3

0,4

0,5

0,6

0,7

0,8

0,9

1,0

0 5 10 15 20 25 30

Deg

ree

of

cure

[-]

Time [min]

Experimental 80 °C

Model 80 °C

Experimental 100 °C

Model 100 °C


Model 120 °C


Model 140 °C


54

55

Chapter 4

Modeling of the Degradation Ki-

netics

In this chapter, the modeling of the void formation of paste

adhesives is detailed. This study defines the relation between

void content, degree of cure, temperature and time for the

paste adhesives under study.

The results presented in this chapter have been published as

an Article in the Journal of Adhesion Science and Technology (see

Sánchez Cebrián, A., Klunker, F., Zogg M., Modeling of void

formation during the curing process of paste adhesives, accepted at

the Journal of Adhesion Science and Technology on Novem-

ber 19.)

4.1. Introduction

The increment of the curing temperature is limited by the

thermal degradation of the paste adhesive due to void for-

mation as higher temperatures are applied [20, 26, 59, 60].

Voids are formed due to the evaporation of moisture, the deg-

radation of chemical components and the expansion of air

trapped during the mixing process, leading to a decrease of

the mechanical performance for the bonded system [21-23].

Maximum void content values can be found in literature for

composite panels and limitations about positioning of voids in

the edge in order to avoid delamination problems [61] How-

CHAPTER 4: MODELING OF THE DEGRADATION KINETICS

56

ever, there are no explicit considerations for bonding systems

[62]. The detection of void content in paste adhesives is lim-

ited by the minimum defect size detected by NDI techniques.

If the voids are small enough they may not be detected by NDI

[20]. Today, no reliable NDI method exists to assess quality of

a bonded joint. 2 % void content is generally accepted as max-

imum value in aerospace [63, 64]. In this context, the modeling

of the void formation in relation the temperature applied

would allow to optimize the process with regard to the curing

time, assessing the impact on the mechanical performance.

Previous research has been carried out to model the thermal

degradation kinetics in thermosets due to the widespread use

of these materials in engineering. Current models define the

thermal degradation through the mass evaporation of fully

cured samples heated at high temperatures under isothermal

and non-isothermal conditions [65, 66]. These models show

that epoxy resins have a complex degradation process, over-

lapping the degradation of different components in a wide

temperature range [67, 68]. The models are mainly used to

observe the application limits of such materials and to validate

the performance of additives at high temperature [58, 69-71].

One of the hypotheses of this doctoral dissertation states that

the void formation is not only dependent on the curing tem-

perature and time, but also on the degree of cure of the adhe-

sive. Paste adhesives are more resistant to higher curing tem-

peratures when the degree of cure is higher, especially after

the solidification in the gelation stage [72]. In this chapter, a

novel degradation kinetics model considering the evaporation

of the paste adhesive is defined. This model is based on exper-

imental data obtained with TGA by isothermal curing cycles at

different temperatures. Then, the model is validated compar-

ing the evaporated mass measured by TGA to predicted val-

ues.

4.2. MATHEMATICAL MODEL

57

To complete the model, the relation between void generation

and evaporated mass during the heating process of a paste

adhesive is experimentally established. The final tool predicts

the void content of a paste adhesive produced during the cur-

ing cycle and thus assessing the quality of the bonded joint.

Finally, the repeatability of the research findings is proved by

the modeling of the void formation of the paste adhesive LME

10625 / LME 10626.

4.2. Mathematical model

The degradation kinetic model considered in this investigation

is based on three hypotheses:

The evaporation of the paste adhesive depends on the

curing temperature.

The evaporation of the paste adhesive depends on the

degree of cure of the paste adhesive [-], described in

Chapter 3: Modeling of the Cure Kinetics.

The paste adhesive evaporates, even after the curing

process is completed, if temperatures higher than

[K] are applied.

In order to consider these hypotheses, the model is based on a

multistep decomposition model, considering the evaporation

of components at different temperatures [65, 66]. In this case,

the mass reduction of the paste adhesive is taken into account

in two independent processes occurring at different tempera-

ture ranges: one starting at low temperature and the other oc-

curring at temperatures higher than [K] of the completely

cured system. The model is described as follows:

4.1


58

where

[-] is the evaporation rate of the paste adhesive,

and are kinetic constants associated with the

thermal degradation of the adhesive. Additionally,

are fitting parameters associated to the degra-

dation order of the components and [– ] is a fitting parame-

ter associated to the curing reaction rate order. Finally, [-] is

the relative paste adhesive mass and [-]is the relative

mass remaining at the end of the process, which is a function

of the temperature and modeled by following linear relation-

ship with the fitting parameter [1/K] as a constant:

4.2

The kinetic constants [– ] are based on an Arrhenius Equa-

tion [65]. In this case, as there are two evaporation processes

modeled, two kinetic constants are used:

(

) { } 4.3

where [1/s] are the pre-exponential factors and [J/(mol)]

are the activation energies, both of them fitting parameters.

4.3. Experimental

The model of the thermal degradation is firstly applied to the

paste adhesive LME 10049-4 / LMB 6687-2 from Huntsman

Advanced Materials. TGA curves are used to fit the parame-

ters of the theoretical model: Different isothermal curing cy-

cles with temperatures between 60 °C and 160 °C in steps of

20 °C are applied. The initial ramp from 30 °C to the desired

isothermal temperature is conducted with a heating rate of 25

°C/min. Higher isothermal temperatures are not considered

because they cause an excessive evaporation of the adhesive

and should not be used in the curing process. The TGA curves

4.3. EXPERIMENTAL

59

are obtained with a Perkin Elmer TGA-1 using a nitrogen flow

rate of 50 ml/min, applying the desired curing cycle, which are

previously programmed. The initial mass of the samples is in

the rage of 20 ± 5 mg, measured with the TGA (having a preci-

sion of 0.0001%, a sensitivity of 0.1 μg and accuracy better than

0.02%). All the samples are performed twice, showing good

repeatability with a maximum relative error of 4% between

samples heated with the same cycle. No effect from the exo-

thermal reaction was noticed during the measurements: the

temperature difference between the sample and the program

at the dwell was always lower than the furnace’s temperature

precision. However, in order to consider this effect, the input

data used for the fitting corresponds to the sample tempera-

ture.

Each run is performed for 30 minutes after the initial heating

because the goal is to accelerate the overall curing process to

less than 30 minutes. Additionally for each curve, two base-

lines are subtracted from the original measurement:

An isothermal baseline at 40 °C, corrected with an

empty sample pan baseline. This is used to reference

the measurements with a curing cycle that influences

the weight measurement but does not generate voids.

The second baseline is carried out with an empty

sample pan to consider the gas circulation inside of

the TGA when the temperature is changed.

In order to identify the final [°C] of the cured sample to

complete the theoretical model Eq. 4.1, Dynamic Mechanical

Analysis (DMA) is conducted, using a Perkin Elmer DMA-7.

The sample used is previously cured at 80 °C for 4 hours in the

oven, following the recommendations from the supplier. The

experiment carried out is a 3 point bending test, performed

within a temperature range from 0 °C to 250 °C with a heating

rate of 3 °C/min, applying a nitrogen flow rate of 40 ml/min.


60

The samples used have a size of 18 mm x 8 mm x 2 mm. As the

sample is previously cured, it should not be affected by fur-

ther void formation. The force applied was 110 mN of preload

with a sinusoidal dynamic force of ±100 mN at 1 Hz. The point

used to measure the [°C] is the peak of tan (δ) curve, located

at 100.0 °C.

4.4. Mass evaporation modeling

4.4.1. Determination of model parameters

TGA data is fitted considering the model described in Sec-

tion 5.2, using the “least squares” algorithm from Matlab. This

procedure finds the coefficients that best fit the nonlinear

function to the experimental data. The parameters to model

the thermal degradation kinetics of the paste adhesive are

listed in Table 4.1.

Table 4.1: Thermal degradation model parameters.

Parameter Value Parameter Value

61’419 [1/s] 0.52 [-]

59’990 [J/mol] 3.19 [-]

93’719 [1/s] 2.87·10-5 [1/K]

64’932 [J/mol] 897.63 [-]

The comparison between experimental TGA measurements

and fitting is shown in Figure 4.1.

4.4. MASS EVAPORATION MODELING

61

Figure 4.1: Fitting of the thermal degradation model.

The experimental data and the model data show a good

agreement and confirm the hypotheses 1 and 3, stated in Sec-

tion 2. The model shows an increase of the evaporated mass

with a higher temperature applied during the curing process.

Additionally, it is shown that curing with a temperature high-

er than [°C] produced an additional evaporation of the

paste adhesive despite the curing process is completed.

In the lower temperature range, a significant difference be-

tween the model and the measured evaporated mass is ob-

served. The biggest difference is found at 60 °C and in the first

10 minutes at 80 °C, where the relative error is around 30%.

However, this model is designed to be more accurate at tem-

peratures higher than 100 °C. In this region there is a good

agreement between experimental and predicted values. The

maximum relative error of 7% is occurring in the curve of 140

°C. The reason to prefer a better agreement of the model at

higher temperatures and not at low temperatures is that this

0.994

0.995

0.996

0.997

0.998

0.999

1

0 5 10 15 20 25 30

Rel

ativ

e m

ass

[-]

Time [min]

TGA 60 °C

Model 60 °C

TGA 80 °C

Model 80 °C

TGA 100 °C

Model 100 °C

TGA 120 °C

Model 120 °C

TGA 140 °C

Model 140 °C

TGA 160 °C

Model 160 °C


62

model aims to accelerate the process to less than 30 minutes

and therefore temperatures higher than 100 °C are needed.

4.4.2. Validation of the model

The model is validated for different curing cycles shown in

Table 4.2, applied to the paste adhesive by comparing the pre-

dicted and measured evaporated mass.

Table 4.2: Summary of samples used for validation.

Sample

[°C]

[min]

[min]

[°C]

[min]

[min]

1 100 2.8 10 160 2.4 15

2 100 2.8 15 160 2.4 10

3 120 3.6 15 160 1.6 5

4 100 2.8 5 140 1.6 15

5 100 2.8 10 160 2.4 5

6 100 2.8 10 180 3.2 5

7 100 2.8 5 160 2.4 5

8 130 4 15 - - -

9 Heating ramp at 25 °C/min from 30 °C to 190 °C

The comparison between experimental and the predicted

evaporated mass for all the cycles considered are shown in

Figure 4.2 to Figure 4.4.


63

Figure 4.2: Validation of the heating cycles 1 to 3.


0.995

0.9955

0.996

0.9965

0.997

0.9975

0.998

0.9985

0.999

0.9995

1

0 5 10 15 20 25 30

Rel

ativ

e m

ass

[-]

Time [min]

Sample 1: 100.15.160.10 TGA

Sample 1: 100.15.160.10 Model

Sample 2: 100.10.160.15 TGA

Sample 2: 100.10.160.15 Model

Sample 3: 120.15.160.5 TGA

Sample 3: 120.15.160.5 Model

0.9955

0.996

0.9965

0.997

0.9975

0.998

0.9985

0.999

0.9995

1

0 5 10 15 20 25

Rel

ativ

e m

ass

[-]

Time [min]

Sample 4: 100.5.140.15 TGA

Sample 4: 100.5.140.15 Model

Sample 5: 100.10.160.5 TGA

Sample 5: 100.10.160.5 Model

Sample 6: 100.10.180.5 TGA

Sample 6: 100.10.180.5 Model


64


In general, the samples show a good agreement between

measurement and prediction, always with relative errors low-

er than 10%. As stated in hypothesis number 1, the higher the

temperature applied in the curing process, the higher the mass

reduction rate. This fact can be observed in Figure 4.2 and

Figure 4.4 for different initial curing process or in Figure 4.3,

for different second heating processes. All the samples show a

steeper slope when applying a higher curing temperature.

Additionally, it is shown that the paste adhesive is more sensi-

tive to thermal degradation at earlier stages of the curing pro-

cess, as stated in the model hypothesis number 2. This fact can

be observed by the logarithmic decrease on the evaporation

curve when the paste adhesive reaches a higher degree of

cure. All samples show a steeper slope on the evaporation

curve at earlier curing stages, e.g. if comparing samples 1 and

2 in Figure 4.2. Sample 2 increases the temperature earlier than

sample 1, thus achieving a higher degree of cure at this point

0.996

0.9965

0.997

0.9975

0.998

0.9985

0.999

0.9995

1

0 2 4 6 8 10 12 14

Rel

ativ

e m

ass

[-]

Time [min]

Sample 7: 100.5.160.5 TGA

Sample 7: 100.5.160.5 Model

Sample 8: 130.15 TGA

Sample 8: 130.15 Model

Sample 9: Ramp 30-190 °C TGA

Sample 9: Ramp 30-190 °C Model


65

than sample 1. Therefore, at the end of the curve, sample 2

shows a less pronounced evaporation slope on the curve.

Additionally, it can also be observed that after the curing pro-

cess is completed, typically considered when α > 95%, the

paste adhesive continues evaporating. This fact confirms the

hypothesis number 3 of this investigation.

In Figure 4.5, the evolution of the degree of cure and the evap-

oration of the paste adhesive as a function of time are shown

for sample 1.

Figure 4.5: Modeling of samples 1.

Finally, the values of the relative evaporated mass after the

first and second isothermal are summarized in Figure 4.6. It

shows the amount of measured and predicted evaporated

mass ( ) for the samples under study after completing the

first isothermal heating (marked with X) and the heating pro-

cess is completed (marked with dots).

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.995

0.9955

0.996

0.9965

0.997

0.9975

0.998

0.9985

0.999

0.9995

1

0 5 10 15 20 25 30

Deg

ree

of

cure

[-]

Rel

ativ

e m

ass

[-]

Time [min]

Relative mass TGA [-]

Predicted rel. Mass [-]

Degree of cure [-]

Degree of cure = 0.95


66

Figure 4.6: Experimental and predicted degradation.

The model presented in this Chapter predicts with good accu-

racy the evaporation of the paste adhesive during the curing

process. The maximum relative error between the experi-

mental and predicted relative evaporated mass is about 10 %

after the first isothermal step and lower than 5 % after the sec-

ond isothermal step.

4.5. Relation between degradation and void

formation

The relation between void formation and evaporation of the

paste adhesive is established by comparing the void content to

the predicted evaporated mass of samples cured with different

cycles by induction heating. The procedure to produce the

samples and to analyze them is detailed in Appendix B.4. Re-

sults of the curing processes considered are summarized in

Table 4.3.

0.002

0.0025

0.003

0.0035

0.004

0.0045

0.002 0.0025 0.003 0.0035 0.004 0.0045

Rel

ativ

e ev

apo

rate

d m

ass

mo

del

[-]

Relative evaporated mass TGA [-]

4.5. RELATION BETWEEN DEGRADATION AND VOID

FORMATION

67

Table 4.3: Void formation with different curing cycles. 𝐓

[

°C]

𝐭 [

min

] 𝐓

[

°C]

𝐭 [

min

] α

[%

] 𝐕

𝐜 [

%]

𝐓 [

°C]

𝐭 [

min

] 𝐓

[

°C]

𝐭 [

min

] α

[%

] 𝐕

𝐜 [

%]

90

12.5

16

0 5

97.2

2.

4 10

0 15

16

0 5

97.5

2.

9

90

15

160

5 97

.3

2.2

100

15

180

5 98

.9

3.1

90

17.5

16

0 5

97.3

2.

0 10

0 17

.5

160

5 97

.6

2.4

90

20

160

5 97

.4

1.7

100

20

140

10

97.3

1.

2

100

5 18

0 5

98.9

6.

8 10

0 20

14

0 15

98

.2

1.2

100

7.5

160

10

98.7

5.

3 10

0 20

16

0 2.

5 96

.0

1.4

100

10

160

10

98.8

3.

8 12

0 5

140

10

96.9

4.

4

100

10

160

15

99.3

4.

2 12

0 5

160

10

98.8

4.

8

100

10

180

5 98

.9

4.0

120

5 18

0 5

98.9

8.

0

100

15

120

15

95.0

1.

2 12

0 10

14

0 5

95.7

4.

3

100

15

120

20

96.1

1.

3 12

0 10

16

0 10

98

.9

4.1

100

15

130

10

95.4

1.

4 12

0 10

18

0 5

99.0

5.

4

100

15

130

15

96.9

1.

3 12

0 15

14

0 5

96.6

3.

9

100

15

140

10

97.0

1.

4 12

0 15

16

0 5

98.2

3.

7

100

15

140

15

98.0

1.

5 12

0 15

18

0 5

99.1

4.

4

100

15

150

5 96

.3

2.0*

12

0 20

16

0 2.

5 97

.6

3.5

100

15

150

10

98.1

2.

1*

120

60

98.6

3.

9

100

15

150

15

98.8

2.

1*

140

30

99.1

6.

1

100

15

160

2.5

95.6

2.

6 16

0 10

18

0 5

99.5

13

.4


68

According to the model, after completing the curing process

(α ≥ 95 %) the paste adhesive continues evaporating. Howev-

er, this does not necessarily lead to further void formation be-

cause the solidification process is completed. This fact is

proved by the samples in Table 4.3 marked with an asterisk

next to the void content value. They show constant void for-

mations despite the samples are cured for a longer time at the

same temperature. Under this consideration, all the samples

must be compared having the same degree of cure in the

model for the evaporated mass. Therefore, the model is used

to calculate the relative evaporated mass of the different cur-

ing cycles at the point where the degree of cure is 95 %, shown

in Figure 4.7.

Figure 4.7: Void content vs. relative evaporated mass.

As expected, the results show an increasing void content with

higher evaporated mass. Nevertheless, the results show cer-

tain evaporated mass for all the samples investigated. This fact

can be explained by the evaporation of some solvents and

moisture, which does not necessary lead to a degradation of

the adhesive and to a generation of voids. One example can be

observed, for instance, on the recommended curing cycle. In

0

2

4

6

8

10

12

14

16

0.0031 0.0032 0.0033 0.0034 0.0035 0.0036 0.0037 0.0038 0.0039

Vo

id c

on

ten

t [%

]

Rel. evaporated mass calculated for α = 95 % [-]

4.5. RELATION BETWEEN DEGRADATION AND VOID

FORMATION

69

this case, it should not lead to thermal degradation because

the curing process is carried out always under the final [°C]

of the adhesive. To complete the void formation model, a

trend line is added. This regression is selected by comparing

several regression models and considering the one with higher

R-squared value:

4.4

This relation, with an R-squared value of 91.6 %, is used to

complete the evaporation kinetics model. Finally, the compari-

son between modeled and experimental void formation is

summarized in Figure 4.8.

Figure 4.8: Summary of the model accuracy.

The final model, predicts the void content of the paste adhe-

sive under study with good accuracy especially for the curing

cycles under consideration with low void content.

0

2

4

6

8

10

12

14

0 2 4 6 8 10 12 14

Vo

id c

on

ten

t m

easu

red

[%

]

Void content predicted [%]


70

4.6. Degradation of the LME 10625 / LME

10626

The modeling of the void formation is repeated with the paste

adhesive system LME 10625 / LME 10626 from Huntsman Ad-

vanced Materials in order to show the repeatability of the in-

vestigation carried out. In this case, TGA curves are used to fit

the parameters of the theoretical model with temperatures

between 80 °C and 140 °C in steps of 20 °C. The fitting param-

eters obtained are listed in Table 4.4.

Table 4.4: Fitting parameters of the thermal degradation model of the LME

10625 / LME 10626.

Parameter Value Parameter Value

A1 45’147 [1/s] p 4.68 [-]

E1 76’464 [J/mol] n 3.51 [-]

A2 743’709 [1/s] b 6.8·10-3 [1/K]

E2 97’500 [J/mol] q 121.13 [-]

Figure 4.9 shows the variation of experimental and modeled

mass fractions.

0.992

0.993

0.994

0.995

0.996

0.997

0.998

0.999

1

0 5 10 15 20 25 30

Rel

ativ

e m

ass

[-]

Time [min]

Model 80 °C

TGA 80 °C

Model 100 °C

TGA 100 °C

Model 120 °C

TGA 120 °C

Model 140 °C

TGA 140 °C

4.6. DEGRADATION OF THE LME 10625 / LME 10626

71

Figure 4.9: Experimental and modeled evaporation curves.

The predicted values show a good accuracy with the TGA da-

ta, having even a better accuracy than for the LME 10049-4 /

LMB 6687-2.

For the establishment of the relation between evaporated mass

and void formation, samples are produced by induction heat-

ing. In this case, the curing cycles applied have been adapted

to achieve a final degree of cure of exactly 95 %. The curing

conditions of the samples, average measured void content and

modeled evaporated mass are shown in Table 4.5.


72

Table 4.5: Void formation with different curing cycles.

Tem

per

atu

re 1

[°C

]

Tim

e 1

[min

]

Tem

per

atu

re 2

[°C

]

Tim

e 2

[min

] V

oid

co

nte

nt

[-]

Ev

apo

rate

d m

ass

mo

del

[-]

80

15

140

2.73

1.

66 ±

0.2

9 0.

0020

8

80

20

100

10.0

7 0.

83 ±

0.2

1 0.

0020

0

80

20

120

4.13

1.

26 ±

0.1

2 0.

0020

1

80

20

140

1.86

1.

37 ±

0.3

3 0.

0020

4

100

5 12

0 6.

08

3.38

± 0

.05

0.00

337

10

5 14

0 3.

18

3.59

± 0

.14

0.00

343

100

5 16

0 1.

6 3.

89 ±

0.1

9 0.

0035

2

100

10

120

3.28

2.

95 ±

0.0

6 0.

0033

4

100

10

140

1.33

3.

08 ±

0.2

4 0.

0033

6

100

10

160

0.25

3.

19 ±

0.3

3 0.

0033

8

100

15

120

0.9

2.76

± 0

.17

0.00

334

120

5 14

0 1.

57

9.08

± 1

.47

0.00

468

120

5 16

0 0.

42

9.76

± 0

.99

0.00

470

4.6. DEGRADATION OF THE LME 10625 / LME 10626

73

The relation between void formation and relative evaporated

mass is shown in Figure 4.10.

Figure 4.10: Relation between voids and evaporated mass.

A trend line is added in order to obtain the relation between

evaporated relative mass and void formation:

4.5

In this case, the R-squared value is 91.3 %., slightly lower than

for the other paste adhesive under study but still over 90 %.

This relation is used to complete the evaporation kinetics

model. Finally, the comparison between modeled and experi-

mental void formation of the samples tested is summarized in

Figure 4.11.

0

2

4

6

8

10

12

14

16

18

20

0 0.001 0.002 0.003 0.004 0.005 0.006 0.007

Vo

id c

on

ten

t [%

]



74

Figure 4.11: Modeled and experimental void formation.

The void formation model of this paste adhesive system

shows a good accuracy with the experimental data. It predicts

with accuracy the void content generated for any curing pro-

cess.

4.7. Conclusions

In this chapter, the modeling of the mass reduction of a paste

adhesive during the curing process is studied considering the

following hypotheses:

The evaporated mass increases when the curing tem-

perature is increased.

At the same curing temperature the evaporated mass

rate is higher at earlier stages of the curing process.

The paste adhesive further evaporates mass, even af-

ter the curing process is completed, if temperatures

higher than [°C] are applied.

0

1

2

3

4

5

6

7

8

9

10

0.00 2.00 4.00 6.00 8.00 10.00

Vo

id c

on

ten

t m

easu

red

[%

]

Void content predicted [%]

4.7. CONCLUSIONS

75

The theoretical model stated in section 4.2 is defined fitting the

model parameters to experimental data obtained by TGA-

measurements. The model is validated by comparing TGA

data obtained combining curing cycles with dwell at two dif-

ferent temperatures. The results show a good agreement with

the predicted values for all the curing cycles analyzed and

proving all the hypotheses stated in this investigation.

Additionally, the relation between evaporated mass of a paste

adhesive and void formation is determined. In this context,

this relation is used to complete the modeling tool to predict

the void formation for a certain curing cycle applied and thus

assessing the quality of the bonded joint. Finally, the repeata-

bility of the process is proved by modeling the thermal degra-

dation of the paste adhesive LME 10625 / LME 10626.


76

77

Chapter 5

Simulation of the curing process

of paste adhesives by induction

heating

In this chapter, the simulation tool to predict the curing pro-

cess of CFRP bonded systems by induction heating is de-

scribed. The complete model is validated comparing the de-

gree of cure and temperature under different conditions. Fi-

nally, a sensitivity analysis is carried out in order to analyze

the impact of different properties on the curing process.


an Article in the Journal of Composite Materials (see Sánchez

Cebrián, A., Klunker, F., Zogg M., Simulation of the Cure of

Paste Adhesives by Induction Heating, first published on May 17,

2013 as doi:10.1177/0021998313487933.)

5.1. Introduction

Several simulation models have been developed for induction

applications, particularly for optimizing the geometry of the

coil [36, 73, 74]. Additionally, work has been carried out in

sensitivity analysis of different parameters e.g. frequency of

the current applied [75] or parameters of the materials used as

susceptors [76]. Regarding the curing kinetics of paste adhe-

sives, most common models existing today are further de-

CHAPTER 5: SIMULATION OF THE CURING PROCESS OF PASTE

ADHESIVES BY INDUCTION HEATING

78

scribed in Chapter 3: Modeling of the Cure Kinetics. For the

thermal degradation model, the models existing today for the

analysis of fully cured samples are defined in Chapter 4: Mod-

eling of the Degradation Kinetics.

In this chapter, the modeling of induction heating, the curing

and the degradation kinetics of the adhesive system LME

10049-4 / LMB 6687-2 are combined in one single simulation

tool. This allows simulating the degree of cure and void for-

mation in a bonded joint directly depending on the assembly

configuration, the applied electrical current and time, as

shown in Figure 5.1.

Figure 5.1: Scheme of the induction setup modeling.

To perform the simulation, Comsol® Multiphysics is used as a

platform. Considering different input parameters e.g. physical

and material properties, tool geometry or design of the joint; a

combination of four different physical aspects was used for

the simulation:

Electrical current applied to the coil I(t)

Frequency applied to the coilfreq(t)

Geometry of the joining tool

Joint design

Joining tool material properties

Adherent material properties

Adhesive material properties

Cure kinetic parameters

Induction heating

Heat transfer

Paste adhesive cure kinetics Degree of cure of the

paste adhesive α [-]

Inputs Output

Thermal degradation kinetics

Void content of the paste adhesive [-]

Thermal degradation parameters

5.1. INTRODUCTION

79

Magnetic fields generated by the induction process:

They generate the heat in the area of electrical conduc-

tive materials.

