Soldering & Surface Mount Technology, Volume 14, Number 1, 2002

ISSN 0954-0911

Soldering & Surface Mount Technology Volume 14, Number 1, 2002

Computational modelling Guest Editor: Chris Bailey This issue is part of a comprehensive multiple access information service comprising:

Contents

Paper format Soldering & Surface Mount Technology includes three issues in traditional paper format. The contents of this issue are detailed below.

Internet Online Publishing with Archive, Active Reference Linking, Non-article Content, Institution-wide Licence, E-mail Alerting Service and Usage Statistics. Access via the Emerald Web site: http://www.emeraldinsight.com/ft See overleaf for full details of subscriber entitlements.

2

Access to Soldering & Surface Mount Technology online

3

Editorial advisory board

4

Abstracts & keywords

Analysis on solder ball shear testing conditions with a simple computational model

6

French abstracts

S.W. Ricky Lee and Xingjia Huang

8

German abstracts

38

CFD modelling of the flow field inside a reflow oven Hao Yu and Jorma Kivilathti

45

10

Editorial

11

Correlation of solder paste rheology with computational simulations of the stencil printing process

18

49

Optimisation modelling for flip-chip solder joint reliability S. Stoyanov, C. Bailey and M. Cross

59

Internet commentary

61

Book review

R. Durairaj, G.J. Jackson, N.N. Ekere, G. Glinski and C. Bailey

61

SMART Group

Solder paste reflow modeling

62

Association news

62

New products

65

Industry news

67

Appointments

67

Publications

68

International diary

70

Note from the publisher

S.H. Mannan

24

Numerical modelling of scanned beam laser soldering of fine pitch packages P.M. Beckett, A.R. Fleming, J.M. Gilbert and D.G. Whitehead

30

A simplified model of the reflow soldering process David C. Whalley and Stuart M. Hyslop

Soldering & Surface Mount Technology ISSN 0954-0911 # 2002 MCB University Press

Guest Editor Chris Bailey

Soldering & Surface Mount Technology provides an authoritative and independent forum for the exchange of information pertaining to the research, development and application of all branches of soldering and surface mount technologies. It makes an important contribution to the technical body of knowledge and expertise in this vital area.

Editorial Advisory Board David Boswell Associate Consultant, Surface Mount Club, UK

Internet services available worldwide on URL http://www.mcb.co.uk

Editor David C. Whalley

Henrik Hvims Reipur Technology A/S, Copenhagen, Denmark

Emerald 60/62 Toller Lane, Bradford, West Yorkshire, England BD8 9BY Telephone: (44) 1274 777700 Fax: (44) 1274 785200

Department of Manufacturing Engineering, University of Loughborough, Leicestershire, UK

Jennie S. Hwang H-Technologies Group, Cleveland, Ohio, USA

Regional offices

Founding Editor William Goldie

Associate Editor Trevor Galbraith Internet editor Brian Ellis E-mail: ssmt-comment@ protonique.com Book Review Editors Brian Ellis, Protonique, SA Bob Willis, Electronic Presentation Services, UK Managing Editor Sharon Parkinson Produced by Alden Multimedia, Northampton Advertising Agent Trevor Galbraith Electronics Marketing Ltd PO Box 2521, 65 High Street, Glastonbury, Somerset BA6 9DS Tel: +44 (0) 1458 832137 Fax: +44 (0) 1458 832143 E-mail: [email protected] Indexed and abstracted in Chemical Abstracts Current Contents1/ Engineering, Computing and Technology CSA Mechanical & Transport Engineering Abstracts (Selective) ISI Alerting Services Materials Science Citation Index1 SciSearch1

Peter Grundy SMART Group, UK

John Lau Agilent Technologies Ltd, San Jose, California, USA Ning-Cheng Lee Vice President, Indium Technology of America, Clinton, New York, USA S.-W. Ricky Lee Department of Mechanical Engineering, Hong Kong University of Science & Technology, Hong Kong Johan Liu IVF, Sweden A. Martinez Chase Advanced Technologies Ltd, West Yorkshire, UK Kyung W. Paik KAIST, South Korea W.J. Plumbridge Materials Department, Open University, Milton Keynes, UK Wallace Rubin International Consultant in Soldering Technology, Middlesex, UK K.C. Teo Philips Centre for Manufacturing, Singapore Laura Turbini Centre for Microelectronics Assembly and Packaging, University of Toronto, Toronto, Canada Christopher S. Welch College of William and Mary, Williamsburg, Virginia, USA Bob Willis Process Engineering Consultant, Electronic Presentation Services, Essex, UK Sung Yi School of Mechanical and Production Engineering, Nanyang Technological University, Singapore

For North America Emerald, 44 Brattle Street, 4th Floor, Cambridge, MA 02138, USA Telephone: Toll free +1 888 622 0075 Fax: +1 617 354 6875 E-mail: [email protected] For South East Asia Emerald, 7-2, 7th Floor, Menara KLH, Bandar Puchog Jaya, 47100, Puchong, Selangor, Malaysia Telephone: 603-5766 009; Fax: 603-5766 007 E-mail: [email protected] For Japan Emereald, 256-25 Kounosudai, Nagareyama City, Chiba Prefecture, Japan 270-0102 Telephone: 81 (0) 471 54 6245 Fax: 81 471 54 6268 E-mail: [email protected]

Please contact the following for further information: Sample copy requests and missing issue claims* to the Customer Help Desk at Bradford (44) 1274 785280 or your regional office. *Non-delivery of copies must be notified to our Customer Help Desk within four months of despatch if substitute copies are to be issued. All order and subscription enquiries to the Subscription Services Department at Bradford or your regional office. Reprint service enquiries to Dave Manks, Bradford Copyright permissions to Charles Parker, Bradford Printed by Printhaus Graphique, 2 North Postway Close, Round Spinney, Northampton NN3 8RQ

No part of this journal may be reproduced, stored in a retrieval system, transmitted in any form or by any means electronic, mechanical, photocopying, recording or otherwise without either the prior written permission of the publisher or a licence permitting restricted copying issued in the UK by The Copyright Licensing Agency and in the USA by The Copyright Clearance Center. No responsibility is accepted for the accuracy of information contained in the text, illustrations or advertisements. The opinions expressed in the articles are not necessarily those of the editor or the publisher. Pre-press activities, comprising origination and pagination, associated with the publication of management journals

[3]

Abstracts & keywords

Correlation of solder paste rheology with computational simulations of the stencil printing process R. Durairaj, G.J. Jackson, N.N. Ekere, G. Glinski and C. Bailey Keywords Solder pastes, Rheology, Stencils Soldering technologies continue to evolve to meet the demands of the continuous miniaturisation of electronic products, particularly in the area of solder paste formulations used in the reflow soldering of surface mount devices. Stencil printing continues to be a leading process used for the deposition of solder paste onto printed circuit boards (PCBs) in the volume production of electronic assemblies, despite problems in achieving a consistent print quality at an ultra-fine pitch. In order to eliminate these defects a good understanding of the processes involved in printing is important. Computational simulations may complement experimental print trials and paste characterisation studies, and provide an extra dimension to the understanding of the process. The characteristics and flow properties of solder pastes depend primarily on their chemical and physical composition and good material property data is essential for meaningful results to be obtained by computational simulation. This paper describes paste characterisation and computational simulation studies that have been undertaken through the collaboration of the School of Aeronautical, Mechanical and Manufacturing Engineering at Salford University and the Centre for Numerical Modelling and Process Analysis at the University of Greenwich. The rheological profile of two different paste formulations (lead and lead-free) for sub 100 micron flip-chip devices are tested and applied to computational simulations of their flow behaviour during the printing process.

Solder paste reflow modeling S H Mannan Keywords Solder paste, Computational fluid dynamics Solder paste printing and reflow are well established processes for producing solder joints in electronic assemblies. Solder paste consists of a dense suspension of solder particles in a liquid medium (vehicle) that acts as an oxide reducing agent (flux) during reflow, cleaning the metal surfaces of oxides. This paper reports on attempts to model the physical and chemical processes occurring during solder paste reflow using computational fluid dynamics (CRD). Axisymmetric, 2 dimensional and 3-dimensional models are described, and a method of reproducing oxide-like behaviour in these models in introduced.

Soldering & Surface Mount Technology 14/1 [2002] Abstracts & keywords q MCB UP Limited [ISSN 0954-0911]

[4]

Numerical modelling of scanned beam laser soldering of fine pitch packages P.M. Beckett, A.R. Fleming, J.M. Gilbert and D.G. Whitehead Keywords Soldering, Lasers, Electronics assembly Laser soldering provides a useful tool for the electronics manufacturer and has found a number of successful industrial applications. The laser provides highly controllable localised heating in a manner similar to hand soldering and has distinct advantages over other soldering methods. However, the heating processes that occur in laser soldering are complex making it difficult to predict whether a particular operation will be successful. Numerical modelling provides a valuable tool in answering such questions but the modelling process and the assumptions upon which it is based must be understood if the results are to be reliable. This paper explains the assumptions made in deriving a suitable model and describes the use of a commercial finite element modelling package to aid the understanding of laser soldering processes, with a particular emphasis on single and multi-pass scanned beam soldering operations.