Simulation of the heat transfer between the materials:

Once induction generates heat in the electrical con-

ductive materials, it is transferred to other parts of the

setup as well.

Modeling of the cure kinetics of the adhesive: Due to

the transferred heat, the paste adhesive cures accord-

ing to its cure kinetics. The additional exothermal en-

ergy due to the reaction is also considered.

Modeling of the thermal degradation kinetics of the

adhesive: The tool predicts the void formation pro-

duced during the curing process.

The simulation model is validated against experimental data

in the following three steps:

Thermal simulation: Comparison of temperature evo-

lution between simulation and experiments.

Cure kinetics: The evolution of the degree of cure of

samples heated with different temperatures is com-

pared to the results of an experimental DSC analysis.

Measurement of the degree of cure: The degree of cure

of the adhesive used in the bonding process of CFRP

structures under different processing conditions is

measured by DSC and then compared to the values

predicted by the simulation.

The results of the validation are discussed, analyzing the accu-

racy of the model. Then, it is investigated which material pa-

rameters and physical effects significantly influence the degree

of cure of the paste adhesive. Additionally, the potential ap-

plications of this tool and its use to optimize the bonding pro-

cess of paste adhesives are discussed.



80

5.2. Theoretical aspects

The simulation tool is designed with Comsol® Multiphysics,

combining physical relations in the four modules considered:

“magnetic fields”, “heat transfer” and two “PDE” (Partial Dif-

ferential Equations) modules for the cure and degradation ki-

netics.

5.2.1. Magnetic fields

The governing equation is the Ampere’s Law, which relates

the integrated magnetic field around the coil to the electric

current passing orthogonally through it [77]. Therefore Eq. 5.1

is used in all the materials considering a 2 dimensional scenar-

io:

5.1

Where [A/m] is the electric current density with a frequen-

cy [Hz] that generates the alternating magnetic B-field [T].

[-] and [-] are the vacuum and relative electric permittivi-

ty, [-] and [-] are the vacuum and relative magnetic per-

meability. σ [S/m] is the electrical conductivity of the material.

Finally, [V·s·m-1] is the magnetic vector potential. The rela-

tion between the magnetic-B field, [T], and magnetic-H field

[T], is defined in Eq. 5.2 [78]:

5.2

The boundary conditions applied to the model consider

continuity for internal boundaries, the electromagnetic isola-

tion in the geometry frames to limit the model (Eq. 5.3) and

the current source applied (Eq. 5.4):

5.2. THEORETICAL ASPECTS

81

5.3

5.4

Where A [V·s/m] is the magnetic vector potential. The current

density, [A/m], is shown in the Eq. 5.5:

5.5

Where I [A] is the current applied and P [m] and [-] the

perimeter and the efficiency of the coil respectively.

5.2.2. Heat transfer

The exchange of the energy generated by Joule effect due to

the induction of Eddy currents is modeled in the “heat trans-

fer” module by applying the thermodynamic laws for heat

exchange. The first thermodynamic law states that the energy

of an isolated system is conserved [79]. In this case, due to heat

generated by the magnetic fields, there is a heat exchange be-

tween the elements in the model to reach the thermal equilib-

rium, modeled by the Eq. 5.6 [80]:

5.6

Where [Kg/m3] is the density, [J/kg·K] is the heat capacity,

[W/m·K] is the thermal conductivity. Eq. 5.6 is applied to all

materials of the setup. The term [W/ m3] models the energy

generated by the exothermal of the chemical reaction of the

paste adhesive and it is only applied to the paste adhesive

domain. The term [W/ m3] models the heat generated by

the magnetic field, which is applied only to the composite. To

model the interaction between the solids and the surrounding

air, the boundary conditions existing in the thermodynamic



82

model are thermal isolation, forced convection and radiation,

following the Eqs. 5.7 to 5.9.

Thermal isolation:

5.7

Forced convection heat transfer:

5.8

Radiation heat transfer:

5.9

Where [W/m] and [W/m] are the heat transferred by

convection and radiation respectively. [W/m2·K] is the con-

vection coefficient, [K] is the ambient temperature, [-] is

the material emissivity and [W/m2·K4] the Stefan’s- Boltz-

mann constant 5.67·10-8 W /m2·K-4 [81].

5.2.3. Paste adhesive curing kinetics

The curing reaction kinetics is modeled by a Partial Differen-

tial Equation (PDE) module, considering the energy generated

due to the exothermal behavior of the paste adhesive. The

modified n-th order reaction model is used to describe the re-

action of the adhesive is further described in Chapter 3: Model-

ing of the Cure Kinetics:

Finally, to model the energy generated by the exothermal of

the chemical reaction of the paste adhesive, [W/m], the fol-

lowing Eq. 5.10 has been used [82]:

5.10

5.3. MODELING OF THE TOOL

83

Where [Kg/m3] is the density of the paste adhesive and

[J/Kg] is the enthalpy generated by the chemical reaction.

5.2.4. Paste adhesive thermal degradation

As for the cure kinetics, the degradation kinetics is modeled

by a Partial Differential Equation (PDE) module, applying the

model described in Chapter 4: Modeling of the Degradation Ki-

netics:

5.11

Additionally, in order to convert the evaporated mass in void

content, the relations described Section 4.5 are used for the

paste adhesive system LME 10049-4 / LMB 6687-2:

5.12

5.3. Modeling of the tool

The test setup in which the simulation tool is based and the

material properties used in this chapter are described in Ap-

pendix B.1. In this context, the electrical conductivity of the

CFRP laminates is modeled as an isotropic property. Ideally,

the modeling of this property should consider the weave

structure of the fabric, the layup of the laminate, analyzing the

fiber junctions and the loops they generate [83]. However, in

this investigation it has been simplified by measuring this

property for a bigger section of the laminate. The value ob-

tained of the electrical conductivity in the direction of the fi-

bers has been used for the whole material domain. The reason

for this simplification is that in the 2-D model, only the con-

ductivity in z-direction, i.e. the direction out of the plane of the

2-D model, affects the temperature generation.



84

The heat is firstly generated in the electrical conductive adher-

ents ( ), mainly in the area closer to the coil.

Conduction: The CFRP plates are heated by induction

and the heat is transferred to the materials in contact

with the CFRP.

Convection: As mentioned earlier, forced convection

is considered to model the heat exchange between the

plates and the air. Details of the calculation of the

convection coefficient are given in the Appendix B.2.

To create the flow of the air, a commercial fan has

been used to create a flow of 2 m/s at the beginning of

the plate, measured with an anemometer (Voltcraft

PL-130, Conrad, UK).

Radiation: The modeling of radiation is also consid-

ered in the model.

To complete the modeling of the tool, the material parameters

and boundary conditions are defined following the equations

defined in Section 5.2 as shown in Table 5.1.

Table 5.1: Summary of relations applied in the model.

Description Equation relation Type

Solid materials (Tool+ ad-

herents) 6.1, 6.6 Domain

Air 6.6 Domain

Adhesive 6.1, 6.6, 6.10, 6.11 Domain

Materials in contact with air

(excepting copper)

6.8, 6.9 Boundary

Copper-Air 6.5 Boundary

External frame 6.3, 6.7 Boundary

5.4. VALIDATION AND DISCUSSION OF THE TOOL

85

5.4. Validation and discussion of the tool

For the validation of the model, the temperature and the de-

gree of cure are measured under different conditions and then

compared to the results of the simulation tool. The validation

of the void content is not considered because is carried out in

Chapter 4: Modeling of the degradation kinetics.

Previously to the experimental validation, the operational pa-

rameters are defined. The electrical current, frequency applied

and room temperature are measured for every experiment to

have the same value in the simulation. Finally, as the efficien-

cy of the coil [-], defined in Eq. 5.5, is unknown, it has to

be assessed. As a consequence, the temperature at the point A

(see Figure 5.2) is compared to the predicted values from the

simulation tool.

Figure 5.2: Validation points for steady state and transient.

The processing parameters used in these measurements are 44

A at 262 kHz with an ambient temperature of 24.5 °C. The

temperature is measured 20 times with an optical fiber py-

rometer [FTC-DIN-ST-HA-LS by Photon control] at point A

after the steady state is reached and then compared to predict-

A B C D E

Distances betweenmeasuring points: A-B: 10 mmA-C: 20 mmA-D: 30 mmA-E: 40 mm



86

ed values varying the efficiency [-] from 0.86 to 0.94, as

can be observed in Figure 5.3.

Figure 5.3: The best estimation for the efficiency is ηcoil= 0.9.

The temperature at the point A shows a good agreement with

the model, if an efficiency of 0.90 is assumed [84]. This value is

in the range of expected values and it is used for the validation

of the model.

5.4.1. Validation of the magnetic field and heat transfer

module

For the validation of the coupling of magnetic field and heat

transfer, the temperatures in the model are compared to the

temperatures measured on the test rig at the points A to E,

(see Figure 5.2), considering:

Steady state measurements: The temperature distribu-

tion along the CFRP plate is measured 20 times for

each point between A and E every 10 seconds after the

plates reach the maximum temperature.

30

40

50

60

70

80

90

Pyrometer Eff = 0,86 Eff = 0,90 Eff = 0,94

Tem

per

atu

re [

°C]


87

Heating and cooling of the CFRP plate: Temperatures

are measured every second twice for each transient

process (300 s for heating and 200 s for cooling) in the

points A, B and C.

The CFRP plates used for the validation of the interaction of

magnetic field and heat transfer are previously bonded, mean-

ing that curing reaction is completed ( W/m3). The cur-

rent applied by the equipment for these measurements is 44 A

at 262 KHz. The results of steady state temperatures at the dif-

ferent points are shown in Figure 5.4.

Figure 5.4: Temperature validation at steady state.

The temperature distribution measured at the upper side of

the CFRP joint is similar to the predicted one, having a maxi-

mum difference of about 6 °C. This difference can be due to

the simplification assumed on the modeling of the thermal

conductivity of the CFRP plates. The model shows a high ac-

curacy at the center of the bondline, the area of interest, hav-

ing less accuracy with the distance.

To show the effect of this property on the temperature distri-

bution and to measure a potential model inaccuracy, sensitivi-

ty analyses are carried out for the thermal conductivity of the

30

35

40

45

50

55

60

65

70

75

80

0 0.01 0.02 0.03 0.04

Tem

per

atu

re [

°C]

Measuring point

Pyrometer

COMSOL

A B C D E



88

CFRP plates in x-direction ( ) and y-direction ( ). The direc-

tions are defined in Figure B.4. Properties are simulated for

in the range of 1 to 5 W/(m*K) and for in the range of 0.25 to

0.75 W/(m*K), being the reference values 2 W/(m*K) and 0.5

W/(m*K) respectively. Results of both sensitivity tests are

shown in Figure 5.5 and Figure 5.6, with points A to E defined

in Figure 6.2.

Figure 5.5: Temperature for different values of λx.

Figure 5.6: Temperature for different values of λy.

30

40

50

60

70

80

90

0 0.005 0.01 0.015 0.02 0.025 0.03 0.035 0.04

Tem

per

atu

re[°

C]

Measuring point

λx = 1 [W/(m*K)]

λx = 2 [W/(m*K)]

λx = 3[W/(m*K)]

λx = 4 [W/(m*K)]

λx =5 [W/(m*K)]

Pyrometer

A B C D E

30

40

50

60

70

80

90

0 0.005 0.01 0.015 0.02 0.025 0.03 0.035 0.04

Tem

per

atu

re [

°C]

Measuring point

λy = 0,25 [W/(m*K)]

λy = 0,5 [W/(m*K)]

λy = 0,75 [W/(m*K)]

Pyrometer

A B C D E


89

Figure 5.5 and Figure 5.6 prove that only affects the tem-

perature distribution. Considering this fact, a temperature de-

pendency of the thermal conductivity could be considered in

order to get more accurate results. However, this simulation

model is designed to predict the degree of cure of the paste

adhesive, placed at the center of the sample. At this region, in

the steady state, the prediction of the temperature only has a

deviation of 1.2 °C, compared to the experiments. This devia-

tion is inside the standard deviation of the measured values.

For the investigation of the accuracy of the model in the tran-

sient, the heating and the cooling processes of the CFRP are

measured for 10 minutes at the point A and compared to the

simulated values, as shown in Figure 5.7.

Figure 5.7: Heating and cooling at the overlap center.

The simulated heating process is faster than the measured

heating process and the simulated cooling process is slower.

The reason can be explained by an underestimation of the

convection coefficient coming from an error at the air speed

measurement due to the accuracy of the anemometer, equal to

±0.2 m/s.

20

30

40

50

60

70

80

0 100 200 300 400 500 600

Tem

per

atu

re [

°C]

Time [s]

Pyrometer

COMSOL



90

In order to analyze the impact of the convection coefficient on

the measurements, a sensitivity test is carried out from 16.56

to 18.30 W/m2*K, varying 5 % the reference value of 17.43

W/m2*K. Results are shown in Figure 5.8 and Figure 5.9.

Figure 5.8: Heating at Point A for different hc.

Figure 5.9: Cooling at Point A for different hc.

It can be observed that by increasing the convection coefficient

by 5 % the results fit better to the experimental measurements.

This fact indicates that the convection coefficient was slightly

underestimated most probably due to the reasons forehead

mentioned.

20

30

40

50

60

70

80

90

0 100 200 300 400 500 600

Tem

per

atu

re [

°C]

Time [s]

hc = 16,56 W/m^2*Khc = 17,43 W/m^2*Khc = 18,30 W/m^2*KPyrometer

20

30

40

50

60

70

80

90

0 100 200 300 400 500 600

Tem

per

atu

re [

°C]

Time [s]

hc = 16,56 W/m^2*K

hc = 17,43 W/m^2*K

hc = 18,30 W/m^2*K

Pyrometer


91

In order to study with more detail the heating and cooling

stages, the temperature is measured at the points A, B and C

(see Figure 6.2). Then, they are compared to predicted values

for a transient of 300 seconds in the heating and 200 seconds in

the cooling, as shown in Figure 5.10 and Figure 5.11.

Figure 5.10: Heating at different distances.

Figure 5.11: Cooling at different distances.

The measured temperatures have a satisfying agreement with

the simulated values at the center as well as at the other posi-

tions under study. The values differ in the heating and cooling

20

30

40

50

60

70

80

0 50 100 150 200 250 300

Tem

per

atu

re [

°C]

Time [s]

Pyrometer Point A

COMSOL Point A

Pyrometer Point B

COMSOL Point B

Pyrometer Point C

COMSOL Point C

20

30

40

50

60

70

80

0 50 100 150 200

Tem

per

atu

re [

°C]

Time [s]

Pyrometer Point A

COMSOL Point A

Pyrometer Point B

COMSOL Point B

Pyrometer Point C

COMSOL Point C



92

phases less than 6 °C for all the distances measured, and 2 °C

at the center, where paste adhesive is placed.

5.4.2. Validation of the cure kinetics model

To validate the entire model, plates are bonded by induction

with different conditions, measuring the degree of cure by

DSC. The results for 14 plates summarized in Table 5.2.

Table 5.2: Results from the validation of the model.


93

Input parameters DSC [%] COMSOL

[%]

Rel.

error [%]

800 s at 60 A & 257 kHz +

500 s at 66.4 A & 256 kHz 97.3 96.7 0.6

820s at 72 A & 256 kHz 98.5 98.8 0.3

500 s at 60 A & 257 kHz +

800 s at 66.4 A & 256 kHz 98.1 97.3 0.8

500 s at 60 A & 257 kHz +

800 s at 72 A & 256 kHz 99.8 99.4 0.4

200 s at 60 A & 257 kHz +

600 s at 72 A & 256 kHz 97.9 98.2 0.3

800 s at 34.8 A & 266 kHz +

800 s at 44 A & 262 kHz 57.0 58.4 1.4

800 s at 52 A & 259 kHz 75.5 74.9 0.6

500 s at 44 A & 262 kHz +

500 s at 52 A & 259 kHz 71.3 70.5 0.8

350 s at 61.6 A & 257 kHz 78.0 78.2 0.2

1800 s at 34.8 A & 266 kHz 43.5 44.5 2.3

2400 s at 30.8 A& 273 kHz 37.3 36.6 1.9

900 s at 25.8 A& 274 kHz 13.0 12.6 3.1

2400 s at 25.8 A& 274 kHz 30.1 29.5 2.0

3600 s at 25.8 A& 274 kHz 37.8 39.3 3.7

The simulation predicts the degree of cure of the adhesive un-

der different processing conditions with a relative error lower

than 4 %, even in the samples cured less than 50 %. The reason

is the good agreement of the temperatures generated at the

center of the plate with the model, where the paste adhesive is

placed and measured. These results are in agreement with the



94

validation of the cure kinetics model, showing a higher error

for lower curing rates. At the ideal application range of the

tool, with a degree of cure above 80 %, the results show an

excellent accuracy with a maximum relative error of 0.8 %, not

showing any tendency between experimental and predicted

values.

5.5. Impact of parameters on the curing pro-

cess

Finally, a study is carried out to analyze which parameters

used for the simulation significantly affect the curing process

of the paste adhesive. The objective is to point out the im-

portant parameters that have to be controlled in an industrial

environment for robust processing. Additionally, potential

applications of modeling the curing process of paste adhesives

by induction heating are discussed.

The parameters considered for this sensitivity analysis can be

divided into four categories:

Paste adhesive material properties.

Induction processing parameters.

Adherent material properties.

Geometry of the assembly.

The simulation is considered for 1 hour of heating applying an

electrical current of 44 A, a frequency of 262 kHz and an ambi-

ent temperature of 24.5 °C.

5.5.1. Effect of adhesive properties on the chemical reac-

tion

The following adhesive properties are considered for the sen-

sitivity tests:

Thermal conductivity from 0.18 to 0.48 W/m*K. The

reference value is 0.33 W/m*K.

5.5. IMPACT OF PARAMETERS ON THE CURING PROCESS

95

Heat capacity from 1’000 to 1’800 J/kg*K. The refer-

ence value is 1’400 J/kg*K.

Density from 700 to 1’500 kg/m3. The reference value

is 1’100 kg/m3.

Surface emissivity from 0.8 to 1. The reference value is

0.9.

Simulated results of the degree of cure are shown in Table 5.3.

Table 5.3: Sensitivity analysis of adhesive properties.

Adhesive thermal conductivity

[W/m*K] 0.18 0.33 0.48

Degree of cure[%] 83.2 83.2 83.1

Adhesive heat capacity [J/kg*K] 1’000 1’400 1’800

Degree of cure[%] 83.2 83.2 83.1

Adhesive density [kg/m3] 700 1’100 1’500

Degree of cure [%] 83.2 83.2 83.1

Adhesive surface emissivity [-] 0.8 0.9 1

Degree of cure[%] 83.2 83.2 83.2

The physical properties of the paste adhesive do not influence

the degree of cure significantly despite simulating a wide

range of values. The main reason for this is the small paste

adhesive domain. This might be different with a thicker

bondline. In that case, a higher exothermal curing energy

would be generated and a variation on the thermal properties

would have a higher impact on the degree of cure. Neverthe-

less, in this study bondlines thicker than 0.5 mm are not con-

sidered as discussed in Appendix A.5.



96

5.5.2. Effect of induction processing parameters on the

chemical reaction

These processing parameters are analyzed in the following

range of values:

Convection coefficient from 10 to 25 W/m2*K. The ref-

erence value is 17.43 W/m2*K.

Ambient temperature from 10 to 30 °C. The reference

value is 24.5 °C.

Electrical current of the coil from 40 to 48 A. The refer-

ence value is 44 A.

Electrical frequency of the coil from 233’100 to 284’900

Hz. The reference value is 259’000 Hz. Despite this

value is coupled to the current applied in the equip-

ment used in this research, the effect of a variation of

this parameter is also considered.

Efficiency of the coil from 0.86 to 0.94. The reference

value is 0.9.

Simulated results of the degree of cure during the heating pro-

cess are summarized in Table 5.4.


97

Table 5.4: Sensitivity analysis of the processing parameters.

Convection coefficient [W/m2*K] 10 15 20

Degree of cure[%] 98.2 87.2 77.2

Ambient temperature [°C] 10 20 30

Degree of cure[%] 70.3 79.9 86.7

Electrical current [A] 40 44 48

Degree of cure [%] 75.4 83.2 89.8

Frequency [kHz] 235.8 262 288.2

Degree of cure[%] 76.1 83.2 89.0

Efficiency of the coil [-] 0.86 0.90 0.94

Degree of cure[%] 79.7 83.2 86.5

Despite the parameters are varied in different magnitudes, all

of them influence on the final degree of cure and the void

formation. The main output of these experiments is to prove

the importance of a correct measurement of the experimental

conditions. Convection coefficient and ambient temperature

are some of the parameters affecting the most the heat transfer

between the components. For this reason, its accurate meas-

urement is critical in order to get reliable results. Additionally,

it is proved that the efficiency also affects the induced energy

and requires an accurate definition in order to get a correct

simulation tool.

These results show the main outcome of the model: the pro-

cessing parameter optimization for a robust induction heating

process. One of the main potential applications of induction

heating is its application outdoors; for instance in repair oper-

ations. In this context, this simulation has to be used to define

the optimal input parameters (frequency and electrical cur-



98

rent) to cure the paste adhesive considering the change of the

environmental conditions.

5.5.3. Effect of adherent properties on the chemical reac-

tion

The change of the adherent properties is critical to understand

how a change in the layup can affect the heat generated. For

example, increasing the number of layers or changing the ori-

entation of them may affect the thermal and electrical proper-

ties of the laminate, thus affecting the energy generated in the

susceptor. In this context, the following properties are studied:

Electrical conductivity of the adherents from 5’321 to

6’503 S/m. The reference value is 5’912 S/m.

Thermal conductivity of the CFRP plates is changed

from 1 to 5 W/(m*K) for and from 0.25 to 0.75

W/(m*K) for . The reference values are for = 0.5

W/(m*K) and = 2 W/(m*K).

Simulated results of the degree of cure are shown Table 5.5.

Table 5.5: Sensitivity analysis of adherent properties.

CFRP electrical conductivity [S/m] 5’321 5’912 6’503

Degree of cure[%] 80.4 83.2 85.4

CFRP x-thermal conductivity

[W/(m*K)] 1 3 5

Degree of cure[%] 85.3 81.6 79.1

CFRP y-thermal conductivity

[W/(m*K)] 0.25 0.5 0.75

Degree of cure [%] 83.3 83.2 83.1

It is proved that by increasing the electrical conductivity of the

adherents, the curing process of the paste adhesive will accel-


99

erate due to the higher inducted currents. Additionally, the

degree of cure increases when decreasing the thermal conduc-

tivity in the case the x-direction, since the heat dissipates

slower. The impact of varying is higher than a change of ,

due to the small thickness of the plates.

These results show other potential applications of the model

than presented in this Chapter; e.g. the analysis of the perfor-

mance of induction heating for other materials and the estima-

tion of unknown material parameters.

5.5.4. Effect of the geometry on the chemical reaction

Finally, the effect of different geometrical parameters of the

simulation is discussed. The goal is to analyze the effect that

thickness variations and an irregular distance between the

susceptor and the coil would have in the degree of cure of the

adhesive, shown in Figure 5.12. For this reason, the following

properties are studied:



100

Figure 5.12: Distance between the lower coil and the plate.

The distance between the plate and the coil is changed

from 4 to 8. The reference value is 6 mm.

The thickness of the laminate is changed from 0.66 to

1.3 mm, having from 2 to 4 layers of woven fabric ma-

terial all with an orientation [0, 90].

As the thickness is changed, the electrical conductivity of the

laminates is also varied. For this reason this value is measured

following the approach described in Section B.1. The results of

the electrical conductivity for the different plates are summa-

rized in Table 5.6.

DISTANCE Lower coil- plate

35.4 mm

24 mm

10 mm

10 mm

15 mm25 mm

15 mm

LAMINATE THICKNESS

xz

y

5.6. CONCLUSIONS

101

Table 5.6: Electrical conductivity of laminates with different number of layers.

Number of layers CFRP electrical conductivity [S/m]

2 5’527 ± 113

3 5’912 ± 94

4 5’176 ± 85

Simulated results of the degree of cure are shown in Table 5.7.

Table 5.7: Sensitivity analysis of adherent properties.

Distance between coil and plate [mm] 4 6 8

Degree of cure[%] 98.5 83.2 61.6

Thickness of the laminate [mm] 0.66 1 1.33

Degree of cure [%] 72.2 83.2 87.6

Both parameters are proved to have a critical influence on the

temperatures generated and therefore on the chemical reac-

tion. For this reason, the geometrical parameters must be very

accurately controlled in order to guarantee the robustness of

the process.

5.6. Conclusions

In this chapter, the simulation tool to predict the degree of

cure and the void content of paste adhesives is described. This

simulation tool is validated, showing a good agreement with

experimental values, especially at the bondline center. Addi-

tionally, the validation of the degree of cure also shows a good

accuracy for different curing profiles measured, especially for

samples with a degree of cure higher than 90 %.

At this point, the simulation is used to carry out a sensitivity

analysis of material and process parameters. The goal is to



102

analyze the impact of a change in one property on the degree

of cure of the adhesive after one hour of curing process. It is

proved that the parameters affecting most the degree of cure

are mainly the processing parameters (convection coefficient,

ambient temperature, electrical current, frequency and effi-

ciency of the coil). The thermal material properties of the ad-

herent show a minor influence on the degree of cure but still

higher than the influence of the adhesive material properties.

Therefore, the correct measurement and control of the pro-

cessing parameters is important to predict with accuracy the

degree of cure and void formation of the paste adhesive and to

assess an optimal bonding process.

This simulation tool is designed to ensure the robustness of

the induction heating process by calculating the optimal input

parameters. As this process can be applied outdoors, some

parameters may change, affecting the chemical reactions. In

this context, this simulation tool can be used to adapt the in-

put parameters, current and frequency, ensuring a correct

bonding process.

Additionally, it can be used to assess unknown material prop-

erties e.g. the thermal properties of a CFRP laminates with a

different layup or to analyze the impact of other heating strat-

egies or susceptors. This simulation tool shows a wide range

of applications to guarantee the robustness of the curing pro-

cess of adhesives by induction heating.