A simplified model of the reflow soldering process David C. Whalley and Stuart M. Hyslop Keywords Soldering, Reflow Previous models of temperature development during the reflow soldering process have typically used commercially available, general purpose, finite difference/finite element modelling tools to create detailed three dimensional representations of both the product and of the reflow furnace. Such models have been shown to achieve a high degree of accuracy in predicting the temperatures a particular PCB design will achieve during the reflow process, but are complex to generate and analysis times are long, even when using modern high performance computer workstations. This paper will report on the development of a simplified model of the process, which uses less complex representations of both the product and the process, together with a simple numerical solver developed specifically for this application, whilst achieving an accuracy comparable with more detailed models. In the simplified model, the product is divided into elements, which are represented using a two-dimensional mesh of thermal conductances linking thermal masses. The values of these conductances and masses are calculated based on the averaged properties of the PCB material and attached components within the area of each of the elements. The representation of the specific reflow furnace is based on measurements of the temperature and level of thermal convection at each point along the length of the furnace, thereby avoiding the necessity of making detailed measurements of the furnace geometry and air flow velocities. The combination of these two simplification techniques allow the reduction of analysis time for a relatively simple PCB from in the order of an hour on a high performance Unix workstation to under a second on a Pentium class PC running Microsoft Windows.

Abstracts & keywords Soldering & Surface Mount Technology 14/1 [2002] 4–5

CFD modelling of the flow field inside a reflow oven

Optimisation modelling for flip-chip solder joint reliability

Hao Yu and Jorma Kivilathti

S. Stoyanov, C. Bailey and M. Cross

Keywords Lead-free soldering, Reflow, Computational fluid dynamics

Keywords Optimization, Flip chip, Underfill, Reliability

Because of new requirements related to the employment of lead-free manufacturing and the diversity of components and metal finishes on high density printed circuit boards, better understanding and control of the reflow process is required in order to achieve acceptable yields and reliability of SMT assemblies. Accurate control of the temperature distributions within components and boards during the reflow process is one of the major requirements, especially in lead-free assembly. This paper outlines a scheme for reflow modelling and presents an oven-level model of the steady state flow-field inside a reflow oven, which will be needed in subsequent transient analysis and small-scale modelling. The model is constructed by utilising the advanced computational fluid dynamics (CFD) technology using commercial software. The computational results are discussed and compared with measured data.

This paper details and demonstrates integrated optimisationreliability modelling for predicting the performance of solder joints in electronic packaging. This integrated modelling approach is used to identify efficiently and quickly the most suitable design parameters for solder joint performance during thermal cycling and is demonstrated on flip-chip components using “no-flow” underfills. To implement “optimisation in reliability” approach, the finite element simulation tool – PHYSICA, is coupled with optimisation and statistical tools. This resulting framework is capable of performing design optimisation procedures in an entirely automated and systematic manner.

Analysis on solder ball shear testing conditions with a simple computational model S.W. Ricky Lee and Xingjia Huang Keywords Plastic ball grid array, Solder joints, Finite element analysis, Reliability This paper introduces a simple computational model for the analysis on the solder ball shear testing conditions. Both two-dimensional (2-D) and three-dimensional (3-D) finite element models are used to investigate the effect of shear ram speed on the solder ball shear strength of plastic ball grid array (PBGA) packages. An effective thickness is identified for the 2-D finite element analysis. By using this effective thickness as a scale factor, it is shown that the 2D model is feasible for the study of 3-D problems. The computational model is validated by experimental data in terms of load-displacement curves. The results from both testing and modeling indicate that the shear ram speed has a substantial effect on the solder ball shear strength. In general, faster ram speed can result in higher ball shear strength. Therefore, the characterization of solder ball shear strength is loading rate-dependent.

[5]

Abstracts & keywords

Corre´lation entre la rhe´ologie des paˆtes de brasage et les simulations automatiques du processus de se´rigraphie

Modelage nume´rique du brasage au rayon laser scanne´ pour les assemblages de tre`s petites dimensions

R. Durairaj, G.J. Jackson, N.N. Ekere, G. Glinski et C. Bailey

P. M Beckett, A.R. Fleming, J. M. Gilbert et D. G. Whitehead

Mots-cle´s Paˆtes de brasage, Rhe´ologie, Stencils

Mots-cle´s Brasage, Lasers, Assemblage e´lectronique

Les techniques de brasage e´voluent sans cesse afin de re´pondre aux demandes de miniaturisation continue des produits e´lectroniques, surtout dans le domaine des formulations de paˆtes de brasage utilise´es dans le brasage par fusion des dispositifs monte´s en surface. La se´rigraphie continue a` repre´senter un processus important, qui est utilise´ pour de´poser de la paˆte de brasage sur les plaquettes de circuits imprime´s dans la prodution en masse d’assemblages e´lectroniques, malgre´ les difficulte´s qu’il y a d’obtenir une qualite´ d’impression uniforme avec des dimensions minimes. Afin d’e´liminer ces de´fectuosite´s, il importe de bien comprendre les processus implique´s dans l’impression. Les simulations automatiques peuvent comple´ter les essais d’impression expe´rimentaux et les e´tudes de caracte´risation de la paˆte, et permettre de mieux comprendre le processus. Les caracte´ristiques et proprie´tes d’e´coulement des paˆtes de brasage de´pendent tout d’abord de leur composition chimique et physique, et il est essentiel de disposer de bonnes donne´es sur les proprie´te´s des mate´riaux pour pouvoir obtenir des re´sultats significatifs par la simulation automatique. L’article que voici de´crit les e´tudes de caracte´risation et de simulation automatique des paˆtes, qui ont e´te´ entreprises en collaboration entre l’Ecole d’Inge´nieurs Ae´ronautiques, Me´caniques et de Fabrication de l’Universite´ de Salford et le Centre de Modelage Nume´rique et d’Analyse des Processus de l’Universite´ de Greenwich. Le profil rhe´ologique de deux formulations de paˆtes diffe´rentes (avec et sans plomb) pour les dispositifs flip-chip de moins de 100 microns est mis a` l’essai et applique´ a` la simulation automatique de leur comportement d’e´coulement au cours du processus d’impression.