103

Part III

Process Optimization

105

Chapter 6

Isothermal heating process

In this chapter, the influence of high temperatures on the qual-

ity of the paste adhesive is investigated. Different techniques

are used to analyze the quality of the paste adhesive with cer-

tain thermal degradation caused during the curing process.

Finally, the results are compared to today’s state of the art,

SLS; aiming at investigating other valid indicators for qualifi-

cation of bonded joints.


an Article in the International Journal of Adhesion and Adhesives

(see Sánchez Cebrián, A., Zogg, M., and Ermanni P.,

Methodology for optimization of the curing cycle of paste adhesives.

International Journal of Adhesion and Adhesives, 2013. 40: p.

112-119.)

6.1. Introduction

As observed in Chapter 3: Modeling of the Cure Kinetics, the

curing process of a paste adhesive can be simply accelerated if

the temperature is increased [17, 28]. Today in industry, this is

usually not considered by the reasons stated in Chapter 2.

This chapter investigates the effect of high temperatures on

different physical properties of the paste adhesive system

LME 10049-3 / LMB 6687-1. The objective is to analyze the re-

lation between curing temperature and mechanical perfor-

mance of the paste adhesive and to set the process limits.

CHAPTER 6: ISOTHERMAL HEATING PROCESS

106

The samples used in this investigation are completely cured

without pressure, applying different curing cycles, with tem-

peratures from 80 °C to 200 °C [19]. The tests considered in

this study include thermal analysis techniques, e.g. DSC and

TGA, as well as optical microscopy. Additionally, mechanical

testing is considered, including three point bending test and

SLS.

As a consequence, a methodology to assess the quality control

of paste adhesives is discussed, obtaining a useful tool to vali-

date the curing processes investigated thorough this disserta-

tion.

6.2. Experimental

In this investigation, most of the samples to analyze the physi-

cal properties of the pure paste adhesive are produced by ov-

en heating. The reason is that induction heating cannot heat

directly the paste adhesive due to its low electrical conductivi-

ty. For this reason, SLS samples are the only samples bonded

by induction heating. The experiments performed in this chap-

ter can be divided into physical and mechanical testing.

Experiments are carried out to study the change in physical

properties include:

DSC to validate the curing process of the samples.

TGA analysis of epoxy and hardener by separate to

measure the thermal degradation of the paste adhe-

sive components. The experimental conditions are de-

tailed in Section 5.3: Experimental

Analysis of size and quantity of voids, by microscopy

techniques further detailed in Appendix B.4, but tak-

ing in this case 10 images on each sample instead of 3.

Additionally, the void content obtained by optical testing is

compared to the measurement of the density of the pure paste

6.3. PHYSICAL PROPERTIES OF THE PASTE ADHESIVE

107

adhesive samples used in this investigation. The mechanical

tests considered are:

DMA 3 point bending test as detailed in Section 5.3:

Experimental.

3 point bending test of pure adhesive following the

ISO 178.

SLS of CFRP bonded systems, as detailed in Appendix

A.3.

6.3. Physical properties of the paste adhe-

sive

Table 6.1 summarizes the degree of cure of the different sam-

ples used for optical testing measured with the DSC, following

the procedure detailed Section 3.2. The total released energy of

the paste adhesive is 262 J/gr.

Table 6.1: Degree of cure of the samples measured by DSC.

Temperature

[°C]

Time

[min]

Released energy

[J/g]

Degree of

cure [%]

80 240 10.5 96.7

100 60 7.6 97.6

120 60 7.6 97.6

140 45 8.4 97.4

160 30 11.4 96.4

180 15 12.5 96.1

200 10 3.1 99.0

The curing process is complete for all the samples as the de-

gree of cure is higher than 95 %.


108

TGA analysis of the separate components is shown in Figure

6.1.

Figure 6.1: TGA analysis for epoxy and hardener.

TGA shows the onset of the hardener at 123.8 °C and in the

resin at 191.2°C, but a loss of mass can be observed in the

hardener before 120 °C. This means that the samples heated

with this temperature or greater are expected to have a higher

void content. This effect is combined with the expansion of the

voids trapped in the dosing and mixing process. As a result,

samples heated with temperatures higher than 100 °C have

more and larger voids that affect the mechanical performance.

By microscopy, the void content of each curing cycle is meas-

ured, as shown in Figure 6.2 to Figure 6.8.

100

98

96

94

92

90

88

8640 60 80 100 120 140 160 180 200 220 240

Temperature [°C]

Wei

gh

t [%

]

LME 10049-3 HardenerOnset = 151.6°C

LME 6687-1 EpoxyOnset = 214.8 °C


109

Figure 6.2: Sample cured at 80 °C (Avg. void content 1.6 %).

Figure 6.3: Sample cured at 100 °C (Avg. void content 1.4%).

1000 µm

1000 µm


110

Figure 6.4: Sample cured at 120 °C (Avg. void content 2.1 %)

Figure 6.5: Sample cured at 140 °C (Avg. void content 21.4%).

1000 µm

1000 µm


111



1000 µm

1000 µm


112


A summary of the void content measured 10 times at each

sample is shown in Figure 6.9.

Figure 6.9: Void content measurement.

Void content for lower temperatures, between 80 °C and

120 °C, is detailed in Figure 6.10.

1000 µm

0%

10%

20%

30%

40%

50%

60%

70%

80%

90%

100%

80 100 120 140 160 180 200

Vo

idco

nte

nt

[%]

Temperature [°C]


113

Figure 6.10: Void content of samples (80 °C to 120 °C).

The results show a lower void formation at the samples cured

at 120 °C or less compared to samples cured at higher temper-

atures. However, an increase of void content is observed in

samples cured from 120 °C compared to samples cured at 80

°C and 100 °C.

These results are compared to TGA analysis, where it is ob-

served that the hardener evaporated if it is heated with more

than 100 °C. This fact can explain the big difference in the void

formation between 120 °C and 140 °C, when evaporation effect

is higher. Comparing Figure 6.4 and Figure 6.5, it can be ob-

served that not only the size of the voids increase, but also the

quantity of voids increases due to the evaporation of volatiles.

The average diameter of the voids is also determined by

measuring ten different voids for each sample. The density of

the samples used in this investigation is also measured and

then compared to the theoretical value according to the tech-

nical specifications of the paste adhesive. Results are shown in

Figure 6.11 and Figure 6.12 and then summarized in Table 6.2.

0.0%

0.5%

1.0%

1.5%

2.0%

2.5%

3.0%

3.5%

80 100 120

Vo

idco

nte

nt

[%]

Temperature [°C]


114

Figure 6.11: Average bubbles diameter of samples.

Figure 6.12: Relative density of samples.

0

100

200

300

400

500

600

80 100 120 140 160 180 200

Vo

idav

erag

ed

iam

eter

[μm

]

Temperature [°C]

20%

30%

40%

50%

60%

70%

80%

90%

100%

80 100 120 140 160 180 200

Rel

ativ

e d

ensi

ty [

%]

Temperature [°C]

6.4. MECHANICAL PROPERTIES OF THE PASTE ADHESIVE

115

Table 6.2: Summary of optical measurements.

Temp.

[°C]

Time

[min]

Av. void

diameter

[µm]

Av. number of

voids

[voids/sample]

Void

content

[%]

Density

[%]

80 240 49 ± 22 8.4 ± 2.9 1.6 ± 0.5 97.6

100 60 62 ± 32 9.25 ± 2.7 1.4 ± 0.5 98.7

120 60 104 ± 29 10.8 ± 2.0 2.1 ± 0.8 95.1

140 45 229 ± 76 16.8 ± 3.5 21.4 ±

6.2

80.7

160 30 263 ± 132 - 33.5 ±

11.6

61.2

180 15 335 ± 236 - 60.5 ±

17.0

41.0

200 10 266 ± 182 - 75.1 ± 1

2.2

25.1

The results show a similar tendency as void content measure-

ment with a clear increase of the diameter for samples cured

with more than 100 °C. The number of voids increases slightly

at the samples cured with temperatures from 80 °C to 120 °C

and increases strongly at the samples cured at 140 °C. As it can

also be observed, the void content values are approximately

inverted values of density, meaning a homogeneous distribu-

tion of the voids in the different samples can be seen.

6.4. Mechanical properties of the paste ad-

hesive

A 3-point bending test is carried out in the DMA for the dif-

ferent samples measuring the storage modulus from 0 °C to

120 °C. Curves of all the samples are shown in Figure 6.14. It

can be observed a lower mechanical performance in samples

cured with higher temperature. In order to evaluate results

and compare them for all the samples, storage modulus at


116

room temperature (20 °C) as well as the Tg [°C] at the maxi-

mum tan δ um are measured as shown in Figure 6.13.

Figure 6.13: DMA measurement at 100 °C.

The results of the rest of the samples are shown in Figure 6.14

and Figure 6.15 and then summarized in Table 6.3.

180 200 220 2400 20 40 60 80 100 120 140 160

Temperature [°C]

20

40

60

80

100

120

140

160

180

200

Sto

rag

e m

od

ulu

s[M

Pa]

Tan

del

ta [

-]

0.05

0.10

0.15

0.20

0.25

0.30

0.35

0.40

0.45

0.50

0.55

0.60Storage modulus [MPa]Tan delta [-]


117

Figure 6.14: Storage modulus measured by DMA.

Figure 6.15: Storage modulus from DMA at 20 °C.

0 10 20 30 40 50 60 70 80 90 100 110 120

50

100

150

200

250

300

350

400

450

500

Temperature [°C]

Sto

rag

e m

od

ulu

s[M

Pa]

240 min @ 80°C

60 min @ 100°C

60 min @ 120°C

45 min @ 140°C

30 min @ 160°C

15 min @ 180°C

10 min @ 200°C

0.00E+00

5.00E+07

1.00E+08

1.50E+08

2.00E+08

2.50E+08

3.00E+08

3.50E+08

4.00E+08

4.50E+08

5.00E+08

80 100 120 140 160 180 200

Sto

rag

e m

od

. 20

°C [

MP

a]

Temperature [°C]

500

450

400

350

300

250

200

100

150

50

0


118

Table 6.3: Storage modulus at 20 °C and Tg.

Temperature

[°C]

Time

[min]

Storage mod.

at 20 °C [MPa] Tg [°C]

80 240 443 111.8

100 60 329 112.8

120 60 285 116.5

140 45 197 120.2

160 30 105 122.0

180 15 40 109.7

200 10 105 120.4

The storage modulus at 20 °C decreases if the samples are

cured at higher temperature. The Tg [°C] remains constant. The

increase in performance of the sample cured at 200 °C com-

pared to the sample cured at 180 °C can be explained by the

high degree of degradation of both samples.

Results of 3-point bending tests are shown in Figure 6.16 and

Figure 6.17 and then summarized in Table 6.4.

0

200

400

600

800

1000

1200

1400

80 100 120 140 160 180 200

Fle

xu

ral

mo

d. [

MP

a]

Temperature [°C]


119

Figure 6.16: Flexural modulus of samples.

Figure 6.17: Flexural strength of samples.

Table 6.4: Summary of results in 3 point bending test.

Tempera-

ture [°C]

Time

[min]

Flexural modulus

[MPa]

Flexural

strength [MPa]

80 240 1098.5 ± 122.3 41.4 ± 4.0

100 60 1190.4 ± 106.3 44.2 ± 1.9

120 60 1062.3 ± 74.8 41.2 ± 2.3

140 45 649.7 ± 120.0 24.5 ± 5.9

160 30 474.0 ± 168.0 15.4 ± 5.8

180 15 194.5 ± 107.0 5.6 ± 2.9

200 10 115.0 ± 23.3 2.9 ± 0.9

The results of the 3-point bending test show that the mechani-

cal performance decreases with curing at high temperatures. A

decrease of performance can be observed for samples cured

with more than 120 °C. The tendency is similar to the DMA

results, but the results according to ISO 178 show high values

for samples cured with 120 °C or less.

0

5

10

15

20

25

30

35

40

45

50

80 100 120 140 160 180 200

Fle

x. s

tren

gth

[MP

a]

Temperature [°C]


120

Finally, the results of the SLS test are summarized in Figure

6.18 and Table 6.5.

Figure 6.18: SLS test for CFRP bonded samples.

Table 6.5: Results of single lap shear test.

Temperature

[°C]

Time

[min]

Shear

strength

[MPa]

Bondline

thickness

[mm]

Fracture

mode

80 240 24.4 ± 2.9 0.20 ± 0.01 Adherent

100 60 22.1 ± 1.8 0.18 ± 0.02 Adherent

120 60 16.3 ± 1.2 0.16 ± 0.01 Cohesive

140 45 15.5 ± 1.6 0.23 ± 0.02 Cohesive

160 30 6.0 ± 1.3 0.19 ± 0.01 Adhesive

180 15 6.1 ± 0.5 0.20 ± 0.04 Adhesive

200 10 6.6 ± 1.4 0.21 ± 0.02 Adhesive

Results show a decrease of shear strength when the curing

temperature is increased, showing an adhesive failure mode in

samples that are highly degraded. Samples cured with 80 °C

and 100 °C show adherent failure, meaning that the paste ad-

0

5

10

15

20

25

30

80 100 120 140 160 180 200

Sh

ear

stre

ng

th [

MP

a]

Temperature [°C]

6.5. ISOTHERMAL ACCELERATION OF THE PROCESS

121

hesive is well cured. Samples with cohesive failure show a

certain degradation of the adhesive for samples cured with

120 °C or more. Finally, samples with adhesive failure show a

poor adhesion between parts meaning that curing process is

not correct due to the high temperatures applied.

6.5. Isothermal acceleration of the process

At this point, a methodology for the quality control of paste

adhesives is defined. In order to analyze the results of the ap-

plied methods, four parameters are considered.

The size of the sample gives information about the poten-

tial in real applications. In this sense, methods that need a

small amount of adhesive are more suitable because ac-

cess to the bondline can be sometimes difficult.

Dependency on other materials is undesirable because

they can influence the assessment of the adhesive quality.

Sensitivity to predict degradation is needed to assess ac-

curately if a sample is valid.

The possibility to qualify by giving a minimum value

used not only for specific paste adhesive but in general is

also desirable to generalize the methodology.

After considering these parameters, the experiments consid-

ered in this investigation are discussed:

DSC is only used to measure the degree of cure. It does not

give any information about the quality of the paste adhesives.

DMA measurements results do not show a clear point where

degradation is started. They do not differ between curing at

100 °C, where the paste adhesive has good quality, and curing

at 120 °C, when the paste adhesive is already degraded. By

this technique results cannot be compared to other adhesives

because mechanical properties can differ without meaning

degradation.


122

TGA analysis cannot explain the entire decrease of perfor-

mance. The reason is that evaporation of particles can only

explain formation of new voids and not the expansion of voids

trapped in the mixing process. Besides, it does not give the

possibility of qualification. Nevertheless, this analysis is im-

portant to assess the limit temperature for the curing process.

Mechanical and density measurements require big samples to

be tested. This is not always possible to acquire from real ap-

plications, where the access to the bondline can be limited.

Additionally, in the case of 3 point bending there is not a clear

difference between curing at 100 °C and 120 °C.

SLS testing, despite being today’s state of the art measuring

technique, depends on the quality of the adherent used. This

fact can influence the results significantly, leading to misun-

derstanding.

Finally, optical analysis is found to be the most accurate indi-

cator to assess the bonding quality. It has the possibility to

analyze the quality of the bonding system with a small sam-

ple. This method restricts the use of curing at 120 °C for pri-

mary structures, having a void content higher than 2 %, and

bubbles that increase in size compared to samples cured at

lower temperatures.

A summary of the parameters considered for the evaluation

and the performance of each method is shown in

6.5. ISOTHERMAL ACCELERATION OF THE PROCESS

123

Table 6.6: Techniques to validate paste adhesives.

Sample

size

Independency

from other

materials

Sensitivity to

predict degra-

dation

Possibility for

qualification

DSC + + - -

DMA + + - +

TGA + + - +

3 pt

bending - + +- +

Density - + +- +

SLS +- - +- +

Optical

testing + + + +

In order to assess the sensitivity of the measurement of the-

void content, it is compared to SLS test in Figure 6.19.

Figure 6.19: Void content compared to SLS test.

0

0.5

1

1.5

2

2.5

3

3.5

0 5 10 15 20 25 30

Vo

idco

nte

nt

[%]

Shear strength [MPa]

80 C100 C

120 C


124

As shown in the results, curing at 120 °C produces a void con-

tent higher than 2 % that will lead to a lower mechanical per-

formance. The high deviation of the results indicates that there

are areas where the evaporation of particles has started,

whereas in other areas the lower void content is measured.

Observing all the experiments carried out, some of them show

more clearly a decrease of performance on samples cured us-

ing temperatures higher than 100 °C. In order to observe better

the behavior of the adhesive at lower temperatures, from 80 °C

to 120 °C, the results of the selected experiments are shown.

They are divided by the values of the curing reference at 80 °C

and then compared in Figure 6.20.

Figure 6.20: Results referenced to supplier’s curing.

All the tests show a decrease of performance on samples cured

with 120 °C. Nevertheless, void content measurement and sin-

gle lap shear test show more clearly this decrease of perfor-

mance on samples heated with 120 °C compared to the refer-

ence curing at 80 °C and the samples cured at 100 °C.

40%

50%

60%

70%

80%

90%

100%

110%

120%

130%

140%

80 100 120

Pro

per

ty/

Pro

per

ty a

t 80

°C

Temperature [°C]

DMA storage mod.

3-pt-Flexural modulus

3-pt-flexural strength

Density of samples

Bubble's diameter

Porosity

Shear strength

6.6. CONCLUSIONS

125

As a conclusion, four techniques are selected within the tech-

niques applied in this study to define a quality control meth-

odology: TGA, DSC, single lap shear test and void content

measurement by optical testing.

TGA analysis of a non-cured paste adhesive sample deter-

mines the temperature where the evaporation, which will lead

to the increment of void content, begins. Once the onset is cal-

culated, this temperature is used as a first assessment for the

upper limit for the manufacturing of samples.

After curing the samples at different temperatures, DSC must

be carried out to ensure the complete curing of samples. Then

mechanical testing assesses the temperature limit where the

performance of the adhesive is affected. Recommended me-

chanical testing is today’s state of the art, SLS. The limit tem-

perature for the curing process is selected by observing when

the mechanical performance of the different samples decreases

and the fracture mode turns from adherent to cohesive.

Additionally, optical testing is recommended to measure the

void content in order to validate the results from TGA. The

advantage of this testing, compared to SLS, is that it is inde-

pendent from the adherent quality. In this sense, it gives in-

formation only about the paste adhesive quality, which will be

affected if the void content is higher than 2 %.

6.6. Conclusions

A temperature increase is used in this Chapter to accelerate

the curing process, proving that the void formation influences

the mechanical performance of the joint. For this reason this

value must be limited, ensuring that the curing process does

not degrade the paste adhesive. In this chapter, it is proved

that a void formation lower than 2 % does not influence the

mechanical performance of the joint.


126

The analysis of the void content is shown to be an accurate

method to assess the quality of a cured paste adhesive. This

system is simple, needing only one optical study with the mi-

croscope. The technique can be used complementary to single

lap shear tests. This is especially useful because it does not

depend on the quality of the adherent and requires small

samples that can be easily obtained.

The analysis of the void diameter and void content is proved

to represent the quality of the paste adhesive under study.

This technique shows a main advantage that can always be

applied, even when traditional NDI techniques cannot be ap-

plied due to the small void size.

127

Chapter 7

Cure dependent heating process

In this chapter, a multistep heating approach is considered to

accelerate the curing process. This chapter analyzes experi-

mentally the effects of a temperature increase at different cur-

ing stages to the void formation.


an Article in the International Journal of Adhesion and Adhesives

(see Sánchez Cebrián, A., Basler, R., Klunker, F., Zogg M., Ac-

celeration of the curing process of a paste adhesive for aerospace ap-

plications considering cure dependent void formations,

International Journal of Adhesion and Adhesives, 2014. 48: p.

51-58.)

7.1. Introduction

This chapter presents the results of a research investigating the

acceleration of the curing process of the paste adhesive system

LME 10049-4 / LMB 6687-2 by means of a cure dependent

heating process.

In chapter 6, the curing process could be accelerated with a

single isothermal heating stage at 100 °C for 1 hour without

affecting the mechanical performance of the joint. Curing the

paste adhesive with a higher temperature significantly in-

creased the void formation. In this context, the objective of this

chapter is the investigation of a non-isothermal curing strategy

to accelerate and its impact in the curing process of the paste

adhesive.

CHAPTER 7: CURE DEPENDENT HEATING PROCESS

128

As a first approach to non-isothermal curing processes, dual

step heating processes are used in this investigation, i.e. a pro-

cess to cure a paste adhesive consisting of two temperature

steps. This approach is based on the hypothesis that paste ad-

hesives are more vulnerable to void formation at the early

stages of curing, as proved in Chapter 4: Modeling of the degra-

dation kinetics. For this reason, low temperatures are initially

applied. Then, as soon as a certain degree of cure is reached,

the temperature might be raised without increasing the void

formation of the paste adhesive. The dual step heating process

represents a first step towards optimization of curing cycles of

paste adhesives by variable temperature cycles. In this context,

the impact of increasing the temperature at different stages of

the curing reaction must be previously evaluated.

7.2. Approach

The approach to accelerate the curing process of paste adhe-

sives consists of a cure dependent heating process, increasing

the curing temperature when the degree of cure is high

enough to ensure a low void formation. The ideal curing pro-

cess is expected to be a ramp with a variable slope. It would

starting at a low temperature, increasing it when the paste ad-

hesive becomes more resistant to void formation due to the

higher degree of cure: When the degree of cure exceeds the

gelation point, typically with a degree of cure between 55 and

80 %, the solidification occurs [85, 86], therefore the formation

of voids is hindered. As a consequence, such a curing strategy

would ensure a low void formation and thus maintaining a

good mechanical performance.

As a first approach towards an optimum curing cycle with

regards to low final void content, this Chapter considers a du-

al stage heating process. It consists of two isothermal heating

stages, as it is shown in Figure 7.1. The heating rates applied

7.2. APPROACH

129

in this study for reaching the two isothermal stages are set to

25 °C/min. For the used setup this is the maximum heating

rate which can be applied accurately in order to ensure a re-

producible quality of the samples. Lower heating rates are not

considered because they would unnecessarily increase the

overall heating time. 25 °C /min seems to be a good compro-

mise between acceleration and accuracy of the applied tem-

perature. For this reason, this rate is used if not indicated the

contrary in this dissertation to accelerate the curing process of

samples.

Figure 7.1: Two step heating process.

The range of the parameters affecting the void formation and

the degree of cure to be studied are:

Initial temperatures [°C] in the range of 80 °C to 160

°C. Chapter 6 has shown that higher temperatures de-

grade the paste adhesive too much and lower temper-

atures do not improve the quality of the paste adhe-

sive but require longer curing cycles.

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0

20

40

60

80

100

120

140

160

180

0 500 1000 1500 2000

α > 0.95

Temperature [°C]Degree of cure [-]

T1

T2

t1 t2

Tem

per

atu

re[°

C]

Deg

ree

of

cure

[-]

Time [min]


130

The duration of the first heating step [min] is de-

fined in steps of 5 minutes between 5 and 20 minutes.

The reason is to get samples with a wide range of de-

gree of cure after the first heating stage. Then, the in-

fluence of the gelation in the void formation can be

analyzed.

The second stage at [°C] is carried out in the range

of 140 °C to 180 °C to reach a degree of cure of 95 %.

This is the minimum considered for aerospace appli-

cations. Lower temperatures are already considered

for the first step ( [°C]). Higher temperatures are not

considered because the maximum achievable [°C]

of the adherents is around 190 °C, and the bonded

joint could be damaged.

The duration of the second step, is varied in steps

of 5 minutes. As a consequence, samples with a degree

of cure higher and lower than 95 % are obtained. This

allows calculating the necessary heating time at the

second step to achieve a degree of cure of exactly 95

%. Additionally, the void content at this point can be

calculated by linear regression.

The aforementioned processing parameters of the curing pro-

cess affect the degree of cure and the void formation. In order

to assess its influence, samples of pure paste adhesive are pro-

duced applying different curing cycles, as described in Ap-

pendix B.4. Then, the void content of these samples is meas-

ured by microscopy techniques. The cure kinetics model for

this paste adhesive, detailed in Chapter 3: Modeling of the Cure

Kinetics, is used to calculate the resulting degree of cure in the

sample.

The following information about the curing cycles applied is

obtained:

7.2. APPROACH

131

Calculated degree of cure after the first heating step

( )

Total curing time considering the time of the

heating ramps ( )

Calculated degree of cure after the curing process (

Measured void content ( )

Assessment of [min] necessary to achieve 95 % of

degree of cure ( )

Total curing time to get a degree of cure of 95 %

( )

Estimation of the void content of samples with a de-

gree of cure of 95 % ( )

Finally, samples cured with the same variation of curing cy-

cles are mechanically tested to validate the cure dependent

heating process. The selection of samples for mechanical test-

ing is based on two criteria, which are ensuring the accelera-

tion of the process:

The overall curing time is shorter than 30 minutes in-

cluding heating ramps

The degree of cure is greater equal 95 %, as required

in aerospace applications [10].

Mechanical testing is performed with the objective to evaluate

the decrease of the mechanical performance of the paste adhe-

sive by the void formation. In this context, samples are pro-

duced with different curing cycles thus generating different

void formations. Additionally, 5 more sample plates are pro-

duced curing the adhesive isothermally with temperatures

from 80 °C to 160 °C. These plates are used as reference in or-

der to show the potential for the acceleration of the curing

process by means of the cure dependent heating process.


132

7.3. Impact of the processing parameters

7.3.1. Impact of on the curing process

The samples tested for the analysis of are produced with an

initial stage of 15 minutes. As a result, samples with the de-

gree of cure in the range of the expected gelation point (55-80

%) are obtained. The second stage is carried out at 140 °C in

order to minimize the void formation. This stage is carried out

until the curing process is finished. are also varied

afterwards in order to observe their impact on the curing pro-

cess and the void formation. The results of the analysis of

are shown in Table 7.1.

7.3. IMPACT OF THE PROCESSING PARAMETERS

133

Table 7.1: Effect of T1 [°C] on the curing process.

[

°C]

[

min

]

[

-]

[

°C]

[

min

]

[m

in]

[

-]

𝑽𝒄 [

-]

=𝟗𝟓

[m

in]

=

𝟗𝟓

[min

] 𝑽

𝒄𝟗𝟓

% [

-]

80

15

57.0

14

0 5

24.7

92

.9

4.1

7.2

26.8

4.

3 80

15

57

.0

140

10

29.7

96

.5

4.5

100

15

79.9

14

0 5

24.7

94

.1

1.3

6.2

25.8

1.

4 10

0 15

79

.9

140

15

34.7

97

.9

1.5

120

15

92.3

14

0 2

21.7

94

.8

3.8

2.4

22.1

3.

8 12

0 15

92

.3

140

5 24

.7

96.3

3.

9

140

5 90

.9

- -

9.7

90.9

4.

4 -

13.5

5.

5 14

0 15

97

.4

- -

19.7

97

.4

6.1


134

As already discussed in Chapter 6; the initial curing tempera-

ture affects critically the void formation during the curing

process. For this reason, and in order to fulfill the maximum

void content value of 2 % considered in this investigation, the

initial curing temperature must be limited for this paste adhe-

sive to 100 °C.

The results of the analysis of the impact of in the curing

process are shown in Table 7.1 and can be summarized as fol-

lows:

An initial heating stage at 80 °C with a premature in-

crease of temperature leads to a high void formation

(4.3 %). In this case, the increase of temperature is car-

ried out when the paste adhesive has a degree of cure

of 57 %, the lowest point of solidification.

Curing the paste adhesive initially with 100 °C for 15

minutes leads to a degree of cure of 80 %; therefore the

solidification is already at an advanced state. At this

level, if the temperature is increased to 140 °C a low

void content (1.3%) is generated on the paste adhe-

sive. Additionally, the curing time is shorter than 26

minutes.

Applying an initial temperature of 120 °C will shorten

the curing cycle compared to an initial temperature of

100 °C. In this case a degree of cure of 92.3 % is ob-

tained after the first heating step. The total curing time

is about 22 minutes. However, it will increase the void

formation to levels that might reduce its mechanical

performance.

An initial temperature of 140 °C can completely cure

the adhesive in less than 14 minutes. However, the

void content is 5.5 %, indicating a higher initial evapo-

ration of volatiles as stated at the hypotheses of Chap-

ter 4: Modeling of the Degradation Kinetics.


135

This test proves that the initial curing temperature affects crit-

ically the void formation during the curing process: When the

paste adhesive has a low degree of cure, high temperatures

will lead to an increase of void formation of the final paste

adhesive. For this reason, for the paste adhesive under study

the initial curing temperature must be limited to 100 °C.

7.3.2. Impact of on the curing process

According to the aforementioned results, [°C] is set to

100 °C. In this section the initial heating time is varied. The

curing in the second stage takes place at 140 °C. Results are

shown in Table 7.2.


136

Table 7.2: Effect of t1 [°C] on the curing process.

[°C

]

[min

]

[-]

[°C

]

[min

]

[min

]

[-]

𝑽𝒄

[-]

=𝟗𝟓

[min

]

=

𝟗𝟓

[min

]

𝑽𝒄𝟗𝟓

%

[-]

100

5 56

.3

140

5 14

.7

92.3

3.

9 7.

9 17

.6

4.6

100

5 56

.3

140

15

24.7

97

.6

5.2

100

10

72.2

14

0 5

19.7

93

.4

2.8

7 21

.7

3.0

100

10

72.2

14

0 15

29

.7

97.7

3.

2

100

15

79.9

14

0 5

24.7

94

.1

1.3

6.2

25.8

1.

4 10

0 15

79

.9

140

15

34.8

97

.9

1.5

100

20

84.2

14

0 5

29.7

94

.8

1.2

5.4

30.1

1.

2 10

0 20

84

.2

140

10

34.7

97

.0

1.2


137

The best curing strategy consists on increasing the curing

temperature over 100 °C when the degree of cure exceeds the

gelation point and reaches certain solidification. The reason is

that after this point, the paste adhesive can withstand the in-

crement of temperatures without major evaporation of vola-

tiles. Additionally, it is proved that a longer initial curing step

does not decrease significantly the void content, but increases

the processing time. In the case of the paste adhesive under

study, a recommendable degree of cure after the first heating

step is about 80 %, obtained after heating at 100 °C for 15

minutes.

7.3.3. Impact of the second heating stage on

the curing process

In this section, the curing temperature is varied between

140 °C and 180 °C and the duration of this curing stage is var-

ied between 0 and 20 minutes. equal to 0 means that curing

the process is completed after the second ramp-up. Results are

shown in Table 7.3.


138

Table 7.3: Effect of T2 [°C] on the curing process.

[°C

]

[min

]

[-]

[°C

]

[min

]

[min

]

[-]

𝑽𝒄

[-]

=𝟗𝟓

[min

]

=

𝟗𝟓

[min

]

𝑽𝒄𝟗𝟓

%

[-]

100

15

79.9

14

0 5

24.7

94

.1

1.3

6.2

25.8

1.

3 10

0 15

79

.9

140

15

34.7

97

.9

1.5

100

15

79.9

16

0 2

22.5

94

.5

2.6

2.4

22.9

2.

6 10

0 15

79

.9

160

5 25

.5

97.3

2.

9

100

15

79.9

18

0 0

21.3

92

.8

2.9

0.7

21.9

3.

0 10

0 15

79

.9

180

5 26

.3

98.8

3.

1


139

The results prove that a higher [°C] increases void content,

but affecting less to the void formation than an increase of

[°C]. The reason is the more advanced state of the chemical

reaction. These results confirm the conclusions of the TGA

analysis, showing that the effect of high curing temperatures

decrease with the cure progression. For this paste adhesive the

maximum curing temperature in the second heating stage is

140 °C considering a maximum void formation of 2 %.

7.3.4. Mechanical testing

Finally, SLS tests of CFRP plates bonded by induction heating

are carried out to validate the cure dependent heating ap-

proach. From Chapter 6, the samples with a void content low-

er than 2 % are expected to keep the maximum mechanical

performance. Besides, the samples with higher void content

are expected to show a lower mechanical performance. The

results of 5 samples tested for each curing cycle are summa-

rized in Table 7.4.

Additionally, Figure 7.2 shows the relation between mechani-

cal performance and void formation. Two regions are

sketched, describing the difference domains of failure of the

joints. Adherent failure occurs by the delamination of the sub-

strate, meaning that the adhesive is not fully loaded. Cohesive

failure is due to a failure of the paste adhesive. In this case it

means a degradation of the paste adhesive due to the high

temperatures applied.


140

Table 7.4: Summary of SLS test. S

amp

le

Cu

rin

g p

roce

ss

Sh

ear

stre

ng

th [

MP

a]

Fai

lure

mo

de

[

-]

𝑽𝒄 [

%]

[m

in]

1 10

0 °C

, 10

min

+ 1

60 °

C, 1

0 m

in

21.7

± 1

.9

25 %

Ad

her

. 75

% C

oh

98

.7

3.8

25.5

2 10

0 °C

, 15

min

+ 1

40 °

C, 1

0 m

in

26.2

± 0

.5

Ad

her

ent

96.7

1.

5 29

.7

3 10

0 °

C, 1

5 m

in +

160

°C

, 2.5

min

25

.2 ±

1.8

A

dh

eren

t 95

.0

2.6

23.0

4 10

0 °C

, 15

min

+ 1

60 °

C, 5

min

26

.4 ±

1.3

A

dh

eren

t 97

.3

2.9

25.5

5 12

0 °C

, 5 m

in +

160

°C

, 10

min

14

.6 ±

2.7

C

oh

esiv

e 98

.8

4.6

20.5

6 12

0 °C

, 10

min

+ 1

40 °

C, 5

min

20

.9 ±

4.2

25

% A

dh

er. 7

5 %

Co

h.

95.6

4.

3 19

.7

7 12

0 °C

, 10

min

+ 1

60 °

C, 1

0 m

in

19.5

± 2

.5

25 %

Ad

her

. 75

% C

oh

. 98

.9

4.1

25.5

8 12

0 °C

, 15

min

+ 1

60 °

C, 5

min

22

.9 ±

1.8

A

dh

eren

t 98

.1

3.7

25.5

9 14

0 °C

, 5 m

in +

180

°C

, 5 m

in

11.3

± 7

.9

Co

hes

ive

99.0

13

.4

16.3

R80

80

°C

, 240

min

24

.4 ±

1.5

A

dh

eren

t 95

.0

1.3

242.

3

R10

0 10

0 °C

, 60

min

22

.1 ±

2.1

A

dh

eren

t 93

.1

1.4

63.1

R12

0 12

0 °C

, 60

min

15

.2 ±

2.1

50

% A

dh

er. 5

0 %

Ad

hes

. 98

.6

3.9

63.9

R14

0 14

0 °C

, 15

min

9.

6 ±

3.4

Co

hes

ive

97.6

6.

1 49

.7

R16

0 16

0 °C

, 10

min

8.

1 ±

7.8

Co

hes

ive

98.6

15

.6

16.3


141

Figure 7.2: Impact of void formation on the shear strength.

The results of the mechanical performance test show that an

increment of the void formation decreases the mechanical per-

formance of the joint. The reason is that the increment of void

formation reduces the effective bonding area, thus reducing

the mechanical performance of the bonded joint.

Additionally, it shows that the standard deviation of the shear

strength is higher when the void formation is increased. This

mainly results from the fact that, in general, samples with low

void formation fail due to the delamination of the CFRP

plates. In contrast, samples with cohesive failure show a high-

er standard deviation of the mechanical performance of the

joint. The reason is that the failure is caused by the degrada-

tion of the paste adhesive, which can vary from sample to

sample.

As shown in the analysis of the dual step heating process, a

high initial curing temperature or a premature increase of

Adherent

1

2 3 4

5

68

9

R80

R100

R140

R1607

R120

0

5

10

15

20

25

30

0 5 10 15 20

Sh

ear

stre

ng

th [

MP

a]

Void content [%]

Cohesive / Mixed


142

temperature lead to a higher void formation, affecting the me-

chanical performance of the joint. This fact can be observed

comparing data sets 1-4, cured initially at 100 °C and 5-9, ini-

tially cured at 120 °C, showing a lower mechanical perfor-

mance. The reason is that the paste adhesive in a liquid state is

heated by higher temperatures, leading to fast evaporation of

volatiles. The temperature of the second curing stage can also

increase the void formation, despite being carried out after the

gelation point is reached. The reason is the evaporation of vol-

atiles if a too high temperature is used.

Applying the cure dependent heating process considered in

this investigation, the mechanical performance of the adhesive

cured under the recommended curing cycle (R80) can be

maintained. Sample 2 proves that the void formation can be

kept low and the curing of the paste adhesive can be accelerat-

ed by an increment of the curing temperatures when the gela-

tion point is reached.

7.4. Conclusions

In this Chapter, a dual step curing method is applied to accel-

erate the curing process of paste adhesives, consisting of two

dwell stages at different temperatures. It is proved that the

curing process must be carried out at low temperature until

the gelation point is reached to avoid void formation. For the

paste adhesive investigated, this point is about a degree of

cure of 80 %. Afterwards, the second dwell stage is used to

accelerate the curing process, without significantly increasing

the void content or reducing the mechanical performance of

the joint.

This approach is used to accelerate the curing cycle of the

paste adhesive under study, LME 10049-4 / LMB 6687-2 from

Huntsman Advanced Materials. For the best strategy found in

this Chapter, the second heating stage was approximately 15

7.4. CONCLUSIONS

143

°C higher than the final glass transition temperature ( = 126

°C) without significantly affecting the final void content. The

overall processing time is reduced from four hours for the rec-

ommended cure cycle to 30 minutes.

However, the dual step curing approach described in this

Chapter is a first step towards the complete process optimiza-

tion. In this chapter, the curing time is reduced by 87 % still

maintaining the mechanical performance of the joint. This re-

search opens a window for an optimization, as the curing time

could possibly be further reduced if additional curing stages

are considered, leading to a more complex but efficient pro-

cess.


144

145

Chapter 8

Process optimization

In this chapter, the optimization routine to accelerate the cur-

ing process of paste adhesives is described. This routine is

based on the thermal degradation model of Chapter 4: Model-

ing of the Degradation Kinetics. Afterwards, this solution is vali-

dated experimentally by measuring the void formation and

the mechanical performance of bonded joints.

8.1. Introduction

After analyzing the effect of high temperatures in the curing

process and modeling the cure kinetics and void formation,

the optimization of the curing process is carried out. The ob-

jective function under study is depending upon two nested

ODEs: The cure kinetics and the thermal degradation of the

paste adhesive. Additionally, the gradients are not available,

having a single objective function with real, scalar values that

might have local minima. In this context, the optimization al-

gorithm used in this dissertation is the CMA-ES (Covariance

Matrix Adaptation Evolution Strategy). This is a stochastic,

derivative-free method for optimization of non-linear prob-

lems [87-90]. More complex algorithms e.g. genetic algorithms

are not considered due to the relative simplicity of the prob-

lem, having a single objective.

In each one of the iterations, new candidates are evaluated

and the ones with higher fitness value are maintained for the

next generation. The fitness function is just the curing time,

http://en.wikipedia.org/wiki/Stochastic

http://en.wikipedia.org/wiki/Derivative

http://en.wikipedia.org/wiki/Numerical_optimization

http://en.wikipedia.org/wiki/Linear_map

CHAPTER 8: PROCESS OPTIMIZATION

146

aiming to be minimized. Additionally, boundary conditions

are considered in order to accept as candidates for the optimal

solutions those curing cycles completing the curing process

(α ≥ 95 %) with a void content lower than 1.5 %.

A standard deviation of the void content when applying the

optimal curing cycle lower than 0.5 % has been considered.

The reason is to ensure a final void formation always lower

than 2 %. The overall optimization process is shown in the box

diagram of Figure 8.1.

Figure 8.1: Optimization scheme of the curing process.

After the optimal candidate is calculated, it is validated con-

sidering two experimental testing: Microscopy and Single lap

Shear test (SLS). These tests must ensure the robustness of the

process with a high mechanical performance in the range of

the reference curing cycle. Additionally, a lower void for-

mation than 2 % must be guaranteed as defined in Chapter 6.

Next, the other paste adhesive considered in this investigation

is optimized, thus proving the repeatability of the process.

Void formation model

Initial values:

Ti, [K] ti [s]

T(t)T [K]

t [s]

Cure kinetics

Thermal degradation

Evolutionary algorithm

CMA-ES

Objective function:

• Reduce time

Constrains:

• Void content ≤ 1.5 %

• Degree of Cure ≥ 95 %

Void formation – thermal degradation

Optimal solution

Convergence

No Yes

0

2

4

6

8

10

12

14

16

0.0031 0.0032 0.0033 0.0034 0.0035 0.0036 0.0037 0.0038 0.0039

Vo

id c

on

ten

t [%

]


Deg

ree

of

cure

[-]

Time [min]

1.0

0.9

0.8

0.7

0.6

0.5

0.4

0.3

0.2

0.1

0.0

0 20 40 60 80 100


0.994

0.995

0.996

0.997

0.998

0.999

1

0 5 10 15 20 25 30

Rel

ativ

e m

ass

[-]

Time [min]

TGA 60 °C

Model 60 °C

TGA 80 °C

Model 80 °C

TGA 100 °C

Model 100 °C

TGA 120 °C

Model 120 °C

TGA 140 °C

Model 140 °C

TGA 160 °C

Model 160 °C

8.2. OPTIMIZATION OF THE LME 10049-4 / LMB 6687-2

147

8.2. Optimization of the LME 10049-4 / LMB

6687-2

After the analysis of more than 4’000 candidates, the algorithm

converges, meaning that the stop criterion is reached. It is set

to converge when the difference of the fitness function be-

tween two candidates evaluated consecutively is lower than

1e-2. In order to check the robustness of the solution found, the

program is run 10 times obtaining always the same results,


Figure 8.2: Optimal curing cycle of the LME 10049-4 / LMB 6687-2.

The optimal curing cycle accelerates the processing time

89.8 % compared to the reference curing cycle. For the valida-

tion, 26 samples are produced following the details given in

Appendix B.4. The void content is shown in Table 8.1.

Table 8.1: Microscopy test of the optimal curing cycle.

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0

20

40

60

80

100

120

140

160

0 5 10 15 20 25

Deg

ree

of

cure

[-]

Tem

per

atu

re [

ºC]

Time [min]

Temperature [ºC]

Degree of cure [-]


148

Sample Vc [%] Sample Vc [%] Sample Vc [%]

1 1.58 ± 0.20 10 1.59 ± 0.31 19 0.91 ± 0.43

2 1.77 ± 0.35 11 1.54 ± 0.57 20 1.58 ± 0.52

3 1.55 ± 0.39 12 1.49 ± 0.62 21 1.24 ± 0.47

4 1.53 ± 0.33 13 1.52 ± 0.40 22 1.17 ± 0.40

5 1.72 ± 0.19 14 1.31 ± 0.62 23 1.41 ± 0.36

6 1.87 ± 0.30 15 1.81 ± 0.20 24 0.98 ± 0.36

7 1.78 ± 0.41 16 1.99 ± 0.25 25 0.63 ± 0.09

8 1.21 ± 0.64 17 1.73 ± 0.24 26 1.41 ± 0.78

9 1.45 ± 0.47 18 1.56 ± 0.57

The results validate the optimal curing cycle, having an aver-

age void formation of 1.48 ± 0.47 %. The standard deviation

measured is in agreement with the predicted value of 0.5. This

means that the optimal curing cycle ensures an average void

formation for all samples lower than 2 %.

Additionally, 8 SLS samples are mechanically tested following

the Appendix A.3. The results are summarized in Table 8.2.

8.2. OPTIMIZATION OF THE LME 10049-4 / LMB 6687-2

149

Table 8.2: SLS test of the optimal curing process.

Overlap

[mm]

Bondline

thickness

[mm]

Width

[mm]

Force

[N]

Shear

Strength

[MPa]

Fracture

mode

14.7 0.15 22.6 8’133 24.6 Adherent

14.1 0.20 22.6 7’566 23.7 Adherent

14.5 0.19 22.6 8’084 24.7 Adherent

14.5 0.19 22.6 8’091 24.8 Adherent

14.2 0.23 22.3 7’849 24.8 Adherent

14.2 0.19 22.5 7’603 23.9 Adherent

14.0 0.19 22.9 8’041 25.1 Adherent

14.4 0.19 21.9 7’842 25.0 Adherent

The mechanical performance is compared to the reference cur-

ing cycle. The results of SLS test are shown in Figure 8.3.

Figure 8.3: Recommended and optimal curing cycles of the LME 10049-4 /

LMB 6687-2.

0

5

10

15

20

25

30

4h at 80 °C Optimal cycle

Sh

ear

stre

ng

th [

MP

a]


150

It is proved that the original mechanical performance is main-

tained with the optimized curing cycle and the processing

time is reduced 90 %.

8.3. Optimization of the LME 10625 / LME

10626

In order to prove the repeatability of the process, the same

methodology is followed with the paste adhesive system LME

10625 / LME 10626. The results of the optimization routine are


Figure 8.4: Optimal curing cycle of the LME 10625 / LME 10626.

This cycle can be compared to the reference curing cycle, es-

tablished at 2 hours at 80 °C. The processing time is reduced

87.2 %, achieving 95 % of degree of cure. The optimal curing

cycle is validated by measuring the void formation in 4 sam-

ples and 5 measurements each. Additionally, the mechanical

performance with 6 SLS samples according to EN 2243-1. The

results of the microscopy are summarized in Table 8.3 and the

SLS test in Table 8.4.

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0

20

40

60

80

100

120

140

160

180

0 5 10 15 20

Deg

ree

of

cure

[-]

Tem

per

aure

[ C

]

Time [min]

Temperature [°C]

Degree of cure [-]

8.3. OPTIMIZATION OF THE LME 10625 / LME 10626

151

Table 8.3: Microscopy of samples cured with the optimal curing cycle.

Sample Void content [%]

1 1.41 ± 0.71

2 1.84 ± 0.35

3 1.63 ± 0.21

4 1.40 ± 0.53

The void content proves that the optimal curing cycle has a

void content lower than 2 %, being equal to 1.57 ± 0.5., ensur-

ing the quality of the paste adhesive.

Table 8.4: SLS test of the optimal curing process.

Overlap

[mm]

Bondline

thickness

[mm]

Width

[mm]

Force

[N]

Shear

Strength

[MPa]

Fracture

mode

13.2 0.16 25.1 8’906 26.88 Adherent

13.2 0.16 25.1 8’724 26.37 Adherent

13.2 0.06 24.9 7’277 22.14 Adherent

13.2 0.11 25.1 8’274 24.95 Adherent

13.2 0.14 25.1 9’252 27.95 Adherent

13.2 0.10 25.1 8’432 25.45 Adherent

Finally, the mechanical performance is compared to the refer-

ence curing cycle (2 hours at 80 °C with the oven) measured

by Huntsman Advanced Materials (TA 211-03-02-02 of Clean

Sky JTI). The average value for 6 samples tested is 24.38 ± 1.2.

The results are shown in Figure 8.5.


152

Figure 8.5: Comparison between recommended and optimal curing cycles of

the LME 10625 / LME 10626.

Also in this case, the mechanical performance of the optimal

curing cycle is maintained if compared to the reference curing

cycle.

8.4. Conclusions

In this chapter, the curing process of paste adhesives is opti-

mized with respect to the time using the models developed

thorough this doctoral dissertation. The process is carried out

for the two paste adhesive systems under study. The solutions

are experimentally validated, proving that the mechanical per-

formance of the joint is maintained if compared to the recom-

mended curing cycles.

0

5

10

15

20

25

30

2h at 80 °C Optimal cycle

Sh

ear

stre

ng

th [

MP

a]

153

Part IV

Demonstrator

155

Chapter 9

Design and assembly of the de-

monstrator

This chapter discusses the challenges when applying induc-

tion heating to complex structures. Additionally, the design

and the assembly process of the technology demonstrator

used to validate the research of this doctoral thesis are de-

tailed.

9.1. Introduction

The research in previous chapters considers SLS samples to

validate the optimization of the curing process of paste adhe-

sives. In a real environment, a number of phenomena are af-

fecting the temperature generated in the bondline:

Edge effect: Heating a long component may lead to a

temperature decrease at the edge of the structure. The

main reasons are a coil efficiency decrease combined with

a variation of the convective cooling at the edges of the

structure. This effect must be considered in order to avoid

that the bondline at the edges of the structure are under

cured.

Variable laminate thickness: Heating a CFRP susceptor

requires the accuracy of the laminate thickness to guaran-

tee an even heating of the component.

Inhomogeneity of the convection coefficient: Complex

shapes may lead to a variation of the air flow due to

CHAPTER 9: DESIGN AND ASSEMBLY OF THE DEMONSTRATOR

156

holes, closed chambers and changes of the shape of the

structure. This may affect to the convective cooling.

Distance between coil and the susceptor: The heat in-

duced depends on this distance. In some parts it might

happen that this distance varies, leading to bondline lo-

cally under cured if the distance increases or locally with

a higher void formation if the distance decreases.

Accessibility of the temperature sensors: In complex

structures, the access for the temperature sensors is some-

times limited. For this reason, the calibration of the sensor

positioning is a crucial step.

In this context, the challenge when bonding a real component

is to achieve a homogeneous temperature distribution at the

bondline. For this reason, the implications of considering a

real 3 dimensional component are investigated, discussing an

approach to ensure a correct curing process.

Afterwards, the methodology described in this investigation

to accelerate the curing process of paste adhesive is applied to

a real component. The geometry selected to demonstrate the

research findings is the aileron of the Dornier 228. This com-

ponent is described and validated regarding the geometric

technical specifications of the Clean Sky JTI.

9.2. Procedure for bonding complex parts

During the assembly process of complex structures, the effects

described in the previous section might lead to a temperature

difference inside the bondline. As a consequence, if the opti-

mal curing process is applied, there could be sports where ei-

ther the curing process is not completed or the void formation

is higher than 2 %. In order to tackle this problem, the ap-

proach followed summarized in Figure 9.1 is followed.

9.2. PROCEDURE FOR BONDING COMPLEX PARTS

157

Figure 9.1: Approach to optimize the curing cycle in complex structures.

Firstly, an accurate temperature distribution analysis at the

bondlines must be performed. One option is to simulate the

curing process, as described in Chapter 5: Simulation of the cur-

ing process of paste adhesives by induction heating. With an ap-

propriate tool, the temperature difference can be predicted.

The accuracy of a thermal model is depending mainly on the

complexity of the part and the accuracy of the convective cool-

ing. In this sense, a complex part will have a more complex air

flow being more difficult to model and therefore less accurate.

In this dissertation, some guidelines are provided in Appendix

B.3 to improve the model accuracy.

Additionally, a proper experimental calibration is necessary to

ensure that the desired temperature is applied. For this reason,

the temperature distribution is accurately characterized at

each bondline obtaining:

The temperature difference between the measuring

spot and the bondline: The pyrometer measures the

temperature at the surface of the susceptor. In some

cases, especially if the laminate is thick, higher tem-

peratures are generated in the bondline. This value is

used to calibrate the temperature sensors.

Analysis of thetemperaturedistribution

Optimizationof curing cycle

of complexstructures

T(t) ∆Tbondline 0 °C

Chapter 10

Optimizationof curing cycle

of simple structures

T(t) ∆Tbondline= 0 °C

Cure kineticsmodel

α = f (T,t)

Thermal degradation

model

= f (T,t)

Voidformation

modelVc = f (T,t)

Optmization for simple structures

Optmization for complex structures

Simulation Experimental


158

The temperature difference existing inside of the

bondline: This value must be considered as input for

the optimization of the curing cycle for complex ge-

ometries.

The optimization must define two curing profiles linked by

the temperature difference measured. The boundaries are

modified in order to ensure that:

The warmest profile does not lead to a void formation

higher than 1.5 %.

The coldest profile is completely cured, α ≥ 95 %.

9.3. Description of the demonstrator

The geometry used as technology demonstrator in this re-

search is the aileron of the Dornier 228. It is modified in col-

laboration with structural engineers from RUAG, a project

partner of Clean Sky JTI. The objective is to adapt today’s ge-

ometry of the aileron, made of aluminum, to a new design

having CFRP parts bonded by induction heating. The real

component is shown in Figure 9.2 [91].

Figure 9.2: Image of the Dornier 228.