Le brasage au laser repre´sente un outil utile pour le fabricant de pie`ces e´lectroniques; il en existe toute une se´rie d’applications industrielles re´ussies. Le laser permet un e´chauffement localise´ extreˆmement controˆlable, selon une me´thode qui ressemble au brasage manuel; il offre de nets avantages par rapport aux autres me´thodes de brasage. Cependant, les processus d’e´chauffement qui se produisent au cours du brasage au laser sont complexes et cre´ent des difficulte´s, lorsqu’il s’agit de pre´dire si une ope´ration particulie`re va re´ussir. Le modelage nume´rique est un outil pre´cieux qui permet de re´soudre ces questions, mais il faut bien comprendre le processus de modelage et les suppositions sur lesquelles il se fonde, si l’on de´sire obtenir des re´sultats fiables. L’article explique les suppositions qui sont faites lorsque l’on de´rive un mode`le approprie´; il de´crit l’utilisation d’un programme commercial de modelage a` e´le´ments finis, qui permet de comprendre les processsus de brasage au laser, en mettant l’accent tout particulie`rement sur les ope´rations de brasage au rayon scanne´ avec passage unique et re´pe´te´.

Modelage de la fusion des paˆtes de brasage S H Mannan Mots-cle´s Paˆtes de brasage, Dynamique automatique des fluides L’impression et la fusion des paˆtes de brasage sont des processus bien e´tablis pour produire des joints de brasage dans les assemblages e´lectroniques. La paˆte de brasage est une suspension dense de particules de brasage dans un milieu (ve´hicule) liquide qui sert d’agent de re´duction de l’oxyde (e´coulement) au cours de la fusion, e´liminant ainsi les oxydes des surfaces me´talliques. L’article que voici rend compte des essais visant a` modeler les processus physiques et chimiques qui se produisent au cours de la fusion des paˆtes de brasage, qui se servent de la Dynamique Automatique des Fluides (Computational Fluid Dynamics – CFD). Il de´crit des mode`les axisymme´triques a` 2 et 3 dimensions et pre´sente une me´thode permettant de reproduire le comportement semblable a` l’oxyde dans ces mode`les. ‘

Soldering & Surface Mount Technology 14/1 [2002] French abstracts q MCB UP Limited [ISSN 0954-0911]

[6]

Un mode`le simplifie´ du processus de brasage par fusion David C. Whalley et Stuart M. Hyslop Mots-cle´s Brasage, Fusion Les mode`les pre´ce´dents de mise au point de la tempe´rature au cours du processus de brasage par fusion utilisent, de manie`re typique, des outils de modelage disponibles dans le commerce, destine´s a` des applications ge´ne´rales, et ayant des e´le´ments finis/a` diffe´rences finies, pour cre´er des repre´sentations tridimensionnelles du produit ainsi que du four a` fusion. Il a e´te´ prouve´ que ces mode`les permettent d’obtenir un degre´ e´leve´ de pre´cision, lorsqu’il s’agit de pre´dire les tempe´ratures qu’un mode`le particulier de plaquette a` circuits imprime´s atteindra au cours du processus de fusion, mails ils sont complexes a` engendrer et les temps d’analyse sont longs, meˆme lorsque l’on utilise des postes de travail automatise´s modernes a` performance e´leve´e. L’article de´crit la mise au point d’un mode`le de processus simplifie´ qui se sert de repre´sentations moins complexes pour le produit et le processus; il de´crit e´galement un “solver” nume´rique simple, mis au point tout spe´cialement pour cette application, et qui permet un degre´ de pre´cision comparable a` celui des mode`les plus de´taille´s. Dans le mode`le simplifie´, le produit est divise´ en e´le´ments, qui sont repre´sente´s au moyen d’un re´seau bidimensionnel de conductances thermiques reliant les masses thermiques. Les valeurs de ces conductances et masses sont calcule´es sur base des proprie´te´s ge´ne´ralise´es qui caracte´risent le mate´riau de la plaquette a` circuit imprime´ et des composantes attache´es aux alentours de chacun des e´le´ments. La repre´sentation du four de fusion spe´cifique se fonde sur les mesurages de la tempe´rature et du niveau de convection thermique a` chacun des points situe´s le long de la longueur du four, ce qui permet de ne pas devoir proce´der a` des mesurages de´taille´s de la ge´ome´trie du four et des vitesses du flux d’air. La combinaison de ces deux techniques de simplification permet de re´duire le temps d’analyse, pour une plaquette a` circuit imprime´ relativement simple, d’environ une heure, sur un poste de travail Unix de performance e´leve´e, a` moins d’une seconde, sur un ordinateur de cate´gorie Pentium utilisant Microsoft Windows.