Aileron

9.3. DESCRIPTION OF THE DEMONSTRATOR

159

The final aileron proposed for this doctoral thesis is made of

12 CFRP parts, as detailed in Figure 9.3. The material used is

the CFRP detailed in Appendix A.1.

Figure 9.3: Geometry of the aileron of the Dormier 228.

A detail of the rear part consisting of two doublers and the C-

spars is shown in Figure 9.4.

Figure 9.4: Detail of the rear part of the aileron.

The layup sequence of the parts was determined by RUAG,

aiming to keep the mechanical performance of the aluminum

Upper C-spar

Lower C-spar

Rib 1

T- stringer

L- stringer

Upper skin

Lower skin

Outer doubler

Inner doubler

Rib 2

Rib 3

Rib 4

Doubler

End profile

Left edge

Right edge

Upper C-spar

Lower C-spar

Outer doubler

Inner doubler


160

original parts. The layup consists of [± 45, 0/90, ± 45] for all the

parts, excepting for one side of the skins and for the C-Spar,

with a layup of [± 45, 0/90, ± 45, ± 45, 0/90, ± 45]. A detail of the

stepwise layup used in one side of the skins is shown Figure

9.5.

Figure 9.5: Layup sequence of the side of the skins.

All the CFRP parts are produced following the Section A.2.1.

Production of CFRP laminates.

9.4. Assembly process

The assembly sequence consists of 9 steps with 21 bondlines,

summarized in Table 9.1. The adhesive and adherents are pre-

pared as detailed in Appendix A.1. Glass beads of 0.1 mm are

used and pressure is applied during the curing process. The

reason is to ensure a bondline thickness between 0.1 and 0.3

mm.

The optimal curing cycle of the paste adhesive system LME

10049-4 / LMB 6687-2 is applied to all the bondlines. This cur-

ing cycle is selected, although it is predicted that some of the

bondlines might have a higher void content than expected.

The reason is the temperature differences between the measur-

ing spot and the maximum temperature inside of the

bondline. This point is addressed in detail in Chapter 10: Anal-

ysis of the demonstrator.

9.4. ASSEMBLY PROCESS

161

The temperature is controlled with the IR pyrometer [IN 510,

LumaSense] and with the optical fiber pyrometer for high

temperature measurements [FTP-NY2-ST2-2M-BA, Photon

Control, Canada]. The positioning is detailed at each assembly

step. The use of fans to create a forced convection cooling sce-

nario is not considered due to the complexity of the bondlines.

However, its use is discussed in Chapter 10: Analysis of the de-

monstrator. The maximum temperature deviation measured

during the process is always kept under 3 °C compared to the

curing cycle defined in Section 8.2: Optimization and validation

of the LME 10049-4 / LMB 6687-2.

Table 9.1: Summary of the bondlines of the demonstrator.

Bondline

1 T-str. made of 2 L str. 12 Rib 1 to lower skin

2 T-str. to lower skin 13 Rib 2 to lower skin

3 L str. to upper skin 14 Rib 3 to lower skin

4 Rib 1 to inner doubler 15 Rib 4 to lower skin

5 Rib 2 to inner doubler 16 C-spar to inner doubler

6 Rib 3 to inner doubler 17 C-spar to lower skin

7 Rib 4 to inner doubler 18 C-spar to inner doubler

8 Rib 1 to upper skin 19 C-spar to upper skin

9 Rib 2 to upper skin 20 Outer d. to upper C-spar

10 Rib 3 to upper skin 21 Outer d. to lower C-spar

11 Rib 4 to upper skin

The bonding process is carried out by pressing the compo-

nents to be joined with the coil, as shown in Figure 9.6. GFRP

spacers of 1 mm thickness are placed between the coil and the

susceptor to ensure a constant distance.


162

Figure 9.6: Scheme of the procedure to heat the components.

Firstly, the T-stringer is bonded from two L-stringers. The

temperature is measured at the center of the stringer, as

shown in Figure 9.7, using the coil to press the parts together.

Figure 9.7: Assembly of the T-Stringer

Then, the T-stringer is bonded to the lower skin and the L-

stringer to the upper skin, following the same approach. An

image of the process is shown in Figure 9.8, with both pyrom-

eters measuring the temperature at the surface under the skin

in the central area of the stringer.

Coil

Bonding partners

Spacers

F

Assembly rig

Pyrometer measuring spot


163

Figure 9.8: Assembly of the T-stringer to the lower skin.

Afterwards, the 4 ribs are bonded to the inner doubler. The

temperature is measured at the center of the area to be bond-

ed, in the internal surface of the rib. The ribs are clamped with

a sandwich structure, shown in Figure 9.9, made of non-

electrical conductive materials to avoid the heating by induc-

tion. This structure is used to fix the components during the

curing process and to ensure a bondline thickness lower than

0.3 mm.

Figure 9.9: Structure to fix the parts during the process.



164

Once the paste adhesive is placed at the 4 bondlines and the

components are pressed together, the curing process is carried

out, as shown in Figure 9.10.

Figure 9.10: Bonding process of a rib and the inner doubler.

Afterwards, the upper skin is bonded to the ribs as shown in

Figure 9.11. The temperature is measured at the internal sur-

face of the rib, at about 5 mm from the doubler.

Figure 9.11: Upper subassembly bonded to the ribs.




165

The resulting subassembly is shown in Figure 9.12.

Figure 9.12: Subassembly with the ribs and the upper skin.

Next, the lower skin is bonded to the ribs. In this bonding pro-

cess, a shorter coil is used with a curvature to ensure the even

heating along the bondline, as shown in Figure 9.13. The tem-

perature is measured under the bonded area of the rib, at

about 5 mm from the doubler.

Figure 9.13: Assembly of the lower skin to the ribs.



166

Afterwards, the C-spars are bonded. This process requires

bonding both faces of the C-spar in a row, as shown in Figure

9.14 and Figure 9.15. The temperature is measured under the

inner doubler, at the region of the edge where the access is

easier.

Figure 9.14: Assembly of a C-spar to the inner doubler.

Figure 9.15: Assembly of the C-spar to the lower skin.




167

Finally, the outer doubler is bonded to the C-spars having 2

additional bondlines, as shown in Figure 9.16. The tempera-

ture is measured under the inner doubler at one of the edges.

Figure 9.16: Outer doubler bonded to the subassembly.

The resulting component is shown in Figure 9.17



168

Figure 9.17: Demonstrator.

9.5. Evaluation of the aileron

After the assembly, the geometrical tolerances are measured

according to Table 9.1. The results are summarized in Table

9.2 for distances, Table 9.3 for the orientations of the parts and

Table 9.4 for the bondline thicknesses. The maximum devia-

tion allowed according to the technical specifications provided

by RUAG is 1 mm for the positioning and 1° for the orienta-

tion of the parts. Additionally, the maximum bondline thick-

ness allowed is 0.3 mm. All the measures are carried out 5

times at different points along the parts.

9.5. EVALUATION OF THE AILERON

169

Table 9.2: Validation of the positioning of the parts where the maximum devi-

ation according to RUAG specifications are 1 mm.

Distance measured [mm] CAD

[mm]

Deviation

[mm]

Left edge to rib 1 77.55 ± 0.42 77.5 0.05

Rib 1 width 20.12 ± 0.13 20 0.12

Rib 1 to rib 2 142.58 ± 0.27 143 0.42

Rib 2 width 20.26 ± 0.05 20 0.26

Rib 2 to rib 3 142.96 ± 0.24 143 0.04

Rib 3 width 19.99 ± 0.15 20 0.01

Rib 3 to rib 4 143.19 ± 0.13 143 0.19

Rib 4 width 20.13 ± 0.22 20 0.13

Rib 4 to right edge 65.48 ± 0.32 65.5 0.02

Doubler to L-stringer 137.19 ± 0.13 136.42 0.77

End profile to L- stringer 139.76 ± 0.03 140.17 0.41

Doubler to T-stringer 157.08 ± 0.12 156.67 0.41

End profile to T-stringer 115.3 ± 0.26 115.48 0.18

L stringer width 10.02 ± 0.05 10 0.02

T stringer width 12.22 ± 0.10 12 0.22

Doubler to Lower skin 10.96 ± 0.23 11 0.04


170

Table 9.3: Validation of the orientation of the parts where the maximum value

according to RUAG specifications are 1°.

Part Angle [°] Part Angle [°]

Rib 1 - left edge 0.21 T-stringer – End profile 0.43

Rib 2 - left edge 0.13 L-Stringer – End profile 0.02

Rib 3 - left edge 0.06 Doubler – End profile 0.12

Rib 4 - left edge 0.16

9.5. EVALUATION OF THE AILERON

171

Table 9.4: Validation of the bondline thickness where the maximum value

according to RUAG specifications are 0.3 mm.

Bondline Thickness

[mm]

L stringer + L stringer to

form the T-stringer

0.10 ± 0.03

T-stringer to lower skin 0.12 ± 0.04

L stringer to upper skin 0.11 ± 0.02

Rib 1 to inner doubler 0.12 ± 0.03




Rib 1 to upper skin 0.18 ± 0.02




Rib 1 to lower skin 0.18 ± 0.05




Lower C-spar to inner dou-

bler

0.13± 0.03

C-spar to lower skin 0.15± 0.03

Upper C-spar to inner dou-

bler

0.20 ± 0.08

C-spar to upper skin 0.17 ± 0.03

Outer doubler to upper C-

spar

0.15 ± 0.05

Outer doubler to lower C-

spar

0.19 ± 0.04


172

The results show that the parts are well positioned, having a

low deviation compared to the CAD geometry. The deviations

are lower than the requirements set for qualification, thus val-

idating the assembly rig used to fix the parts during the bond-

ing process.

9.6. Conclusions

Induction heating of complex structures requires considering

some effects that might lead to an inhomogeneous tempera-

ture distribution at the bondlines. The accurate of prediction

this phenomenon is critical to ensure the good quality of the

joint.

This chapter defines a strategy to deal with this challenge and

therefore guarantee a robust joining process for complex struc-

tures. This approach is based on the correct modeling of the

thermal process by simulation. With this model, the tempera-

ture distribution can be predicted and considered in the opti-

mization process of the curing process of paste adhesives.

Finally, the complex geometry considered in this dissertation

to validate the methodology investigated. The assembly pro-

cess is described and the test rig is validated considering the

technical specifications of the project.

173

Chapter 10

Analysis of the demonstrator

In this chapter, a thorough temperature analysis is carried out

to generate input data for the optimization of the process.

Then, the curing cycles are corrected according to the tempera-

ture distribution in the bondlines and used to produce a sec-

ond demonstrator.

10.1. Introduction

In this chapter, the temperature difference is experimentally

measured during the heating process with natural convection.

The objective of this investigation is to generate the necessary

input data for the optimization to consider the bondline tem-

perature distribution. With this purpose, all the bondlines of

the demonstrator are divided in groups representing the heat-

ing conditions of all the bondlines. The 21 bondlines are divid-

ed into 4 bondlines types having different heating scenarios

(see A to D, Figure 10.1).

CHAPTER 10: ANALYSIS OF THE DEMONSTRATOR

174

Figure 10.1: The 4 representative bondlines.

Then, the temperature distribution at the bondlines is meas-

ured. At this point, the reasons for an inhomogeneous temper-

ature distribution are investigated, discussing potential design

alternatives to minimize them.

Afterwards, the measured values of the temperature differ-

ence in the bondlines are used as input data in the optimiza-

tion of the curing process. Additionally, the limits for the ap-

plicability of the optimization tool are discussed.

The cycles are optimized considering the temperature distri-

bution. Then, they are applied in the bonding process of a sec-

ond demonstrator. Finally, the void formation and the degree

of cure are measured at all the bondlines, validating the ro-

bustness of the process investigated.

1. L+L to make a T2. T to upper skin3. L to lower skin4. Rib 1 to Inner doubler5. Rib 2 to Inner doubler6. Rib 3 to Inner doubler7. Rib 4 to Inner doubler8. Rib 1 to Upper skin9. Rib 2 to Upper skin10. Rib 3 to Upper skin11. Rib 4 to Upper skin12. Rib 1 to Lower skin13. Rib 2 to Lower skin14. Rib 3 to Lower skin15. Rib 4 to Lower skin16. C- spar to Inner doubler lowerside17. C- spar to lowerskin18. C- spar to Inner doubler upper side19. C- spar to upper skin20. C- spar to outer doubler upper side21. C- spar to outer doubler lower side

A

B

C

D

10.2. TEMPERATURE DISTRIBUTION ANALYSIS WITH NATURAL

CONVECTION.

175

10.2. Temperature distribution analysis with

natural convection.

To measure the temperature distribution inside of the

bondline, small holes with a diameter of 2 mm are drilled at

different distances. Then, they are filled with thermal conduc-

tive paste in order to stabilize the temperature during the

measurement. Afterwards, the temperature distribution is

measured with the optical fiber pyrometer for high tempera-

ture measurements [FTP-NY2-ST2-2M-BA, Photon Control,

Canada]. The current applied for all the measurements is 55.2

A, with a frequency of 198’000 Hz for the longer coil and

270’000 Hz for the curved coil.

The temperature is measured in each case at the center of the

bondline and in the surface of the adherent. The reason is to

measure the differences between the spot where the tempera-

ture is measured and the real temperature inside of the

bondline. At this point, the optimal positioning of the sensor is

defined according to accessibility and good temperature dis-

tribution.

Additionally, the maximum temperature difference inside of

the bondline is measured in order to redefine the optimal cur-

ing cycle. In this section, the measurements are carried out

always 5 times and the average is calculated and shown for

the different bondlines. The standard deviation cannot be ob-

served in the graphs because the values obtained are always

lower than 1.0 °C.

The temperature distribution analysis in Sections 10.2.1 to

10.2.4 is obtained at a constant current. However, the optimal

curing cycle considers a wide temperature range where the

temperature difference may vary depending on the current

applied. In this context, Section 10.2.5 addresses this issue.


176

10.2.1. Bondline type A between upper skin and ribs.

The temperature distribution is measured inside the bondline

between rib 1 and the upper skin (see Figure 9.3) at 9 points, 5

times each. The measurement is carried out inside of the

bondline and under the bondline, as shown in Figure 10.2.

Figure 10.2: Measurement in bondline A.

Figure 10.3: Temperature distribution in bondline A.

The maximum temperature difference in the bondline

is 6.8 °C. This value used to re-define the optimized

curing cycle.

20

30

40

50

60

70

80

90

0 5 10 15 20 25 30

Tem

per

atu

re [ C

]

Measuring point x [cm]

Under bondline

Inside bondline


CONVECTION.

177

The temperature difference between the measurement

spot and the temperature inside of the bondline is 5.7

°C in x= 20 cm.

10.2.2. Bondline type B between lower skin and ribs.

The measurement is carried out again for the curved bondline

between the lower skin and the rib 1, also 5 times each as


Figure 10.4: Measurement in bondline B.

Figure 10.5: Temperature distribution in bondline B.

The temperature distribution shows a maximum difference

about 25 °C. However, between x = 10 cm and x= 13 cm there

20

40

60

80

100

120

140

0 5 10 15 20 25 30

Tem

per

atu

re [ C

]


Inside bondline

Under bondline


178

is no bondline between bibs and skin due to the stringer. For

this reason, these points are not considered.



curing cycle.



°C in x= 20 cm.

10.2.3. Longitudinal bondlines type C.

All the bondlines where the entire coil length is used are

summarized by this case. As a result, the bondline between

the c-spar and the lower skin is measured. The spots where

the temperature is measured 5 times each are shown in Figure

10.6 and the temperature distribution in steady state is shown

in Figure 10.7.

Figure 10.6: Measurement in bondline C.


CONVECTION.

179

Figure 10.7: Temperature distribution in bondline C.



curing cycle.



°C in x= 54 cm, besides rib 1.

10.2.4. Bondline type D between inner doubler and ribs.

Finally, the vertical bondlines between ribs and doubler are

analyzed in the points marked in Figure 10.8, 5 times each.

The bondline measured is the one between the inner doubler

and the rib 1. The temperature distribution is also shown in

Figure 10.9.

2030405060708090

100110120

0 10 20 30 40 50 60

Tem

per

atu

re [

°C]


Inside bondline

Under bondline


180

Figure 10.8: Measurement in bondline D.

Figure 10.9: Temperature distribution in bondline D.



curing cycle.



°C in x= 5 cm.

0

10

5

2.5

7.5

x

20

30

40

50

60

70

80

90

0 2 4 6 8 10

Tem

per

atu

re [ C

]


Inside bondline

Under bondline


CONVECTION.

181

10.2.5. Temperature distribution with a variable current

This section investigates the accuracy of the temperature dif-

ferences when a different current is applied. In this context,

the temperature distribution at bondlines type C is measured,

varying the current applied from 50.4 A to 60 A. These values

are chosen because they lead to a generation of temperatures

between 80 and 130 °C. In this temperature range most of the

chemical reaction takes place according to the optimal curing

cycle defined in Chapter 8. The results are summarized in Ta-

ble 10.1 for the temperature difference in the bondline and in

Table 10.2 for the temperature difference in the measuring

spot.

Table 10.1: Temperature difference inside of the bondline.

Current [A] 50.4 A 55.2 A 60A

Max. temperature at the

bondline (°C) 101.1 ± 0.6 112.9 ± 0.2 126.8 ± 0.3

Min. temperature at the

bondline (°C) 82.6 ± 0.4 93.3 ± 0.4 107.6 ± 0.6

∆T at the bondline (°C) 18.4 19.6 19.2

The results show a higher temperature generated when the

current is increased. The temperature difference varies when

applying difference currents about 1 °C. In this context, it is

assumed that the temperature distribution is not affected by a

current change in the range of values used in this investiga-

tion.

Table 10.2: Temperature difference in the measuring point, x= 54 cm.


182

Current [A] 50.4 A 55.2 A 60A

Max. temperature at the

bondline (°C) 101.1 ± 0.6 112.9 ± 0.2 126.8 ± 0.3

Temperature at the meas-

uring point (°C) 96.02 ± 0.2 107.9 ± 0.3 120.5 ± 0.2

∆T at the measurement

point (°C) 5.1 5.2 5.6

These results show that the temperature difference between

the measuring spot and the temperature inside of the bondline

increases slightly with a higher current applied. However, the

differences measured are inside the values of standard devia-

tion of the measurements. As a consequence, the temperature

difference between affecting the pyrometer is considered con-

stant in this investigation.

10.3. Temperature distribution in complex

shapes heated by induction

In Chapter 9: Design and assembly of the demonstrator, a number

of effects that could influence on the temperature distribution

generated by induction heating are described. This section

analyzes each of them for the demonstrator under study.

10.3.1. Edge effect

When heating the edge of a component, a temperature de-

crease is noticed. In order to analyze this effect, a bondline

type C (see Figure 10.1) is heated varying the offset of the coil

with respect to the susceptor. The temperature is then meas-

ured at the edge of the bondline. A scheme of the experiment

and the results are shown in Figure 10.10.

10.3. TEMPERATURE DISTRIBUTION IN COMPLEX SHAPES HEATED BY

INDUCTION

183

Figure 10.10: Edge effect measurement.

The results prove that the edge effect is caused by a combina-

tion of an increase of heat dissipation of the component and a

reduction of the efficiency of the coil at the edges. In this

sense, having the coil significantly longer compared to the

component reduces remarkably this effect. The results show a

temperature difference of about 20 °C with an offset of 25 mm

and of 10 °C with 100 mm. However, this phenomenon is still

cannot be completely corrected by increasing the offset be-

tween the coil and the susceptor. The reason is the higher sur-

face free area at the edge, leading to a higher cooling.

10.3.2. Effect of the laminate thickness variation

The thickness variation of the laminate affects to the electrical

conductivity of the laminate and therefore to the induced en-

ergy.

80

85

90

95

100

105

110

115

52 54 56 58 60 62 64

Tem

per

atu

re [

°C]


25

50

100

Offset [mm]:


184

An example of this effect is illustrated in Figure 10.11. In this

case, the measurement at x = 11 cm is carried out in a thicker

region due to the added thickness from the t-stringer. As a

consequence, the temperature generated is slightly higher at x

= 11 cm than at x= 10 cm and x= 13 cm.

Figure 10.11: Thickness variation effect.

This effect can be predicted by a simulation tool considering a

proper modeling of the electrical conductivity of the different

regions.

10.3.3. Effect of the convective flow

In order to discuss the use of forced convection in complex

structures, the measurements of Section 10.2 are repeated with

forced convection cooling. In this context, a homogeneous

convection flux of 2 m/s is generated at the beginning of the

part, as detailed in Appendix B.2. The position of the fans is

changed depending on the bondline analyzed, as shown in

Figure 10.12. The temperature gradients for each bondline

20

40

60

80

100

120

140

0 5 10 15 20 25 30

Tem

per

atu

re [ C

]


Inside bondline

Under bondline

10.3. TEMPERATURE DISTRIBUTION IN COMPLEX SHAPES HEATED BY

INDUCTION

185

type (see Figure 10.1) in steady state are measured and sum-

marized in Table 10.3.

Figure 10.12: Fan positioning of forced convection cooling.

Table 10.3: Temperature gradients with different cooling.

Bondline

type

Forced conv.

direction

Forced conv.

Max. ∆T [°C]

Natural conv.

Max. ∆T [°C]

A Sideward 8.6 6.8

B Sideward 25.8 13.9

C Sideward 45.1 19.6

C Backward 23.0 19.6

C Frontward 15.7 19.6

D Sideward 13.4 14.7

The use of forced convection with complex shapes leads to an

irregular air flow meaning a nonhomogeneous cooling. Only

at the vertical bondline cooled sideward and at long bondline

cooled frontward a lower temperature difference is measured

compared to natural convection. The main reason is the dis-

turbances of the laminar air flow generated by the section

changes. As the coil is used to press the joints during the pro-

cess, the air cannot flow between the susceptor and coil, lead-


186

ing to a poor and irregular cooling. As a consequence, the use

of forced convection leads to higher temperature gradients

and a less controllable process.

10.3.4. Effect of the distance between the coil and the sus-

ceptor

The distance between the coil and susceptor is more or less

constant in all the bondline areas analyzed because the coil is

used to press the parts together. This ensures a minimum dis-

tance between coil and susceptor. However, there are some

cases where this contact is not maintained, for instance, in the

heating of the bondlines type A and B (see Figure 10.1). In

those cases, the distance increases at the edges, being the heat

generated in the susceptor lower compared to the regions with

lower distance between the coil and the susceptor. One exam-

ple is shown in Figure 10.13.

Figure 10.13: Effect of the distance between the coil and the susceptor.

This effect can be minimized by a proper coil design that

keeps constant the distance between the coil and the suscep-

tor.

20

40

60

80

100

120

140

0 5 10 15 20 25 30

Tem

per

atu

re [ C

]


Inside bondline

Under bondline

10.4. MODIFICATION OF THE OPTIMAL CURING CYCLE

187

10.4. Modification of the optimal curing cy-

cle

The temperature differences obtained in the analysis of the

temperature distribution are summarized in Table 10.4. The

the optimal location for the measurement is included consider-

ing accessibility and edge effects.

Table 10.4: Temperature distribution analysis summary.

∆T in the

bondline (°C)

∆T between bondline and

measurement point (°C)

Bonline type A 6.8 5.7 (x= 20 cm)

Bonline type B 13.9 4.1 (x= 20 cm)

Bonline type C 19.6 5.2 (x= 54 cm)

Bonline type D 14.7 11.7 (x= 5 cm)

In this context, the optimization tool must consider the value

of the ∆T in the bondline as an input. For this reason, the op-

timization tool of Chapter 8: Process optimization, is modified in

order to optimize 2 temperature profiles (named “h” an “l”

referring to higher and lower temperature respectively). These

profiles have a constant temperature difference (∆T at the

bondline) during the whole process. As a consequence, the

optimization tool provides the optimal solution considering:

Minimum time.

The lower temperature cycle must be completed

(α ≥ 95 % ).

The higher temperature cycle must not have a void

formation higher than 1.5 %.


188

As the measured values of the temperature difference at the

bondline of the investigated demonstrator reach almost 20 °C,

the input temperature is varied until 25 °C in Figure 10.14.

Figure 10.14: Optimum curing cycle for a ∆T in the bondline between 5 and

25 °C.

The curing cycle is adapted with a longer time at higher tem-

perature. The reason is that the model considers that there is

no further degradation when a degree of cure of 95 % is

reached, as mentioned in Chapter 3: Modeling of the Degrada-

tion Kinetics.

However, the tool is not allowed to increase the curing tem-

perature after reaching a degree of cure of 95 %. This is to

avoid the increase of temperatures at spots in a region close to

the spot of maximum temperature. These spots could have

still a degree of cure lower than 95 % that could increase the

void formation. For this reason, the temperature does not in-

crease after the curing process is completed.

Additionally, the applicability of this optimization is investi-

gated. This tool can be used until a point where the tempera-

0

20

40

60

80

100

120

140

160

0 5 10 15 20 25 30 35 40 45 50

Tem

per

atu

re [

°C]

Time [min]

Temperature difference = 5 °C Lower temp.





Temperature difference = 5 °C Higher temp.






189

ture differences generated by induction are too high and the

methodology defined in this dissertation cannot be applied.

In order to obtain this limit, the optimization tool is run for

different values of ∆T until the solution does not converge an-

ymore. The results of the total curing time of the difference

solution with respect to the ∆T used are shown in Figure 10.15.

Figure 10.15: Duration of the optimal curing cycle with a temperature differ-

ence.

It is shown that for values until 35 °C, the temperature cycle

can be readapted. However, it the temperature difference is

higher the process defined in this research cannot be applied.

The reason is that the bondline would be either locally de-

graded in the spot of higher temperature or locally not be

cured in the spot of lower temperature.

At this point, one solution would consist on an improvement

of the coil design to decrease the temperature difference. An-

other solution would be the use of other heating methods that

lead to lower temperature differences at the bondline e.g. oven

heating.

0

10

20

30

40

50

60

70

80

0 5 10 15 20 25 30 35 40

To

tal

curi

ng

tim

e [m

in]

Temperature difference [°C]

Time


190

10.4.1. Curing cycles to be applied at the demonstrator

With the temperature differences measured in Section 10.2, the

optimal curing cycle is recalculated for each representative

bondlines (see Figure 10.1). The results are shown in Figure

10.16 to Figure 10.19.