Abstracts & keywords Soldering & Surface Mount Technology 14/1 [2002] 6–7

Modelage CFD d’un champ d’e´coulement a` l’inte´rieur d’un four de fusion

Optimalisation du modelage pour la fiabilite´ des joints de brasage flip-chip

Hao Yu et Jorma Kivilathti

S. Stoyanov, C. Bailey et M. Cross

Mots-cle´s Brasage sans plomb, Fusion, Dynamique automatique des fluides

Mots-cle´s Optimalisation, Flip-chip, “Underfills”, Fiabilite´

En raison des nouvelles exigences concernant l’emploi de produits sans plomb et de la diversite´ des composantes et des finis me´talliques sur les cartes a` circuit imprime´ de haute densite´, il est ne´cessaire de mieux comprendre et controˆler le processus de fusion, afin d’obtenir une production acceptable et fiable dans la construction des assemblages SMT. Le controˆle pre´cis de la re´partition des tempe´ratures dans les composantes et les plaquettes, au cours du processus de fusion, est l’une des ne´cessite´s principales, surtout dans les assemblages sans plomb. L’article que voici de´crit un programme permettant de modeler la fusion; il pre´sente un mode`le de champs de flux en re´gime permanent, travaillant comme un four, a` l’inte´rieur d’un four de fusion, qui sera requis par la suite dans l’analyse transitoire et le modelage a` e´chelle re´duite. Pour construire le mode`le, on utilise la dynamique automatique des fluides, en se servant d’un logiciel commercial. L’article discute les re´sultats nume´riques et les compare aux donne´es mesure´es.

L’article de´crit en de´tail et de´montre le modelage inte´gre´, avec fiabilite´ optimalise´e, qui permet de pre´dire la performance des joints de brasage dans les assemblages e´lectroniques. Ce modelage inte´gre´ permet d’identifier de manie`re efficace et rapide les parame`tres de dessin les mieux approprie´s pour la performance des joints de brasage pendant le cycle thermique; ceci est de´montre´ sur des composantes flip-chip au moyen d’underfills “no-flow”. Pour mettre en pratique la me´thode “d’optimalisation avec fiabilite´”, l’outil de simulation des e´le´ments finis – PHYSICA – est ajoute´ a` l’optimalisation et a` des outils statistiques. La structure qui en re´sulte est capable de performer des proce´de´s d’optimalisation du dessin de manie`re entie`rement automatique et syste´matique.

Analyse des conditions d’essai de cisaillement des perles de brasage a` l’aide d’un mode`le de calcul simple S.W. Ricky Lee et Xingjia Huang Mots-cle´s “Plastic ball grid array”, Joint de brasage, Analyse des e´le´ments finis, Fiabilite´ L’article pre´sente un mode`le de calcul simple pour l’analyse des conditions d’essai de cisaillement des perles de brasage. Des mode`les a` e´le´ments finis bidimensionnels (2D) et tridimensionnels (3D) sont utilise´s pour examiner l’effet que la vitesse du coulisseau de cisaillement a sur la re´sistance au cisaillement de la perle de brasage dans les montages PGBA (Plastic Ball Grid Array). Il identifie une e´paisseur efficace pour l’analyse des e´le´ments finis bidimensionnels. En se servant de cette e´paisseur efficace comme facteur de graduation, il de´montre que le mode`le bidimensionnel peut convenir pour l’e´tude des proble`mes tridimensionnels. Le mode`le de calcul est valide´ par des donne´es expe´rimentales, en ce qui concerne les courbes de de´placement de la charge. Les re´sultats des essais, ainsi que ceux du modelage, indiquent que la vitesse du coulisseau de cisaillement a un effet important sur la re´sistance au cisaillement de la perle de brasage. De manie`re ge´ne´rale, une vitesse de coulisseau accrue permet d’augmenter la re´sistance de la perle de brasage. La caracte´risation de la re´sistance au cisaillement de la perle de brasage de´pend de`s lors du taux de chargement.