Figure 10.16: Optimal curing cycle at bondline type A.

0

20

40

60

80

100

120

140

160

0 5 10 15 20 25 30

Tem

per

atu

re [

°C]

Time [min]

Optimal curing cycle at the hot spot A

Maximum temperature bondline

Temperature of the pyrometer

Minimum temperature bondline


191

Figure 10.17: Optimal curing cycle at bondline type B.

Figure 10.18: Optimal curing cycle at bondline type C.

0

20

40

60

80

100

120

140

160

0 5 10 15 20 25 30 35 40

Tem

per

atu

re [

°C]

Time [min]

Optimal curing cycle at the hot spot B




0

20

40

60

80

100

120

140

160

0 5 10 15 20 25 30 35 40

Tem

per

atu

re [

°C]

Time [min]

Optimal curing cycle at the hot spot C





192

Figure 10.19: Optimal curing cycle at bondline type D.

10.5. Assembly and analysis of the second

demonstrator

The second demonstrator is bonded as explained at Chapter 9:

Design and assembly of the demonstrator, applying the curing

cycles defined in the previous section. The results of the posi-

tioning of the parts are summarized in Table 10.5 for distances

and in Table 10.6 for orientations of the parts. Finally, the

quality control of the bondlines is shown in Table 10.7.

0

20

40

60

80

100

120

140

160

0 5 10 15 20 25 30 35 40

Tem

per

atu

re [

°C]

Time [min]

Optimal curing cycle at the hot spot D




10.5. ASSEMBLY AND ANALYSIS OF THE SECOND DEMONSTRATOR

193

Table 10.5: Validation of the positioning of the parts where the maximum

deviation according to RUAG specifications are 1 mm.

Distance measured [mm] CAD

[mm]

Deviation

[mm]

Left edge to rib 1 76.86 ± 0.22 77.5 0.64

Rib 1 width 20.21 ± 0.04 20 0.21

Rib 1 to rib 2 143.05 ± 0.21 143 0.05

Rib 2 width 20.25 ± 0.06 20 0.25

Rib 2 to rib 3 142.59 ± 0.30 143 0.41

Rib 3 width 20.27 ± 0.04 20 0.27

Rib 3 to rib 4 143.09 ±0.43 143 0.09

Rib 4 width 19.98 ±0.11 20 0.02

Rib 4 to right edge 65.38 ±0.22 65.5 0.12

Doubler to L-stringer 136.31 ± 0.05 136.42 0.11

End profile to L- stringer 140.86 ± 0.06 140.17 0.69

Doubler to T-stringer 156.93 ± 0.12 156.67 0.26

End profile to T-stringer 116.00 ± 0.20 115.48 0.52

L stringer width 10.37 ± 0.08 10 0.37

T stringer width 12.01 ± 0.22 12 0.01

Doubler to Lower skin 10.91 ± 0.05 11 0.09


194

Table 10.6: Validation of the orientation of the parts where the maximum val-

ue according to RUAG specifications are 1°.

Part Angle [°] Part Angle [°]

Rib 1 - left edge 0.06 T-stringer – End profile 0.01

Rib 2 - left edge 0.11 L-Stringer – End profile 0.04

Rib 3 - left edge 0.01 Doubler – End profile 0.01

Rib 4 - left edge 0.05

10.5. ASSEMBLY AND ANALYSIS OF THE SECOND DEMONSTRATOR

195

Table 10.7: Quality control of the second demonstrator where the maximum

value according to RUAG specifications are 0.3 mm for the bondline thick-

ness, a degree of cure higher than 95 % and a void content lower than 2 %.

Bondline

Degree

of Cure

[%]

Void con-

tent [%]

Thickness

[mm]

L stringer + L stringer to form

the T-stringer

95.52 1.43 ± 0.47 0.14 ± 0.03

T-stringer to lower skin 95.22 1.69 ± 0.24 0.16 ± 0.03

L stringer to upper skin 95.62 1.52 ± 0.23 0.18 ± 0.03

Rib 1 to inner doubler 95.05 1.62 ± 0.03 0.21 ± 0.07




Rib 1 to upper skin 95.54 1.19 ± 0.73 0.18 ± 0.05

Rib 2 to upper skin 95.44 1.05 ± 0.23 0.19 ± 0.03

Rib 3 to upper skin 95.01 1.18 ± 0.45 0.18 ± 0.06

Rib 4 to upper skin 95.20 1.59 ± 0.35 0.24 ± 0.03

Rib 1 to lower skin 95.69 1.65 ± 0.31 0.17 ± 0.03

Rib 2 to lower skin 95.58 1.88 ± 0.04 0.20 ± 0.04

Rib 3 to lower skin 95.24 1.28 ± 0.14 0.14 ± 0.02

Rib 4 to lower skin 95.22 1.39 ± 0.39 0.2 ± 0.04

Lower C-spar to inner doubler 95.44 1.15 ± 0.35 0.18 ± 0.03

C-spar to lower skin 95.01 1.25 ± 0.12 0.21 ± 0.04

Upper C-spar to inner doubler 95.14 1.21 ± 0.45 0.15 ± 0.04

C-spar to upper skin 95.57 1.61 ± 0.25 0.15 ± 0.04

Outer doubler to upper C-spar 95.20 1.60 ± 0.25 0.19 ± 0.06

Outer doubler to lower C-spar 95.42 1.39 ± 0.26 0.15 ± 0.02


196

The results show that the parts are well positioned, showing a

lower deviation than in the first demonstrator due to the expe-

rience gained. The quality control shows that the curing pro-

cess is completed in all the bondlines (the degree of cure was

always higher than 95 %). Additionally, it shows that the as-

sembly process is correct (the bondline thickness is below 0.3

mm). Finally, is proved that the void formation measured in

all the bondlines is lower than 2 %, with an average of all the

bondlines of 1.46 %, showing a good agreement with the ob-

jective value of the optimization of 1.5 %.

10.6. Conclusions

In this chapter, it is shown that a proper tool design is of high

importance to avoid an inhomogeneous temperature distribu-

tion at the bondline. The results also suggest that the edge ef-

fect is caused by a combination of the decrease of efficiency of

the coil at the edge, but mainly by an increase of convective

cooling at the bondline edge.

When the optimization tool is modified considering these ef-

fects, the new cycles require a longer curing time to guarantee

the bondline quality. As a consequence, the efficiency of the

process decreases.

The modified curing optimization cycles are applied to a sec-

ond demonstrator, bonded following the same process de-

scribed in Chapter 9: Design and assembly of the demonstrator.

The final demonstrator shows a good quality in all the

bondlines, validating the approach followed to ensure a ro-

bust curing process at the demonstrator.

197

Part V

Conclusions & Outlook

199

Chapter 11

Concluding remarks

11.1. Conclusions

The application paste adhesive bonding in industry is today

matter of research. Despite having a big number of ad-

vantages, the application of paste adhesive in critical parts still

is not commonly established. The main reason is that the

bonding process of paste adhesive has been shown to be very

sensitive to many processing parameters. The surface treat-

ment, geometry of the bondline or the difference between

measured and real temperature applied can lead to a quick

drop of the mechanical performance with catastrophic conse-

quences. The robustness of bonding processes is still one of the

reasons why this technique is not more commonly extended.

In this challenging context, the research in this doctoral disser-

tation was undertaken with the aim of investigating a robust

methodology to accelerate the curing process of paste adhe-

sives. With this objective in mind, a novel approach to opti-

mize this process is developed, discussed and validated. Addi-

tionally, the process robustness has been thoroughly investi-

gated with different adhesive systems and with components

with different shapes and sizes. As a conclusion of this re-

search work, a number of guidelines to ensure a robust bond-

ing process and to accelerate the curing process of paste adhe-

sives are summarized in this chapter.

CHAPTER 11: CONCLUDING REMARKS

200

11.1.1. Guidelines to ensure a good mechanical perfor-

mance in a bonded joint.

There are mainly 3 factors influencing the bondline quality

that have to be tightly controlled to ensure the robustness of

the process: The surface quality, the bondline geometry, and

the curing of the paste adhesive.

11.1.1.1. Surface treatment

Before the bonding process is carried out, the adherent can be

affected by different sources of contamination. This may re-

duce the strength of the joint, leading to a premature failure.

For this reason, a surface treatment is typically carried out,

ensuring a correct adhesion between the joining partner and

the paste adhesive.

1. Ensure a good surface treatment

The research carried out in this dissertation, summarized in

Section A.5, has proved that a mechanical sanding treatment

followed by a thorough surface cleaning process ensures a

good adhesion between the bonding partners and the paste

adhesive. The best surface treatment for single lap shear test

samples investigated in this dissertation is:

Manual sanding with P100 grain size

Manual sanding with P150 grain size

Acetone cleaning

Water cleaning

De-ionized water cleaning.

Drying of the partners in the oven at 65 °C for 1 hour.

It is proved that sanding just with P100 produce a sample with

more peaks of resin, which can be smoothed with the second

sanding process. Finally, the cleaning ensures the robustness

of the process, having a low variability of the results.

11.1. CONCLUSIONS

201

11.1.1.2. Bondline geometry

The geometry of the bondline affects the load carried by the

bonding system. A thick bondline induces a higher peeling

stress in a single lap shear joint loaded in tension. This leads to

a premature failure of the joint. For this reason, the bondline

thickness must be controlled.

2. Ensure a correct bondline thickness.

In Section A.5, it is proved that the bondline thickness affects

the mechanical performance of the joint in SLS samples. The

results show that SLS samples bonded with LME 10049-4 /

LMB 6687-2 from Huntsman Advanced Materials and loaded

in tension reduce its mechanical performance when the

bondline is thicker than 0.3 mm.

Additionally, it is proved that a very thin bondline thickness

does not reduce the performance of the joint. However, as it

must be guaranteed that the adhesive is placed everywhere to

avoid dry spots, a minimum bondline thickness of 0.1 mm

must be used. For this reason, in this research glass beads with

a nominal diameter of 0.1 mm are used to ensure a minimum

distance between the partners and pressure is applied during

the curing process to ensure that this distance is always lower

than 0.3 mm.

11.1.1.3. Curing of the paste adhesive

The curing process of the paste adhesive is a crucial step to

ensure the mechanical performance of the joint. For this rea-

son, this process must be controlled.

3. Ensure a correct curing process

According to the adhesive supplier, the curing process of the

pastes adhesives investigated in this dissertation is completed

when a degree of cure of 95 % is reached. Additionally, in


202

Chapter 6: Isothermal heating process is proved that a correct

curing cycle can be ensured when the void formation is lower

than 2 %.

In this dissertation is proved that the void formation depends

on the curing profile applied. For this reason, in order to en-

sure that the curing process is correct, the applied temperature

must be accurately controlled. Depending of the heating tech-

nique used to cure the paste adhesive, the temperature distri-

bution must be taken into account, as shown in Chapter 2:

State of the art.

In this context, an accurate simulation model of the heating

process is of interest to ensure the robustness of the process by

defining the optimal curing temperature profile and tempera-

ture sensor location.

11.1.2. Guidelines for a fast curing of paste adhesives

In this dissertation, a methodology to accelerate the curing

process of paste adhesives is described. This novel methodol-

ogy requires a number of guidelines to ensure the process ro-

bustness.

The curing process of paste adhesives can be accelerated by

considering a cure dependent heating process. This approach

requires a heating method that can vary the curing tempera-

ture faster than traditional heating methods e.g. oven heating.

4. Apply a variotherm curing process with induction heating

In Chapter 8: Cure dependent heating process, it is proved that a

variothermal process leads to an acceleration of the process

without decreasing the mechanical performance of the joint.

As the process investigated in this dissertation is carried out

with a variable temperature profile, the temperature differ-

ence in the bondline must be considered

11.1. CONCLUSIONS

203

5. Ensure low temperature difference in the bondline.

For induction heating the temperature differences in the

bondline are generated due to different phenomena detailed

in Chapter 9: Design and assembly of the demonstrator. They have

to be accurately controlled and reduced as much as possible in

order to lead to a faster process.

Edge effects: It is temperature decrease at the end of the coil.

It is caused by a higher heat dissipation combined with a low-

er efficiency of the coil at the edge, as shown in Section 14.4.

This effect can be minimized by an optimization of the coil

design, e.g. considering a longer coil than the susceptor, leav-

ing a distance around 10 cm at each side.

Variation of the laminate thickness: A region locally thicker

could have a different electrical conductivity, leading to an

inhomogeneous temperature generation.

Convective air flow: The convection coefficient affects the en-

ergy dissipated in the susceptors. A higher convection can be

used to accelerate the process, because the transients are faster

and therefore easier to control. The use of forced convection

can lead to a reduction of the temperature difference generat-

ed in the bondline if its homogeneity can be ensured along the

bondline.

In the approach under study, the use of forced convection is

considered for simple shapes, where it can be ensured that the

air flow is homogeneous in all the plate. In this context, the

use of a convective laminar air flow to cool the joint has re-

markable advantages in terms of process controllability, de-

spite requiring more power. As example, Figure 11.1 shows

the simulation of the heating process with natural and forced

convection from room temperature to 80 °C. In this case 40.8 A

and 203’000 Hz are applied for natural convection and 57.6 A

and 198’000 Hz are applied for forced convection. It is shown


204

how the use of forced convection reduces the transient periods

from 600 s to 150 s, leading to a more controllable process.

Figure 11.1: Forced and natural convection.

By the local adaptation of the processing parameters the tem-

perature gradient could be reduced. This would lead to an

optimal process that would also ensure the bondline quality of

the joint. However, this approach is not considered in this re-

search for complex shapes. The reason is the complexity to get

a homogeneous convection coefficient along the bondline due

to the shape of the components.

Distance between coil and the susceptor: A longer distance

decreases the temperature generated. This phenomenon can

be minimized by a proper coil design and by using well de-

fined spacers between the coil and susceptor in order to keep

the distance constant along the bondline.

All those phenomena can be taken into account by a simula-

tion tool to optimize the process and assess the temperature

distribution at the bondline. As proved in Chapter 4: Simula-

tion of the curing process of paste adhesives by induction heating,

the input process parameters affect to the temperature gener-

20

30

40

50

60

70

80

0 100 200 300 400 500 600

Tem

per

atu

re [ C

]

Time [s]

Natural convection heating

Natural convection cooling

Forced convection heating

Forced convection cooling

11.2. OUTLOOK

205

ated at the bondline and therefore the curing process. In this

case, as induction heating is used, the parameters affecting the

process are:

a. Material parameters of the susceptor:

The electrical conductivity affect to the heat gener-

ated by induction.

The conduction coefficient affects the heat transfer

between the susceptor and paste adhesive.

b. Process parameters:

The convection coefficient affects the heat dissipat-

ed by the joining parts.

The ambient temperature affects the heat dissipat-

ed by radiation and convection.

Equipment input parameter: The input electrical

current, efficiency of the coil and frequency of the

magnetic field affect the energy generated in the

susceptors.

c. Geometric parameters:

The distance between the coil and the susceptor af-

fect the energy generated in the susceptor.

The thickness of the laminate will affect the electri-

cal conductivity of the material.

The correct definition of those parameters is necessary to en-

sure a robust curing process optimized with respect to the

time.

11.2. Outlook

The use of a local heating system to cure paste adhesive joints

is a remarkable solution with a huge number of advantages

over riveting. For this reason, this must field should be further

investigated. At this point, there are still multiple open possi-

bilities to continue this research regarding modeling, pro-

cessing and applicability of the process.


206

With respect to the modeling, in this research an accurate two

dimensional tool is developed and validated. It is used to

measure the impact of the processing parameters in the quali-

ty of the process. In this area of investigation, an eventual ex-

tension of the simulation tool to 3D scenarios is of a great in-

terest. This would allow the process simulation of complex

structures and thus simulating the temperature differences

generated in the bondline. For instance, effects as the variable

distance between the coil and the susceptor, a variable thick-

ness of the susceptor could be accurately predicted. This more

sophisticated tool could also be used to optimize the coil in the

design phase, leading to a more homogenous temperature

generated and a thus a reduction of the curing processing

time.

With respect to the processing, there are several topics that

should be further investigated to consider the industrialization

of this methodology. One of these topics is the automation to

improve the accuracy of the process.

The development of a repeatable and robust process that does

not depend on the operator’s experience is the key to the in-

dustrial application of the methodology developed in this doc-

toral thesis. As a consequence, the design of an automated

control of the induction equipment would be of great value.

This would permit the fully automated application of vari-

otherm curing cycles, ensuring the accuracy of the curing

temperatures, leading to a more robust and controllable pro-

cess.

Regarding the applicability of the process investigated, there

is a huge potential for further research. In this dissertation, the

methodology to accelerate the curing process of paste adhe-

sives has been carried out only with structural aerospace ad-

hesives to bond CFRP joints. In this sense, other materials

could be considered for both, adherents and adhesives. This

11.2. OUTLOOK

207

work would extend this research to other fields where it could

mean a clear step forward compared to today’s state of the art

e.g. automotive and construction industries.

Regarding the applicability to different adherents, this re-

search has been focused in the induction bonding of CFRP

materials. These materials have remarkable electrical proper-

ties, which make them ideal candidates for induction heating.

However, there are many materials that do not have such elec-

trical conductive properties and therefore induction heating

cannot be applied today. In this thesis, this topic has been in-

vestigated considering a modification of the paste adhesive

with electrical conductive particles, ferromagnetic particles

and metallic meshes. This investigation, summarized in Ap-

pendix B.5, has shown very promising results, identifying

some of the limitations of this approach. In this sense, a more

comprehensive investigation of this topic would be of great

importance to extend the research carried out in this thesis to

other fields.

Finally, regarding the applicability of this process to other

paste adhesive systems, there is an open challenge. In this doc-

toral thesis, this methodology has been successfully applied to

two different structural adhesive systems to ensure its repeat-

ability. However, the extension of this research to other com-

pletely different paste adhesive systems would be of interest

not only to extern its use to other fields but also to identify

other process limitations.

One potential research field could be to investigate the lami-

nates curing process of by induction heating. In this case, the

benefits of a fast and robust curing process could be used to

accelerate the process. One example could be in the curing

process out of autoclave prepreg systems. This would reduce

the energy consumption and the environmental impact of the


208

process, proposing a solution to one of the major concerns of

composite technologies.

Another example of a field in which this investigation could

mean a high added value is automotive industry. Today,

many automotive components are bonded. The paste adhesive

systems used today cure faster than the ones used in aero-

space, taking place in minutes instead of hours. However, the

acceleration of the curing processes is not considered. The ap-

plication of the methodology investigated in this research

could reduce the processing time leading to a higher produc-

tivity, as was achieved in this doctoral investigation.

209

APPENDIX

APPENDIX A: MECHANICAL TESTING OF COMPOSITE SAMPLES

211

Appendix A

Mechanical testing of composite

samples

Mechanical testing is today state of the art to validate CFRP

bonded joints. In aerospace industry, there are several me-

chanical standards describing the methods used for validation

of composite bonded joints. One of the most important tests in

aerospace is Single Lap Shear test (SLS), where the shear

strength of the joint is measured in small coupons previously

bonded. This mechanical test is considered in this dissertation

as reference for the investigation of the acceleration of the cur-

ing process of a paste adhesive. The testing conditions used

are described in EN 2243-1 regarding preparation, cutting and

testing of samples.

However, for the correct application of this standard to com-

posite materials, there are some parameters that must be pre-

viously investigated. These parameters can influence on the

mechanical performance and lead to confusion on the inter-

pretation of the results. For this reason, the change of these

parameters is analyzed, defining the optimal conditions to

maximize the mechanical performance of the joint:

Thickness of the adherent

Thickness of the bondline

Surface treatment of the adherent


212

As a result, all the experiments carried out in this dissertation

are carried out under the same conditions, detailed in this Ap-

pendix.

A.1: Adherents and adhesives

In this dissertation, four materials are mainly used to investi-

gate the acceleration of the curing process of paste adhesives:

Paste adhesives

o LMB 6687-1 / LME 10049-4

o LME 10625 / LME 10626

Composite adherents

o CFRP

o GFRP

A.1.1: Paste adhesives

All the research in this doctoral thesis is carried out with two

different paste adhesive systems supplied by Huntsman Ad-

vanced Materials (Basel, Switzerland), partner of the Clean

Sky project. The main research is firstly carried out with the

adhesive system named LMB 6687 / LME 10049. Then, in or-

der to prove the repeatability of the research findings, the

paste adhesive system LME 10625 / LME 10626 is used.

Both systems are toughened bi-component paste adhesives

composed of an epoxy resin and a hardener. The mixing ratio

of this paste adhesive is 100:55, referring to the weight content

of epoxy:hardener, for the LMB 6687 / LME 10049 and 100:43

for the LME 10625 / LME 10626 [92].

During the period of this dissertation, the supplier changed

slightly the composition of the LMB 6687 / LME 10049, chang-

ing from LMB 6687-1 LME 10049-3 to LMB 6687-2 LME 10049-

4. However, the mechanical properties of the paste adhesive as

well as the curing conditions were not modified. For this rea-

A.1: ADHERENTS AND ADHESIVES

213

son, in Phase 1: Heating Concepts and in Chapter 7: Isothermal

heating process, the adhesive used is the LMB 6687-1 LME

10049-3, and for the rest of the research the adhesive used is

the LMB 6687-2 LME 10049-4.

A.1.2: Composite adherents

A.1.2.1: Carbon Fiber Reinforced Polymer (CFRP)

The CFRP adherents used in this dissertation are made of wo-

ven fabric prepreg supplied by CYTEC® (resin system MTM

44-1®, Cytec, UK, fibers Sigmatex CF 5804A, Benicia, CA,

USA), partner of the Clean Sky project. Each layer has a thick-

ness of c.a. 0.33 mm. The plates typically produced, if not indi-

cated the contrary, are made of three layers of fabric with ori-

entation [0, 90].

This system, despite being an out of autoclave system (OoA),

showed a low mechanical performance due to a high porosity

of the laminates. As a result, the tool surface was rather rough,

as can be observed in Figure A.1. For this reason, in this dis-

sertation samples are always cured in autoclave, following

supplier recommendations applying 180 °C for two hours

with 3 bars of pressure.


214

Figure A.1: OoA produces a bad surface quality.

A.1.2.2: Glass Fiber Reinforced Polymer (GFRP)

Some mechanical testing of GFRP bonded samples is carried

out in this dissertation. The GFRP adherents, already cured,

are supplied by Swiss Composites AG.

A.2: Production of bonded samples

A.2.1: Production of CFRP laminates

If not indicated the contrary, the plates produced for the me-

chanical testing in this dissertation are made using a typical

vacuum bag lay-up on a tool made of steel that is covered by a

release film (WL 3900R, Airtech Advanced Materials, Hun-

tington Beach, CA). Then, on top of the covered mold, the wo-

ven CFRP lay-up is made always with an orientation [0, 90],

placing glass fiber cords as recommended by the supplier.

Then, the laminate is covered by another release film (200 P3

150 500, Flugzeug Union Süd GmbH, Ottobrunn, Germany).

Finally, the entire layup is covered with a cotton breather (N4

Airwave, Airtech Advanced Materials, Huntington Beach, CA)

Sample cured in autoclave: Smooth

surface

Sample cured out of autoclave:

Deficient resin flow

A.2: PRODUCTION OF BONDED SAMPLES

215

and then sealed with a vacuum bag (WL 7400 001 48’ 1000 LFT

48, Airtech Advanced Materials, Huntington Beach, CA). Dur-

ing this dissertation, a special layup was manufactured with a

peel ply (CS, Interglass Technologies AGGmbH, Erbach, Ger-

many), in order to evaluate its influence on the mechanical

performance of the bonded joint. The lay-up used in this dis-

sertation to produce the samples for mechanical testing is

schemed in Figure A.2.

Figure A.2: CFRP layup scheme for the CFRP processing.

The curing process of the panels is carried out in autoclave

under supplier’s recommendation. The conditions are:

Vacuum: -0.9 bar

Autoclave pressure: 3 bar

Temperature: 180 °C

Time: 2 hours

Heating rate: 2 °C/min

A.2.2: Bonding process

Previously to the bonding process, the adhesive is mixed. The

dosing of the paste adhesive is carried out under supplier

specifications. Then, the mixing is carried out with a centri-

fuge mixer (SpeedMixer DAC 150.1 FV, SPX, Charlotte, North

Carolina) for three periods of 1 minute. The two first ones

stages are carried out at 1500 rpm and the last one at

2500 rpm. This mixing process is carried out for all the sam-

Release film

Laminate

(Peel ply)

Release film

Steel mould

Breather

Vacuum bag


216

ples used in this dissertation in order to make the process re-

peatable [93]. Sometimes, by the recommendation of the sup-

plier, glass beads of 0.1 mm are mixed to the paste adhesive in

order to ensure a minimum bondline thickness. They are

mixed in 0.5 wt%, a small quantity that does not influence the

mechanical performance of the adhesive.

After mixing the adhesive is placed on the bondline area. Fi-

nally the parts to be bonded are manually pressed together

and finally cured.

The bondline thickness is controlled by spacers placed under

one the upper bonding partner, creating the necessary gap of

c.a. 1.1 mm (1 mm for the CFRP plate and 0.1 mm for the de-

sired bondline thickness). Then, some weight is placed on the

upper bonding partner during the process to ensure a minimal

bondline thickness.

A.3: Single lap shear testing

The standard considered in this dissertation to validate the

bonded samples is the single lap shear test described in the

Intentional Standard EN 2243-1. This method is typically used

for validation of adhesive bonded samples by Airbus, partner

in the Clean Sky project.

This standard recommends bonding big plates, which after-

wards are cut into small coupons for testing. In this disserta-

tion the adherent plates are always cut, if not indicated the

contrary, with a size of 100 mm x 300 mm x 1 mm. Afterwards,

the surface treatment is applied and then the plates bonded

with an overlap of 12.5 mm and a bondline thickness of less

than 0.3 mm.

The curing process of the paste adhesives in this dissertation is

carried out by different techniques and curing cycles that al-

ways are detailed. The samples are cut after bonding with a

A.4: FRACTURE MODES

217

diamond saw (Diadisc 4200, Mutronic GmbH & Co Kg,

Rieden, Germany) with 25 mm of width. Samples are then

placed in the testing machine (Zwick Roell 1474, Ulm, Germa-

ny) and tested under the following conditions:

Distance between grips: 100 mm.

Testing speed: 2 mm/min.