[7]

Abstracts & keywords

Korrelation zwischen Lo¨tpastenrheol ogie und der numerischen Simulation des Siebdruckprozesses

Numerische Modellierung des Scanned-Beam-Laserlo¨tens von Fine-Pitch-Bauelementen

R. Durairaj, G.J. Jackson, N.N. Ekere, G. Glinski und C. Bailey

P.M. Beckett, A.R. Fleming, J.M. Gilbert und D. G. Whitehead

Stichworte Lo¨tpasten, Rheologie, Schablonen

Stichworte Lo¨ten, Laser, elektronische Montage

Im Bereich der Lo¨ttechniken findet eine sta¨ndige Weiterentwicklung statt. Bedingt ist dies durch die Anforderungen der zunehmenden Miniaturisierung von elektronischen Bauteilen, vor allem an die Zusammensetzung der Lo¨tpasten, die zur Reflow-Lo¨tung von oberfla¨chenmontierten Bauelementen verwendet werden. Der Siebdruck ist bei der Massenproduktion von elektronischen Baugruppen weiterhin ein wichtiger Prozess fu¨r die Aufbringung der Lo¨tpaste auf die Leiterplatten, auch wenn die Erzielung einer einheitlichen Druckqualita¨t bei sehr kleinen Dimensionen problematisch ist. Zur Eliminierung dieser Defekte ist ein gutes Versta¨ndnis der am Druckverfahren beteiligten Prozesse erforderlich. Numerische Simulationen ko¨nnen dabei neben experimentellen Druckprobela¨ufen und Lo¨tpastencharakterisierungs-Studien eingesetzt werden und tragen weiter zum Versta¨ndnis des Prozesses bei. Die Merkmale und Fliebeigenschaften von Lo¨tpasten ha¨ngen in erster Linie von ihrer chemischen und physikalischen Zusammensetzung ab. Zuverla¨ssige Daten zu den Materialeigenschaften sind daher unerla¨sslich, wenn die numerische Simulation aussagekra¨ftige Ergebnisse liefern soll. Der Artikel beschreibt Lo¨tpastencharakterisierungs und numerische Simulations-Studien, die durch Zusammenarbeit der School of Aeronautical, Mechanical and Manufacturing Engineering der Universita¨t Salford und des Centre for Numerical Modelling and Process Analysis der Universita¨t Greenwich durchgefu¨hrt wurden. Das rheologische Profil der beiden verschiedenen Pastenformeln (bleihaltig bzw. bleifrei) fu¨r weniger als 100 Mikrometer große Flip-Chip-Bauelemente wird gepru¨ft und auf die numerische Simulation des Fließverhaltens der Pasten beim Siebdruck angewandt.

Die Laserlo¨tung ist fu¨r Elektronikhersteller ein sehr nu¨tzliches Verfahren und hat erfolgreich Eingang bei einer Reihe von industriellen Anwendungen gefunden. Der Laser sorgt fu¨r eine gut kontrollierbare lokale Erhitzung, a¨hnlich wie dies beim Lo¨ten von Hand der Fall ist, und zeichnet sich gegenu¨ber anderen Lo¨tmethoden durch klare Vorteile aus. Die Erhitzungsprozesse beim Laserlo¨ten sind jedoch komplex und es ist daher schwer abscha¨tzbar, wie gut oder schlecht eine bestimmte Lo¨tung ausfallen wird. Die numerische Modellierung hilft dabei, diese Fragen zu beantworten, um jedoch zuverla¨ssige Ergebnisse zu erhalten, ist ein gutes Versta¨ndnis des Modellierungsprozesses und der Annahmen, auf denen dieser beruht, erforderlich. Der Artikel erkla¨rt die Annahmen, die bei der Erarbeitung eines geeigneten Modells zugrunde gelegt wurden, und beschreibt die Anwendung eines kommerziellen Finite-ElementeModellierungsprogramms zur Untersuchung von Laserlo¨tprozessen, insbesondere von Scanned-BeamLo¨tvorga¨ngen mit einem oder mehreren Durchga¨ngen.

Modellierung des Lo¨tpasten-Reflows S.H. Mannan Stichworte Lo¨tpasten, numerische Stro¨mungsdynamik Lo¨tpastendruck und reflow sind seit langem etablierte Prozesse zur Herstellung von Lo¨tverbindungen bei elektronischen Baugruppen. Die Lo¨tpaste besteht aus einer dichten Suspension von Lo¨tpartikel in einem flu¨ssigen Medium (Vehikel), das beim Ru¨ckfluss als Oxidreduktionsmittel (Flussmittel) fungiert und die Metalloberfla¨chen von Oxiden befreit. Der Artikel berichtet u¨ber Versuche, die beim Lo¨tpastenru¨ckfluss auftretenden physikalischen und chemischen Prozesse durch numerische Stro¨mungsdynamik (Computatinal Fluid Dynamics – CFD) zu modellieren. Beschrieben werden achssymmetrische, zwei und dreidimensionale Modelle. Zudem wird eine Methode zur Reproduktion des oxidartigen Verhaltens bei diesen Modellen vorgestellt.

Soldering & Surface Mount Technology 14/1 [2002] German abstracts q MCB UP Limited [ISSN 0954-0911]

[8]

Vereinfachtes Modell des ReflowLo¨tprozesses David C. Whalley und Stuart M. Hyslop Stichworte Lo¨ten, Reflow Fru¨here Modelle, die sich mit der Temperaturentwicklung beim Reflow-Lo¨ten befassten, verwendeten normalerweise kommerziell erha¨ltliche, allgemeine FiniteDifferenz/Finite-Elemente-Modellierungsinstrumente, um detaillierte, dreidimensionale Darstellungen des Produkts und des Reflow-Ofens zu erzeugen. Derartige Modelle sagen sehr genau die Temperatur vorher, die ein bestimmtes Leiterplattendesign beim Reflow-Prozess erreicht, ihre Erstellung ist jedoch auch bei Verwendung moderner, leistungsstarker Computer komplex und die Analysezeiten sind lang. Der Artikel berichtet u¨ber die Entwicklung eines vereinfachten Prozessmodells, das eine weniger komplexe Darstellung des Produkts und Prozesses und einen einfachen, speziell fu¨r diese Anwendung entwickelten numerischen ”Solver“ entha¨lt und dessen Genauigkeit mit der von detaillierteren Modellen vergleichbar ist. Bei dem vereinfachten Modelle wird das Produkt in Elemente eingeteilt, die mit einem zweidimensionalen Raster aus Wa¨rmekonduktanz und thermisch wirksamer Masse dargestellt werden. Die Werte dieser Konduktanzen und Massen werden basierend auf den gemittelten Eigenschaften des Leiterplattenmaterials und der hiermit verbundenen Komponenten im Bereich der einzelnen Elemente berechnet. Die Darstellung des Reflow-Ofens beruht auf der Messung der Temperatur und der Wa¨rmekonvektion an jedem Punkt entlang der La¨nge des Ofens, so dass keine genaue Messung der Ofengeometrie und der Luftstro¨mungsgeschwindigkeit erforderlich ist. Die Kombination dieser beiden Vereinfachungstechniken ermo¨glicht eine Verku¨rzung der Analysezeit bei relativ einfachen Leiterplatten, wobei die Zeitersparnis zwischen einer Stunde bei einer hochleistungsfa¨higen UnixWorkstation und einer Sekunde bei einem Pentium-PC mit Microsoft Windows liegen kann.

Abstracts & keywords Soldering & Surface Mount Technology 14/1 [2002] 8–9

CFD-Modellierung des Flussfeldes in einem Reflow-Ofen

Optimierungsmodell fu¨r die Zuverla¨ssigkeit von Flip-Chip-Lo¨tverbindungen

Hao Yu und Jorma Kivilathti

S. Stoyanov, C. Bailey und M. Cross

Stichworte Bleifreies Lo¨ten, numerische Stro¨mungsdynamik

Stichworte Optimierung, Flip-chip, Underfill, Zuverla¨ssigkeit

Aufgrund der zunehmenden Forderung nach Verwendung von bleifreien Materialien und der Vielfa¨ltigkeit der Komponenten und Metalloberfla¨chen auf dicht bestu¨ckten Leiterplatten ist ein besseres Versta¨ndnis und eine genauere Kontrolle des Reflow-Prozesses erforderlich, um bei der Herstellung von SMT-Baugruppen eine akzeptable Ausbeute und Zuverla¨ssigkeit zu erzielen. Eine genaue Kontrolle der Temperaturverteilung in den Komponenten und Leiterplatten wa¨hrend des Reflow-Prozesses ist dabei vor allem bei bleifreien Baugruppen eine der wichtigsten Forderungen. Der Artikel beschreibt einen Rahmen fu¨r die Reflow-Modellierung und stellt ein auf Ofenebene operierendes Modell des Steady-State-Flussfeldes in einem Reflow-Ofen vor, das bei der anschließenden Transientanalyse und klein angelegten Analyse beno¨tigt wird. Das Modell wird unter Zuhilfenahme der fortschrittlichen numerischen Stro¨mungsdynamik (CFD)Technik mit einer handelsu¨blichen Software konstruiert. Die numerischen Ergebnisse werden diskutiert und mit den gemessenen Daten verglichen.

Der Artikel beschreibt und demonstriert, wie die integrierte Optimierungs/Zuverla¨ssigkeits-Modellierung dazu verwendet werden kann, um das Verhalten von Lo¨tverbindungen in elektronischen Baugruppen vorherzusagen. Die integrierte Modellierung wird dazu verwendet, um effizient und schnell die geeignetsten Designparameter fu¨r das Verhalten von Lo¨tverbindungen bei Temperaturschwankungen zu identifizieren, und wird anhand von Flip-Chip-Komponenten mit ”No-Flow“Underfills demonstriert. Um den ”Optimierungs/Zuverla¨ssigkeits“-Ansatz zu implementieren, wird die Finite-Elemente-Simulation PHYSICA mit Optimierungs und Statistikinstrumenten gekoppelt. Der hieraus resultierende Rahmen ermo¨glicht die vollkommen automatisierte, systematische Anwendung von Designoptimierungsverfahren.