Overlap between plates: 12.5 mm

The samples are tested until fracture. Then, the maximum

force is used to calculate the shear strength of the samples,

dividing the maximum value of the force obtained by the

bondline area.

A.4: Fracture modes

Apart from measuring the maximum shear strength of the

samples, the failure mode is characterized. In adhesive bond-

ing, three main types of failure can be observed depending on

the location of the failure. If the fracture is located in one side

of the adhesive, having most of the bondline in one of the

sides, then is called adhesive failure. An example of this fac-

ture mode is shown in Figure A.3.


218

Figure A.3: Adhesive failure in a wedge sample.

This failure mode typically indicates adhesion problems be-

tween adhesive and adherents. Most of the times, this failure

mode is produced by a deficient surface treatment of the ad-

herents [94]. This fracture mode can also happen if the adhe-

sive is completely degraded by the application of high tem-

peratures in the curing process.

If the failure is located at the center of the bondline, as shown

in Figure A.4, it is called cohesive failure.

A.4: FRACTURE MODES

219

Figure A.4: Cohesive failure in a wedge sample.

This kind of failure indicates that the fracture is produced be-

cause the adhesive could not carry the loads. Typically, this

fracture mode indicates certain degradation of the paste adhe-

sive, mostly caused by the application of too high tempera-

tures in the bonding process.

If the fracture is placed in the adherent, in this case by delami-

nation of the composite, as shown in Figure A.5, the fracture is

called adherent.


220

Figure A.5: Adherent failure in a wedge sample.

This fracture mode means a good preparation of the samples

and a good quality of the bondline. Bonding systems showing

this failure mode can take more loads than the adherents.

Finally, it can occur that there is a mixed fracture mode, as

shown in Figure A.6.

Figure A.6: Mixed failure in a wedge sample.

A.5: PREPARATION OF SAMPLES FOR MECHANICAL TESTING

221

In this research, this fracture mode can occur when the curing

process is carried out at a high temperature which is not high

enough to produce a clear thermal degradation.

A.5: Preparation of samples for mechanical

testing

A.5.1: Adherent and bondline thickness effect on the me-

chanical performance

In order to study the influence of these parameters, four plates

are bonded following the process detailed in Section A.2: Pro-

duction of bonded samples. The following geometrical parame-

ters are considered:

Adherent thickness 1 mm, bondline thickness 0.25 mm




Additionally, in order to measure the shear strength of the

CFRP independently from the adhesive, two more CFRP joints

are manufactured by co-curing. Instead of paste adhesive, ex-

tra layers of CFRP are placed in the overlap, joining the plates

during the curing process of the laminates. The thickness of

the “bondline” in this case is around 0.5 mm when 1 layer is

placed and 0.7 mm with 2 layers in the bondline area.

The results of the samples tested by SLS test are shown in Fig-

ure A.7.


222

Figure A.7: Effect of bondline and adherent thicknesses.

It can be observed that the bondline thickness affects the me-

chanical performance of the joint. The samples with a bondline

thickness lower than 0.4 mm show an adherent failure. Sam-

ples with higher values show adhesive failure due to the high-

er peeling stress induced by the bondline thickness. This leads

to a lower mechanical performance. For this reason, the sam-

ples bonded in this dissertation are validated only if the

bondline thickness is lower than 0.3 mm.

The thickness of the adherent shows a smaller influence on the

mechanical performance of the joint. It can be observed that a

thicker laminate shows less dependence on the bondline

thickness. The reason is that a thicker laminate bends less dur-

ing the testing, inducing a lower peeling. Nevertheless, in or-

der to have a validation system more sensible to the bonding

quality, laminates with 1 mm thickness are used for the me-

chanical testing considered in this dissertation.

Additionally, co-cured samples delaminate at about 17 MPa.

These results are obtained with a thick bondline of more than

10

12

14

16

18

20

22

24

26

0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8

Sh

ear

stre

ng

th[M

Pa]

Bondline thickness [mm]

Adherents thickness 1 mm

Adherents thickness 2 mm

Co-curing


223

0.4 mm, affecting the results. A thinner SLS sample would

probably delaminate at a higher load.

A.5.2: Selection of the adherent surface treatment

In the production of adherents, the new parts are affected by

different sources of contamination that can affect the strength

of the joint [92]. Another problem in the manufacturing pro-

cess of the CFRP laminates is that the composite tool side has a

higher resin percentage, while carbon fibers remain some mi-

crometers under this first layer. This fact affects the shear and

peel strength of the joint. To avoid these undesirable effects,

some treatments can be applied in order to minimize the

amount of resin [15].

The surface treatments are typically divided in three main

blocks: mechanical abrasion, chemical pretreatment and spe-

cial methods for CFRP such as the use of a peel ply [95]. Addi-

tionally, the use of release agents in the manufacturing process

of the CFRP plates is also considered in this study.

The highlights of today’s state of the art set abrading methods

as common approach for surface treatment. However, several

companies prefer the use of peel ply over abrading methods

due to its low cost and reduced human factors [96]. In either

any case, a cleaning, chemical or just manual scoring must be

applied. This research compares different approaches for sur-

face treatment of CFRP bonded plates, proving the importance

of surface treatment before bonding.

In this section, different surface treatments are applied to the

CFRP adherents. Additionally, the surface treatments are

combined with different manufacturing process, release film

(noted as RF) and peel ply (noted as Pp) are considered for the

production of the laminates. Then, the mechanical perfor-

mance of the different samples is compared.


224

The abrading methods consist on removing the layer of resin

placed at the top of the composite plate. Different methods

following this principle have been applied:

Manual sanding P100 (noted as S100) and P150 (noted as

S150): The election of the size of the sand plays an important

role. With a bigger size, more resin is removed being the pro-

cess faster and simpler, but as drawback, the quality of the

bond can decrease due to fracture of fibers. Typical sizes rec-

ommended for composites in literature are from P80 to P180

for rough sanding [17].

Air-grit blasting (noted as GB): This technique, commonly

used in industry as surface treatment, consists on particles

thrown with a pressurized air gun [97]. These small particles

hit the surface to be treated removing the resin. The equip-

ment used is a Mikromat 760s, (Mikromat GmbH, Dresden,

Germany), using quartz sand and a jet pressure of 2 bar.

Water jet: The equipment used is a Bystronic water jet cutting

machine (Waterjet AG, Aarwachen, Switzerland). In order to

get the desired impact of the jet, removing the resin and not

breaking the fibers, the water jet is used without abrasive ma-

terial (noted as WJ). Samples are treated with a flat jet nozzle

(Procer SAS, Möhlin, Switzerland) with a spraying angle of 30

degrees and a flow rate equivalent of a 0.3 mm diameter noz-

zle. The operation conditions are a traverse speed of

300 mm/min and a distance around 19 mm.

Finally, after the treatment is carried out, a cleaning process

(noted as C) is carried out in some of the samples. In this pro-

cess, acetone is applied and afterwards is dried with a paper.

Then, water is used to remove the remaining acetone from the

surface to be bonded. Additionally, de-ionized water is ap-

plied to the surface and finally the samples are dried in a con-

vection oven for 60 minutes at 65 °C.


225

The CFRP laminates are bonded with is 100 °C for 1 hour,

achieving more than 95 % of degree of cure. The results of SLS

for different combinations of surface treatments are summa-

rized in Table A.1 and Figure A.8.

Table A.1: SLS for different surface treatments.

Sample

name

Manufacturing

process

Surface treat-

ment

Shear

strength

[MPa]

RF Release film - 16.4 ± 3.1

RF+C Release film Cleaning 22.1 ± 1.1

RF+S100+C

Release film Sanding P100 +

Cleaning 22.0 ± 1.8

RF+S100+S150+C

Release film Sanding P100 +

P150 + Cleaning 24.4 ± 1.5

RF+GB+C Release film Grit blasting +


RF+WJ Release film Water jet FN 12.3 ± 1.5

RF+WJ+C Release film Water jet FN +


Pp+C Peel ply Cleaning 20 ± 2.1

Pp+GB+C Peel ply Grit blasting +



226

Figure A.8: SLS test for different surface treatments.

The fracture modes are summarized in Table A.2.

Table A.2: Fracture mode of SLS samples.

It is proved that a simple cleaning process improves the me-

chanical performance of the joint by about 50 % compared to

0

5

10

15

20

25

30S

hea

r st

ren

gth

[M

pa]

Sample name Fracture mode

RF 50 % Adhesive 50 % Cohesive

RF+C 20 % Adhesive 80 % Adher-

ent RF+S100+C Adherent

RF+S100+S150+C Adherent

RF+GB+C Adherent

RF+WJ 70 % Adherent 30 % Cohesive

RF+WJ+C 90 % Adherent 10 % Cohesive

Pp+C 30 % Adhesive 70 % Cohesive

Pp+GB+C 70 % Adherent 30 % Cohesive


227

non-treated samples. For this reason, cleaning is shown to be

the simplest method to improve the mechanical performance.

The best surface treatment method for single lap shear test

samples is sanding with P100 followed by a second sanding

with P150 and then cleaned. The reason is that sanding just

with P100 produce a sample with more peaks of resin, which

can be smoothed with the second sanding, reducing as well

the variability of the results.

Grit blasting and water jet show a low mechanical perfor-

mance with adherent failure. Both methods are difficult to

control due to the dependence on the operator. This can result

in the destruction of a high percentage of fibers, meaning a

decrease of the mechanical properties.

Samples produced with a peel ply have a lower performance

than samples produced with release film. The reason is the

load concentration in the resin peaks formed during the pro-

duction of the plate. Additionally, the application of abrading

methods in samples produced with a peel ply will mean a de-

struction of fibers, decreasing its mechanical performance.

For these reasons, the surface treatment method to be applied

on the preparation of the samples in this dissertation is a sand-

ing with P100 followed by a second sanding with P150 and

finally cleaned with acetone, water and de-ionized water.


228

APPENDIX B: INDUCTION SETUP

229

Appendix B

Induction setup

In this Appendix, the equipment for induction heating setup is

shown. The materials used in the assembly rig are detailed in

Section B.1, Geometry and Materials. Then, the effect of the dis-

tance between the coil and susceptor is measured in Section

B.2, Effect of the distance between the coil and the susceptor. After-

wards, Section B.3: Adaptation of the simulation tool for natural

convection conditions, describes how to model a natural convec-

tive flow in a 2 dimensional model. Afterwards, the produc-

tion process of samples for the void content analysis by induc-

tion heating is described in Section B.4, Production of samples by

induction heating. Finally, the modification of paste adhesives is

investigated in Section B.5, Modification of the paste adhesive for

bonding nonelectrical conductive materials by induction heating.

B.1: Geometry and materials

The induction setup consists of an AC power supply that pro-

vides high current and high frequency (maximum of 385 A

and 400 kHz) by an EasyHeat unit of 2.4 kW (Ambrell, Scotts-

ville, NY, USA). The frequency applied by the equipment is

set automatically depending on the current applied. Neverthe-

less, the frequency applied is detailed for all the experiments

carried out in this dissertation and considered as an important

parameter. The power supply is combined with a “Fluxtron”

coil supplied by Plustherm Point Ltd. (Wettingen, Switzer-

land), placed at the end of the inductor head. The equipment

is connected to a cooling unit Hyfra Chilli-15 (HYFRA Indus-


230

triekühlanlagen GmbH, Krunkel, Germany) as shown in Fig-

ure B.1.

Figure B.1: Induction setup used.

The test setup is designed using non-electrical conductive ma-

terials to fix the plates during the bonding process of SLS

samples. As a consequence, the coil is designed having a sin-

gle-turn copper beam and a magnetic field concentrator at-

tached to the lower part of the coil made of a material called

“Ferrotron 559H” (Fluxtrol, Auburn Hills, MI, USA). The test

setup supports susceptors on a sandwich structure made of

GFRP with a core used as base. The core is made of Aramid

honeycomb lightweight honeycomb made from Nomex pa-

per® (Kevlar® paper) Additionally, wood is used to fix the

plates to be bonded and to support the sandwich structure, as

shown in Figure B.2.

Inductor head

Coil

Coolingunit

Powergenerator

B.1: GEOMETRY AND MATERIALS

231

Figure B.2: Test rig designed to bond SLS samples.

The materials of the different components of the test rig are

detailed in Figure B.3. Additionally, the material properties

used to model the simulation tool described in Chapter 5:

Simulation of the curing process of paste adhesives by induction

heating are listed in Table B.1.

Inductor head

Coil


232

Figure B.3: Geometry of the test setup.

Co

pp

erFe

rro

tro

nC

FRP

Ad

hes

ive

Air

GFR

PSa

nd

wic

hco

reW

oo

dx

zy

B.1: GEOMETRY AND MATERIALS

233

Table B.1: Material properties of the test rig.

Ele

ctro

mag

net

ic p

rop

erti

es

Hea

t tr

ansf

er p

rop

erti

es

Ele

ctri

cal

con

du

ctiv

ity

𝝈 [

S/m

]

Rel

ativ

e

per

me

abil

ity

𝝁 [

-]

Rel

ativ

e

per

mit

tiv

ity

𝜺 [-

]

Th

erm

al

con

du

ctiv

ity

λ [

W/m

·K]

Hea

t ca

pac

ity

[

J/K

g·K

]

Den

sity

ρ

[Kg

/m3 ]

Em

issi

vit

y

ε [-

]

Co

pp

er

58.1

·10

6 1

18.1

38

4 34

0 8’

960

0.96

Fer

rotr

on

6.

7·10

-7

18

14.2

4

450

5’90

0 0.

8

CF

RP

5’

912.

1 1

3.4

3(

T-2

73)+

960

1’54

0 0.

9 2

0.5

Ad

hes

ive

2.5·

10-1

2 1

3.6

0.33

14

00

1’10

0 0.

9

GF

RP

1·

10-1

0 1

3.8

0.35

94

0 1’

700

0.8

San

dw

ich

core

1·

10-1

0 1

2.5

0.2

800

100

0.91

Wo

od

1·

10-9

1 2.

6 0.

13

1381

84

0 0.

91


234

Most properties considered in this Appendix are common

values well documented in literature. Others are obtained

from the supplier [10, 98]. Finally, some other properties re-

quire be either explaining or measuring experimentally e.g.

density.

The electrical conductivity of the CFRP is measured for the

whole laminate by applying an electrical current of 1 A from a

power source [KEPCO 20-5M] to samples of CFRP with a size

of 40 mm * 10 mm * 5 mm. The voltage is measured with a

multimeter [Mastech MY-67] having a separation between the

contacts of 20 mm. Then, 20 values are measured having an

average of 5912.1 ± 94 S/m.

Regarding the thermal properties of CFRP, a theoretical calcu-

lation of the thermal properties is carried out. It is considered

that the microstructure of CFRP determines its material prop-

erties [79]. In this context, the CFRP is modeled as homoge-

nized anisotropic material. In literature the thermal conductiv-

ities of carbon fibers can vary from 2.5 to 10 W/m·K, e.g. 5

W/m·K [82], the one of epoxy resins is varying from 0.2 to 0.35

W/m·K [99]. As the real values are not measured, assessed

values of 2.0 W/m·K for x-direction (see Figure B.3) and 0.5

W/m·K for y-direction are used. These values lie within the

range which is achieved by applying the rule of mixture tak-

ing into account fiber orientation.

The heat capacity of the CFRP plates is modeled as tempera-

ture dependent using a model that considers the fiber volume

content and the individual properties of the resin and fibers

[82]. Using the mixing rule with the individual temperature

dependent heat capacities of resin and fibers, the final heat

capacity of the composite is calculated.

B.2: CALCULATION OF THE CONVECTION COEFFICIENT

235

B.2: Calculation of the convection coefficient

In the modeling of the heat transfer module, forced convection

is considered in order to to make the convection coefficient

more reliable and controllable compared to natural convec-

tion. The flow of the air is considered in z-direction, perpen-

dicular to the model plane, making the convection coefficient

independent from the geometry and homogenous in the

section of the model. The forced convection coefficient can be

calculated by Eq. B.1.

. B.1

[W/m·K] is the thermal conductivity of the air, [m] is the

distance between the beginning of the plate and the measuring

point of the coefficient and [-] is the Nusselt number,

shown in Eq. B.2, which measures the increase of heat transfer

from a surface surrounded by a fluid compared to the heat

transfer by pure conduction [100].

B.2

Nusselt number considered for this model is for laminar fluid

conditions [101], depending on the Reynolds and Prandtl

numbers, shown in Eqs. (B.3 and B.4):

B.3

B.4

[kg/m3] is the density of the air, [m/s] is the speed of the

fluid, [Pa·s] is the dynamic viscosity of the fluid and

[J/Kg·K] is the heat capacity of the fluid.


236

The thermal parameters of the air are calculated with the

equations shown in Table B.2 [27].

Table B.2: Summary of dry air properties.

Thermal conductivity

λ [W/m·K] (3.807 + 0.074·T)·10-3

Density ρ [Kg/m3]

Dynamic viscosity

μ [Pa·s] (

)

Heat capacity

Cp [J/Kg·K]

As it can be observed, the convection coefficient depends on

the properties of dry air that are function of the air pressure

[Pa] and the temperature [K]. To determine this parame-

ter, the air pressure is considered constant and equal to

101’300 Pa, the fluid speed 2 m/s at the beginning of the plate

and the distance between the beginning of the plate and the

simulation section is 0.1 m as shown in Figure B.4.

Figure B.4: Conditions for the calculation of hc.

Then, the air properties as well as Reynolds and Prandtl num-

bers are calculated as function of the temperature to prove

laminar conditions. Finally, the convection coefficient is calcu-

lated considering certain room temperature. If the value of the

Section A-A’

A-A’

B.3: ADAPTATION OF THE SIMULATION TOOL FOR NATURAL

CONVECTION CONDITIONS

237

ambient temperature is considered 24.5 °C, the convection co-

efficient is equal to 17.43 W/m2·K with a Prandtl number equal

to 0.716 and a Reynolds number of 12'864, what are values

typically considered in the range of laminar forced convection.

B.3: Adaptation of the simulation tool for nat-

ural convection conditions

Due to the shape complexity of the demonstrators described in

this doctoral thesis, they are cooled by natural convection. In

this section, a simulation tool is described, determining the

convection coefficient when there is no forced air flow. Addi-

tionally, as the coil is changed (heating from one side only) to

heat the long bondlines, the new value of the efficiency must

be determined previously.

In this context, a simple experiment is carried out consisting of

heating one of the C-spars used in the demonstrator and com-

paring to the temperature generated. The part is initially heat-

ed applying forced convection, with the air conditions detailed

in Appendix B.1. The parameters used for the simulation are

57.6 A at 198’000 Hz, with an ambient temperature of 21.1 °C.

The temperature is measured at point C of Figure B.5 in steady

state conditions for 10 times with an optical fiber pyrometer

[FTC-DIN-ST-HA-LS by Photon control, CA]. The average is

then compared to the predicted temperatures generated by

induction, shown in Figure B.6.


238

Figure B.5: Geometry to validate natural convection.

Figure B.6: Efficiency of the coil.

The results show that the efficiency of the coil is around 82 %.

In order to validate the simulation tool, additional test are car-

ried out. The temperature distribution at points A to E and the

heating and cooling processes for 150 s at point C are also

compared. The results are shown in Figure B.7 and Figure B.8.

5 mm5 mm5 mm5 mm

70

72

74

76

78

80

82

84

Tem

per

atu

re [ C

]



239

Figure B.7: Temperature distribution of the C-spar.

Figure B.8: Heating and cooling processes of the C-spar.

The simulated results show a good agreement with experi-

mental data, thus validating the modeling of the pancake-

shaped coil used in this part of the dissertation. The maximum

temperature difference is located at the edges, where there is a

temperature difference lower than 3 °C.

In a next step, the use of natural convection is considered. As

the simulation tool is validated for a forced convection scenar-

62

64

66

68

70

72

74

76

78

80

A B C D E

Tem

per

atu

re [ C

]

COMSOL

Experimental

20

30

40

50

60

70

80

90

0 20 40 60 80 100 120 140

Tem

per

atu

re [ C

]

Time [s]

Experimental heatingExperimental coolingCOMSOL heatingCOMSOL cooling


240

io, the temperature is measured with no forced convection.

Then, it is compared to the simulation results with different

convection coefficients around 5 W/m2*K, which are typical

values for natural convection conditions [27]. The parameters

used for the simulation are 40.8 A at 203’000 Hz, with an am-

bient temperature of 20 °C. The temperature is measured at

point C of Figure B.5 in steady state conditions for 10 times

with an optical fiber pyrometer [FTC-DIN-ST-HA-LS by Pho-

ton control]. The results are shown in Figure B.9

Figure B.9: Assessment of the convection coefficient.

The results show a good agreement between simulation and

experimental data when using a convection coefficient of

4 W/m2*K. This value is in the range of natural indoor convec-

tion coefficient values (between 3 and 10 W/m2*K) and there-

fore used to complete the simulation tool for complex shapes.

Finally, the tool is validated previously to be applied to the

simulation of the optimal curing process. For this reason, the

60

62

64

66

68

70

72

74

76

78

80

Experimental hc = 3

W/m2*K

hc = 4

W/m2*K

hc = 5

W/m2*K

hc = 6

W/m2*K

Tem

per

atu

re [ C

]



241

tool is validated with steady state conditions by applying dif-

ferent currents and for heating and cooling processes. The

temperatures are measured at the point C (see Figure B.5) and

then compared to the predicted values (see Figure B.10).

Figure B.10: Results for different currents applied.

The measurement of the temperature generated by different

currents proves that the tool can be applied to predict the op-

timal curing process with natural convection in all the temper-

ature range.

The validation of the heating and cooling for 600 s is shown in

Figure B.11.

0

10

20

30

40

50

60

70

80

90

20 25 30 35 40 45

Tem

per

atu

re [ C

]

Input current [A]

Experimental

COMSOL


242

Figure B.11: Heating and cooling with natural convection.

The results show a temperature difference lower than 3 °C in

both processes, proving the accuracy of the simulation tool

modeled.

B.4: Production of samples by induction heat-

ing

For the validation of the different models, several samples

have been produced by induction heating. In this section, the

details to produce them are given.

The samples produced typically consist of different isothermal

stages at different temperatures applied by induction. Most of

the samples produced consist on a dual step curing process, as

shown in Figure B.12, combining different process parameters

on the heating stages ( , ).

20

30

40

50

60

70

80

90

0 100 200 300 400 500 600

Tem

per

atu

re [ C

]

Time [s]

Experimental heating

Experimental cooling

COMSOL heating

COMSOLcooling

B.4: PRODUCTION OF SAMPLES BY INDUCTION HEATING

243

Figure B.12: Cycle to produce samples for validation.

The CFRP plates are dosed with 0.3 mm of paste adhesive as

follows:

The paste adhesive is placed on the plate with a wood

stick.

The adhesive is compressed over the desired surface with

a steel tool that has a cavity of 0.3 mm, by which the adhe-

sive passes along the bondline.

The exceeding adhesive is removed from the plate.

Then, the samples are placed under the coil and the input elec-

tric current is manually tuned. The temperature profile is then

generated in the CFRP thus curing the paste adhesive. A

scheme is shown in Figure B.13.

The temperature is measured with different systems:

IR pyrometer [IN 510, LumaSense].

Optical fiber pyrometer for low temperature meas-

urements [FTP-NY2-ST1-2M-BA, Photon Control,

Canada]

Optical fiber pyrometer for high temperature meas-

urements [FTP-NY2-ST2-2M-BA, Photon Control,

Canada]

0

20

40

60

80

100

120

140

160

180

0 500 1000 1500 2000

α > 0.95

T1

T2

t1 t2

Tem

per

atu

re[°

C]

Time [min]


244

The method used depends on the experiment conditions and

is specified for each case.

Figure B.13: Scheme to produce samples by induction.

The samples are then grinded and polished with a rotatory

polishing machine (Streuers, GE). The time applied for the

polishing steps varies depending on the sample, guaranteeing

that the bondline is observable in the final sample. 6 steps are

applied with the following grain sizes for the grinding and

diameter of the diamond suspension for the polishing.

Step 1: 120 grain polishing.



Step 4: 9 μm diameter for the diamond suspension for pol-

ishing for 5 minutes.





Then, the void content is measured by microscopy techniques.

To carry out an accurate measurement, at least 3 images are

taken from each sample with the microscope [Leica DM RXA,

Inductor head

Optical fiber

pyrometerCoil

CFRP

Paste adhesive

B.5: BONDING OF NON-ELECTRICAL CONDUCTIVE MATERIALS BY

INDUCTION

245

Leica Microsystems GmbH, Wetzlar, Germany]. The void con-

tent is then measured inside of a defined frame, as shown in

Figure B.14.

Figure B.14 Sample for the microscopy measurement.

B.5: Bonding of non-electrical conductive ma-

terials by induction

B.5.1: Introduction

As commented in Chapter 2, the performance of induction

heating depends on the electrical or magnetic properties of the

susceptors. In this appendix, the possibility to modify the

paste adhesive in order to generate heat internally is investi-

gated [36, 37]. The approach consists on the addition of electri-

cal conductive and ferromagnetic materials permitting to bond

non-electrical conductive adherents [14, 102-104].

Previous research has been carried out using a conductive ma-

terial as one of the adherents e.g. aluminum [105]. The use of a

conductive mesh made of copper applied in the adhesive

bondline has also been investigated. In both cases, big temper-

ature gradients appeared in the curing process of the paste

adhesive leading to a degradation of the paste adhesive and a

decrease of the mechanical performance on the joint [40].

1’000 μm


246

The mixture of magnetic particles has also been investigated in

paste adhesive systems curing at high temperature [34, 41]. It

has been proved that curing paste adhesive with 50 wt% of

particles is possible.

In order to investigate the bonding process of non-electrical

conductive materials, the following experimental approach is

considered:

1. Measure the temperature generated in the modified

paste adhesive with meshes made of different materi-

als and different particles.

2. Bond and test SLS structures made of non-electrical

conductive materials by induction heating.

B.5.2: Experimental

The modification of the paste adhesive is investigated for the

system LME 10049-4 / LMB 6687-2. This paste adhesive is nei-

ther electrically conductive nor ferromagnetic. As a conse-

quence, some materials affected by induction are mixed with

the paste adhesive in order to generate the heat directly in it.

The mixing process is detailed in Appendix A.2. In this chap-

ter, two different types of modification are considered:

Addition of particles with different weigh content, di-

ameter and materials.

Addition of meshes with different geometrical param-

eters e.g. different diameter of the wire.

The major properties of the materials used in this study are

summarized in Table B.3.


INDUCTION

247

Table B.3: Particles used for adhesive modification.

Ty

pe

M

ater

ial

Bra

nd

nam

e S

up

pli

er

Par

ticl

e

dia

met

er

[μm

]

Mag

net

ic

satu

rati

on

[A·m

2 /K

g]

Ele

ctri

cal

con

du

ctiv

ity

[S/m

]

Par

ticl

es

Iro

n

117.

0160

S

wis

s C

om

po

site

s A

G

60

218

1·10

7

Par

ticl

es

Ste

el

117.

8001

S

wis

s C

om

po

site

s A

G

80

1.8

1.5·

106

Par

ticl

es

Co

ated

FeO

M

agS

ilic

a E

vo

nik

0.

05

20

7·10

-12

Par

ticl

es

Alu

min

um

74

4.00

01

Sw

iss

Co

mp

osi

tes

AG

10

0 -

3.5·

107

Mes

h

Alu

min

um

A

lMg

5 H

aver

& B

oec

ker

-

- 3.

5·10

7

Mes

h

Nic

kel

-

Hav

er &

Bo

eck

er

- 55

1.

4·10

7

Mes

h

Ste

el

X6C

r17

H

aver

& B

oec

ker

-

1.8

6·10

5

Mes

h

Co

pp

er

- H

aver

& B

oec

ker

-

- 5.

9·10

7


248

The induction equipment used is detailed in Appendix B.1.

The temperature generated is measured in the modified paste

adhesive. The device used is an optical fiber temperature sen-

sor model FTC-DIN-ST-HA-LS (Photon Control, Burnaby,

Canada). This test is carried out by adjusting the current in

order to generate 80 °C at the different samples.

B.5.3: Results

B.5.3.1: Temperature generation

The heating of samples of 100 mm x 25 mm x 0.35 mm modi-

fied with particles and meshes is carried out in steady state. It

is compared to the heating of the pure adhesive, not affected

by induction. However, the sample of pure paste adhesive is

heated by convection from the warm coil due to losses during

the heating process, as shown in Figure B.15. Samples with

wt% 50 are marked red, wt% 25 are marked orange and wt%

15 are marked green.

Figure B.15: Temperature generation in modified adhesives.

20

30

40

50

60

70

80

90

0 100 200 300 400 500

Tem

per

atu

re [

°C]

Current [A]

Steel

Nickel

Copper

Aluminium

MagSilica

Steel AS80

Steel RZ60

Aluminium

Pure Paste

adhesive

Steel 80 μm

Iron 60 μm

Meshes

Pure adhesive

Particles

Aluminum

Aluminum

Meshes

Particles


INDUCTION

249

In general, samples with meshes generate more heat than the

samples with particles. For this reason, the samples with

meshes and particles are compared independently in Figure

B.16 and Figure B.17.

Samples with meshes are compared, analyzing the role of dif-

ferent geometrical parameters summarized in Table B.4. The

loop area is also included, being the fraction of empty area at

each loop formed by the wires.

Table B.4: Geometric parameters of the meshes used.

The current necessary for each sample to achieve 85 °C is plot-

ted in Figure B.16 considering the numbering of Table B.4.

Material

Distance

between

wires L

[μm]

Diameter

of the

wires D

[μm]

Skin

depth

δ [μm]

Loop

area [%]

1 Aluminum 75 53 242 34.3

2 Copper 140 112 129 30.9

3 Copper 25 140 129 41.1

4 Nickel 160 100 11 37.9

5 Nickel 224 80 11 54.3

6 Nickel 500 140 11 61.0

7 Steel 150 112 45 32.8

8 Steel 250 125 45 44.4

9 Steel 160 90 45 41.0

10 Steel 224 95 45 49.3

11 Steel 500 140 45 61.0

12 Copper 1’000 40 129 92.5


250

Figure B.16: Current needed to achieve 85 °C with different meshes.

Steel meshes require less current to generate the same temper-

ature, followed by the samples with nickel meshes. The reason

both materials are ferromagnetic and therefore, affected by

hysteresis. Comparing both, samples with nickel meshes per-

form worst despite having better magnetic properties. The

reason is the skin effect, which generates a non-homogeneous

heating at the mesh wire. Samples with copper and aluminum

meshes require more power to reach the temperature range

under study due to the worse magnetic properties. Samples

with copper meshes perform better than with aluminum due

to their higher electrical conductivity.

Additionally, the particle modification is analyzed comparing

the temperature generated for the different weight contents,

when a fixed current of 130 A is applied. Results in steady are

reported in Figure B.17, obtaining a deviation in the measure-

ments lower than 1 °C.

0

10

20

30

40

50

60

70

80

1 2 3 4 5 6 7 8 9 10 11 12

Cu

rren

t n

eed

ed t

o a

chie

ve

85 ℃

Material number

Material sets:1 : Aluminum2, 3, 12: Copper4–6: Nickel7-11: Steel


INDUCTION

251

Figure B.17: Heating of adhesives with particles.

Samples with MagSilica show the best performance among the

particles tested. The reason is the special magnetic properties

of these particles, especially designed to generate heat by the

hysteresis effect combined with the small size of the particles.

This fact ensures that skin effect is not affecting the heating

process, which does not happen for the steel and especially

with iron particles. Nevertheless, the use of these particles

presents a drawback: that not more than 25 wt% can be mixed

with the adhesive due to the particle volume.

For the rest of the particle modifications, samples with steel

particles perform better than samples with iron and aluminum

particles. The reason is the skin effect, which affects more the

iron particles due to its higher electrical conductivity and

magnetics permeability (typically orders of magnitude higher

than for steel). In this context, iron particles are only heated at

20

30

40

50

60

70

80

90

0% 10% 20% 30% 40% 50%

Tem

per

atu

re [

°C]

wt%

MagSilica

Steel AS80

Steel RZ60

Al

Steel 80 μm

Iron 60 μm


252

the shell, thus reducing the efficiency of this material for in-

duction heating. Finally, aluminum particles are only residual-

ly heated. The reason is that in the induction heating of parti-

cles, hysteresis plays a dominant role. The induction of Eddy

currents requires a conductive domain. This would require

that all the particles would touch each other inside of the ad-

hesive forming an electrical conductive network. As this is not

the case, only ferromagnetic particles are significantly heated.

B.5.3.2: Mechanical performance of the paste adhesive

Finally, SLS test samples with non-electrical conductive sus-

ceptors made of GFRP are bonded. Details for the samples

preparation and mechanical testing applied are given in Ap-

pendix A.

8 samples with meshes and 5 samples with particles are select-

ed for this test, considering only steel and MagSilica for the

particles testing due to the better heating ability. The samples

are cured isothermally at about 100 °C for 60 minutes. In the

case of the sample with 15 wt% of MagSilica, as only 68 °C are

achieved at full power (384 A and 250’000 Hz), the curing pro-

cess is extended for 4 hours. The curing conditions and the

results are summarized in Table B.5 and Figure B.18, in which

the grey columns mean 50 % Cohesive 50 % Adherent fracture

mode, blue columns mean adherent fracture and black column

means cohesive fracture.


INDUCTION

253

Table B.5: Summary of SLS test of GFRP bonded samples. S

amp

le

Cu

rren

t [A

] C

uri

ng

tem

p.

[°C

]

Bo

nd

lin

e th

ick

nes

s

[mm

]

Sh

ear

stre

ng

th

[MP

a]

Fra

ctu

re m

od

e

Pu

re A

dh

esiv

e 0

100.

0 0.

15

18.2

± 1

.5

Ad

her

ent

3 C

u

74.4

10

1.5

0.36

14

.7 ±

1.7

A

dh

eren

t

5 N

i 34

.4

99.9

0.

27

16.7

± 0

.4

Ad

her

ent

7 S

t 28

.8

100.

7 0.

39

15.6

± 1

.9

Ad

her

ent

8 S

t 24

.0

99.9

0.

30

16.6

± 1

.5

Ad

her

ent

9 S

t 29

.6

100.

4 0.

34

15.3

± 1

.1

Ad

her

ent

11 S

t 31

.2

103.

7 0.

36

16.3

± 1

.2

Ad

her

ent

12 C

u

70.4

98

.0

0.23

11

.0 ±

0.5

C

oh

esiv

e

Mag

Sil

ica

15 %

38

4.0

68.3

0.

14

19.7

± 1

.1

Ad

her

ent/

Co

hes

ive

Mag

Sil

ica

25 %

32

0.0

99.9

0.

35

13.2

± 1

.4

Ad

her

ent/

Co

hes

ive

Ste

el 1

5 %

33

6.0

99.9

0.

19

14.4

± 0

.9

Ad

her

ent

Ste

el 2

5 %

25

2.0

98.1

0.

13

14.2

± 3

.8

Ad

her

ent

Ste

el 5

0 %

18

4.0

100.

7 0.

13

14.6

± 0

.8

Ad

her

ent


254

Figure B.18: SLS test results.

The samples with a mesh show mostly an adherent fracture

mode but a lower mechanical performance than the pure ad-

hesive samples. This means that the decrease of mechanical

performance is not due to a thermal degradation of the paste

adhesive by a temperature gradient between mesh and adhe-

sive. This fact can be explained by the stiffening effect of the

particles and meshes that lead to higher load peaks and there-

fore a decrease of the mechanical performance of the joint.

Another reason is the bigger bondline thickness due to the

diameter of the wire used in the meshes, thicker than 0.3 mm

in most of the cases. This fact is proved with CFRP bonded

samples in Section A.5.1.

The samples with particles show a certain decrease on the me-

chanical performance, especially with MagSilica having a

0

2

4

6

8

10

12

14

16

18

20

22S

hea

r st

ress

[M

Pa]

Meshes Particles


INDUCTION

255

mixed failure mode. The reason is the big volume of particles

mixed, mostly affecting the effective bonding area. It can be

observed that MagSilica samples with 15 wt% may not be fully

cured because the temperature generated was below the rec-

ommended by the supplier. This could lead to a tougher resin

and to a decrease of the load peaks, leading to a higher me-

chanical performance. Additionally, it is proved that the me-

chanical performance of samples with steel particles is not af-

fected by the amount of particles mixed in.

B.5.4: Conclusions

The modification of paste adhesives to bond non-electrical

conductive adherents by induction is investigated, observing:

Bonding non-conductive composite structures by induction

is possible by adding either ferromagnetic particles or

meshes.

Meshes show a superior heating ability than particles due

to the combination of induction of Eddy currents and hys-

teresis effect.

The electromagnetic material properties are affecting the

most the heating ability of the modified paste adhesive, es-

pecially the magnetic properties influencing the hysteresis

effect.

The meshes affect the mechanical performance of the joint

only depending on the diameter of the wires affecting the

bondline thickness.

The volume of particles mixed has an influence on the me-

chanical performance of the joint due to the loss of effective

bonding area.


256

257

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269

List of Tables

Table 2.1: Parameters used to cure SLS samples. ...................... 22 Table 2.2: Results of the SLS test. ................................................ 22 Table 2.3: Time necessary to heat up a CFRP plate. ................. 23 Table 2.4: Summary of the performance of the heating

methods investigated. ................................................................... 24 Table 3.1: Results of the overall heat enthalpy measurement. 34 Table 3.2: n-th order model parameters by lineal regression. . 37 Table 3.3: Parameters of the n-th order model. ......................... 38 Table 3.4: n-th order model parameters in the chemical

controlled part. .............................................................................. 39 Table 3.5: Fitted parameters at the DiBenedetto equation. ...... 43 Table 3.6: αonset at different temperatures. ................................... 43 Table 3.7: Fitted parameters of αonset. ........................................... 44 Table 3.8: Values of C1 and KDonset. ............................................... 45 Table 3.9: Summary of all the fitted parameters. ...................... 45 Table 3.10: Summary of all the fitted parameters. .................... 47 Table 3.11: Samples for cure kinetics model validation. .......... 50 Table 3.12: Results of the overall heat enthalpy. ....................... 52 Table 3.13: Cure kinetics parameters of the LME 10625 / LME

10626................................................................................................ 52 Table 4.1: Thermal degradation model parameters. ................. 60 Table 4.2: Summary of samples used for validation................. 62 Table 4.3: Void formation with different curing cycles. ........... 67 Table 4.4: Fitting parameters of the thermal degradation model

of the LME 10625 / LME 10626. ................................................... 70 Table 4.5: Void formation with different curing cycles. ........... 72 Table 5.1: Summary of relations applied in the model. ........... 84 Table 5.2: Results from the validation of the model. ................ 92 Table 5.3: Sensitivity analysis of adhesive properties. ............. 95

LIST OF TABLES

270

Table 5.4: Sensitivity analysis of the processing parameters. .. 97 Table 5.5: Sensitivity analysis of adherent properties. ............. 98 Table 5.6: Electrical conductivity of laminates with different

number of layers. ......................................................................... 101 Table 5.7: Sensitivity analysis of adherent properties. ........... 101 Table 6.1: Degree of cure of the samples measured by DSC. 107 Table 6.2: Summary of optical measurements. ........................ 115 Table 6.3: Storage modulus at 20 °C and Tg. ............................ 118 Table 6.4: Summary of results in 3 point bending test. .......... 119 Table 6.5: Results of single lap shear test. ................................ 120 Table 6.6: Techniques to validate paste adhesives. ................. 123 Table 7.1: Effect of T1 [°C] on the curing process. ................... 133 Table 7.2: Effect of t1 [°C] on the curing process. .................... 136 Table 7.3: Effect of T2 [°C] on the curing process. ................... 138 Table 7.4: Summary of SLS test. ................................................ 140 Table 8.1: Microscopy test of the optimal curing cycle. ......... 147 Table 8.2: SLS test of the optimal curing process. ................... 149 Table 8.3: Microscopy of samples cured with the optimal

curing cycle. ................................................................................. 151 Table 8.4: SLS test of the optimal curing process. ................... 151 Table 9.1: Summary of the bondlines of the demonstrator.... 161 Table 9.2: Validation of the positioning of the parts where the

maximum deviation according to RUAG specifications are 1

mm. ............................................................................................... 169 Table 9.3: Validation of the orientation of the parts where the

maximum value according to RUAG specifications are 1°. ... 170 Table 9.4: Validation of the bondline thickness where the

maximum value according to RUAG specifications are 0.3

mm. ............................................................................................... 171 Table 10.1: Temperature difference inside of the bondline. .. 181 Table 10.2: Temperature difference in the measuring point, x=

54 cm. ............................................................................................ 181 Table 10.3: Temperature gradients with different cooling. ... 185 Table 10.4: Temperature distribution analysis summary. ..... 187

LIST OF TABLES

271

Table 10.5: Validation of the positioning of the parts where the

maximum deviation according to RUAG specifications are 1

mm. ............................................................................................... 193 Table 10.6: Validation of the orientation of the parts where the

maximum value according to RUAG specifications are 1°. ... 194 Table 10.7: Quality control of the second demonstrator where

the maximum value according to RUAG specifications are 0.3

mm for the bondline thickness, a degree of cure higher than 95

% and a void content lower than 2 %. ...................................... 195 Table A.1: SLS for different surface treatments. ...................... 225 Table A.2: Fracture mode of SLS samples. ............................... 226 Table B.1: Material properties of the test rig. ........................... 233 Table B.2: Summary of dry air properties. ............................... 236 Table B.3: Particles used for adhesive modification. .............. 247 Table B.4: Geometric parameters of the meshes used. ........... 249 Table B.5: Summary of SLS test of GFRP bonded samples. .. 253

LIST OF TABLES

272

273

List of Figures

Figure 1.1: Image of the carbon fiber reinforced polymer

fuselage section of the Boeing 787 [2]. .......................................... 4 Figure 1.2: Adhesive bonding introduces a uniform load. ........ 5 Figure 1.3: Adhesive sample with high void formation. ........... 9 Figure 2.1: Hot press heats the adherents by conduction. ....... 14 Figure 2.2: Oven heats all the components ................................ 15 Figure 2.3: Hysteresis loop of a ferromagnetic susceptor. ....... 18 Figure 2.4: Induction heats the CFRP area close to the coil. .... 20 Figure 2.5: SLS results. .................................................................. 22 Figure 3.1: DSC curve with a heating rate of 10 °C/min. ......... 31 Figure 3.2: DSC measurements. .................................................. 35 Figure 3.3: Conversion rate of the different measurements. ... 36 Figure 3.4: Measurement and fitting of ln (k). ........................... 37 Figure 3.5: Comparison of n-th order and experimental data. 38 Figure 3.6: n-th order model in the chemical controlled part. . 40 Figure 3.7: Chemical and diffusion controlled parts. ............... 41 Figure 3.8: Conversion rate vs. Tg in the measurements. ......... 42 Figure 3.9: Fitting of αonset .............................................................. 44 Figure 3.10: n-th order model fitting at the chemical region. .. 46 Figure 3.11: Fitting of the final model. ....................................... 47 Figure 3.12: Scheme for the modeling of the cure kinetics. ..... 48 Figure 3.13: The temperature is increased from T1 to T2. ......... 49 Figure 3.14: The process will have a faster curing process. ..... 49 Figure 3.15: Validation of the cure kinetics model. .................. 51 Figure 3.16: Model validation for the cure kinetics model. ..... 53 Figure 4.1: Fitting of the thermal degradation model. ............. 61 Figure 4.2: Validation of the heating cycles 1 to 3. .................... 63 Figure 4.3: Validation of the heating cycles 4 to 6. .................... 63 Figure 4.4: Validation of the heating cycles 7 to 9. .................... 64

LIST OF FIGURES

274

Figure 4.5: Modeling of samples 1. ............................................. 65 Figure 4.6: Experimental and predicted degradation............... 66 Figure 4.7: Void content vs. relative evaporated mass. ............ 68 Figure 4.8: Summary of the model accuracy. ............................ 69 Figure 4.9: Experimental and modeled evaporation curves. ... 71 Figure 4.10: Relation between voids and evaporated mass. .... 73 Figure 4.11: Modeled and experimental void formation. ........ 74 Figure 5.1: Scheme of the induction setup modeling. .............. 78 Figure 5.2: Validation points for steady state and transient. ... 85 Figure 5.3: The best estimation for the efficiency is ηcoil= 0.9. .. 86 Figure 5.4: Temperature validation at steady state. ................. 87 Figure 5.5: Temperature for different values of λx. ................... 88 Figure 5.6: Temperature for different values of λy. ................... 88 Figure 5.7: Heating and cooling at the overlap center. ............ 89 Figure 5.8: Heating at Point A for different hc. .......................... 90 Figure 5.9: Cooling at Point A for different hc. .......................... 90 Figure 5.10: Heating at different distances. ............................... 91 Figure 5.11: Cooling at different distances. ............................... 91 Figure 5.12: Distance between the lower coil and the plate. . 100 Figure 6.1: TGA analysis for epoxy and hardener. ................. 108 Figure 6.2: Sample cured at 80 °C (Avg. void content 1.6 %).

....................................................................................................... 109 Figure 6.3: Sample cured at 100 °C (Avg. void content 1.4%).

....................................................................................................... 109 Figure 6.4: Sample cured at 120 °C (Avg. void content 2.1 %)

....................................................................................................... 110 Figure 6.5: Sample cured at 140 °C (Avg. void content 21.4%).

....................................................................................................... 110 Figure 6.6: Sample cured at 160 °C (Avg. void content 33.5 %).



....................................................................................................... 112

LIST OF FIGURES

275

Figure 6.9: Void content measurement. ................................... 112 Figure 6.10: Void content of samples (80 °C to 120 °C). ......... 113 Figure 6.11: Average bubbles diameter of samples. ............... 114 Figure 6.12: Relative density of samples. ................................. 114 Figure 6.13: DMA measurement at 100 °C. .............................. 116 Figure 6.14: Storage modulus measured by DMA.................. 117 Figure 6.15: Storage modulus from DMA at 20 °C. ................ 117 Figure 6.16: Flexural modulus of samples. .............................. 119 Figure 6.17: Flexural strength of samples. ............................... 119 Figure 6.18: SLS test for CFRP bonded samples. .................... 120 Figure 6.19: Void content compared to SLS test...................... 123 Figure 6.20: Results referenced to supplier’s curing............... 124 Figure 7.1: Two step heating process. ....................................... 129 Figure 7.2: Impact of void formation on the shear strength. . 141 Figure 8.1: Optimization scheme of the curing process. ........ 146 Figure 8.2: Optimal curing cycle of the LME 10049-4 / LMB

6687-2. ........................................................................................... 147 Figure 8.3: Recommended and optimal curing cycles of the

LME 10049-4 / LMB 6687-2......................................................... 149 Figure 8.4: Optimal curing cycle of the LME 10625 / LME

10626.............................................................................................. 150 Figure 8.5: Comparison between recommended and optimal

curing cycles of the LME 10625 / LME 10626. ......................... 152 Figure 9.1: Approach to optimize the curing cycle in complex

structures. ..................................................................................... 157 Figure 9.2: Image of the Dornier 228. ....................................... 158 Figure 9.3: Geometry of the aileron of the Dormier 228. ....... 159 Figure 9.4: Detail of the rear part of the aileron. ..................... 159 Figure 9.5: Layup sequence of the side of the skins. .............. 160 Figure 9.6: Scheme of the procedure to heat the components.

....................................................................................................... 162 Figure 9.7: Assembly of the T-Stringer ..................................... 162 Figure 9.8: Assembly of the T-stringer to the lower skin. ...... 163 Figure 9.9: Structure to fix the parts during the process. ....... 163

LIST OF FIGURES

276

Figure 9.10: Bonding process of a rib and the inner doubler.164 Figure 9.11: Upper subassembly bonded to the ribs. ............. 164 Figure 9.12: Subassembly with the ribs and the upper skin. . 165 Figure 9.13: Assembly of the lower skin to the ribs. ............... 165 Figure 9.14: Assembly of a C-spar to the inner doubler......... 166 Figure 9.15: Assembly of the C-spar to the lower skin........... 166 Figure 9.16: Outer doubler bonded to the subassembly. ....... 167 Figure 9.17: Demonstrator. ......................................................... 168 Figure 10.1: The 4 representative bondlines. ........................... 174 Figure 10.2: Measurement in bondline A. ................................ 176 Figure 10.3: Temperature distribution in bondline A. ........... 176 Figure 10.4: Measurement in bondline B. ................................ 177 Figure 10.5: Temperature distribution in bondline B. ............ 177 Figure 10.6: Measurement in bondline C. ................................ 178 Figure 10.7: Temperature distribution in bondline C. ............ 179 Figure 10.8: Measurement in bondline D. ................................ 180 Figure 10.9: Temperature distribution in bondline D. ........... 180 Figure 10.10: Edge effect measurement. ................................... 183 Figure 10.11: Thickness variation effect. .................................. 184 Figure 10.12: Fan positioning of forced convection cooling. . 185 Figure 10.13: Effect of the distance between the coil and the

susceptor. ...................................................................................... 186 Figure 10.14: Optimum curing cycle for a ∆T in the bondline

between 5 and 25 °C. ................................................................... 188 Figure 10.15: Duration of the optimal curing cycle with a

temperature difference. .............................................................. 189 Figure 10.16: Optimal curing cycle at bondline type A. ......... 190 Figure 10.17: Optimal curing cycle at bondline type B. ......... 191 Figure 10.18: Optimal curing cycle at bondline type C. ......... 191 Figure 10.19: Optimal curing cycle at bondline type D. ......... 192 Figure 11.1: Forced and natural convection. ............................ 204 Figure A.1: OoA produces a bad surface quality. ................... 214 Figure A.2: CFRP layup scheme for the CFRP processing. ... 215 Figure A.3: Adhesive failure in a wedge sample. ................... 218

LIST OF FIGURES

277

Figure A.4: Cohesive failure in a wedge sample. .................... 219 Figure A.5: Adherent failure in a wedge sample. ................... 220 Figure A.6: Mixed failure in a wedge sample. ........................ 220 Figure A.7: Effect of bondline and adherent thicknesses. ...... 222 Figure A.8: SLS test for different surface treatments. ............. 226 Figure B.1: Induction setup used. ............................................. 230 Figure B.2: Test rig designed to bond SLS samples. ............... 231 Figure B.3: Geometry of the test setup. .................................... 232 Figure B.4: Conditions for the calculation of hc. ...................... 236 Figure B.5: Geometry to validate natural convection. ............ 238 Figure B.6: Efficiency of the coil. ............................................... 238 Figure B.7: Temperature distribution of the C-spar. .............. 239 Figure B.8: Heating and cooling processes of the C-spar. ..... 239 Figure B.9: Assessment of the convection coefficient. ............ 240 Figure B.10: Results for different currents applied. ................ 241 Figure B.11: Heating and cooling with natural convection. .. 242 Figure B.12: Cycle to produce samples for validation. ........... 243 Figure B.13: Scheme to produce samples by induction.......... 244 Figure B.14 Sample for the microscopy measurement. .......... 245 Figure B.15: Temperature generation in modified adhesives.

....................................................................................................... 248 Figure B.16: Current needed to achieve 85 °C with different

meshes. ......................................................................................... 250 Figure B.17: Heating of adhesives with particles. ................... 251 Figure B.18: SLS test results. ...................................................... 254

LIST OF FIGURES

278

279

About the Author


was born on July 20th, 1985

in Barcelona, Spain where

he spent most of all his edu-

cation time. He enrolled at

the Escola Tècnica Superior

d’Enginyeria Industrial de

Barcelona (ETSEIB), in 2003.

In 2008 he made his Master

Thesis at the Swiss Federal

Institute of Technology

(ETH Zürich) in the frame of

vibration damping as a part of the Erasmus exchange pro-

gram. In 2009 he obtained his Diploma in Industrial Engineer-

ing from the Universitat Politècnica de Catalunya (UPC) with

the specialization in Vehicles and Transportation Science. He

continued working as research assistant in the frame of novel

approaches for structural damping at the Centre of Structure

Technologies, in the Department of Mechanical and Process

Engineering at the Swiss Federal Institute of Technology (ETH

Zürich) in the early 2009. He began his doctoral studies on

September 1st, 2009, at the Centre of Structure Technologies, in

the Department of Mechanical and Process Engineering at the

Swiss Federal Institute of Technology (ETH Zürich) in Swit-

zerland and defended his thesis, contained in this volume, on

the 11th of April of 2014.

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