Analyse der Lo¨tperlen-Scherversuchbedingungen mit einem einfachen Berechnungsmodell S.W. Ricky Lee und Xingjia Huang Stichworte Plastic Ball Grid Array, Lo¨tverbindung, Finite-Elemente-Analyse, Zuverla¨ssigkeit Der Artikel stellt ein einfaches Rechenmodell zur Analyse der Lo¨tperlen-Scherversuchbedingungen vor. Zweidimensionale (2D) und dreidimensionale (3D) FiniteElemente-Modelle werden verwendet, um den Effekt der Schersto¨ßelgeschwindigkeit auf die Lo¨tperlenScherfestigkeit von PBGA (Plastic Ball Grid Array)Baugruppen zu untersuchen. Es wird eine effektive Dicke fu¨r die 2D-Finite-Elemente-Analyse identifiziert. Durch Verwendung der effektiven Dicke als Skalierungsfaktor wird gezeigt, dass das 2D-Modell zur Untersuchung von 3DProblemen eingesetzt werden kann. Das Berechnungsmodell wird durch experimentelle Daten aus Ladungsverdra¨ngungskurven validiert. Die Ergebnisse der Tests und Modellierung zeigen, dass die Schersto¨ßelgeschwindigkeit einen wesentlichen Effekt auf die Lo¨tperlen-Scherfestigkeit hat. Eine ho¨here Schersto¨ßelgeschwindigkeit kann allgemein zu einer ho¨heren Lo¨tperlen-Scherfestigkeit fu¨hren. Die Charakterisierung der Lo¨tperlen-Scherfestigkeit ist daher von der Ladungsrate abha¨ngig.

[9]

Editorial

This special issue of SSMT brings together seven papers demonstrating the latest achievements in the applications of computational modelling technology to soldering processes and solder joint reliability.

both during the soldering process and subsequently during the lifetime of the solder interconnects.

How are models currently used? Why use computational models? The performance of soldering materials during product assembly is governed by complex interacting physical phenomena. Although experimentation can be used to provide an insight into the mechanisms taking place it is often impossible to use such techniques for all situations. The development and use of computational models provides a way of understanding solder material behaviour that would be impossible or far too expensive to achieve through experimentation alone. Computational models also provide the ability to reduce the overall cost of designing and manufacturing new products. This is due in part to increased accuracy in the mathematical models, larger and faster computer systems, and better graphical interfaces allowing easy use of modelling technology by engineers. Using these technologies to test new ideas and analyse current processes helps take the guesswork out of industrial process design and offers very attractive benefits to industry – in terms of financial cost savings. The use of such software tools in the context of soldering and assembly is the theme behind this special issue of SSMT.

What is computational modelling?

Soldering and Surface Mount Technology 14/1 [2002] 10 q MCB UP Limited [ISSN 0954-0911]

[ 10 ]

At the heart of any computational modelling analysis are the mathematical equations that relate, for example, the heat transfer properties of a material to its surroundings. Ever since ancient times people have used mathematics to help understand the world around them. For example, the mathematical masters of Greece provided the world with Euclidean geometry, which provided the foundations for navigation and exploration. The 17th century saw a major leap in mathematical representation of physical phenomena with the development of calculus by Newton. This provided the basis for other scientists to develop the classical equations of mathematical physics throughout the 18th and 19th centuries, such as: Heat Transfer (Fourier), Fluid Flow (Bernoulli, Navier, Stokes), Solid Mechanics (Navier, Hooke), and Electromagnetism (Faraday and Maxwell). The arrival of commercially available digital computers in the late 1950’s and the introduction of high level computing languages such as Algol and Fortran marked the true birth of computational modelling as we know it today. With the introduction of computers it became possible to solve many simultaneous equations whose number was only limited by the memory available. In the early days the number of simultaneous equations (representing, for example, the unknown temperatures across a printed circuit board) was in the 100’s. If the growth in computing power had remained linear over the last forty years, then the scientific community will still be very limited in what it could accomplish today with computers. Fortunately advances in computer chip architecture have grown at such a rate that computing power has doubled every 18 months. In fact, some estimates state that our problem solving capability has increased by a factor of a million during the last thirty years. Solving the governing equations of physics using numerical techniques embedded in software provides the opportunity to investigate, in detail, phenomena taking place

This issue of SSMT demonstrates how computational modelling tools are currently being used to investigate different aspects of the soldering process from solder paste print to solder joint reliability. The first paper by Durairaj and co-workers demonstrates models that predict the flow of solder paste material during the printing process. Given appropriate rheology data such models, based on computational fluid dynamics, can be used to predict the velocity, pressure and shear rates in solder as it rolls across a stencil. Of interest here is the different performance predicted for both tin-lead and a lead-free solder paste and the pressures along the stencil. Solder paste coalescence is the topic discussed in the second paper by Mannan. This illustrates the use of computational fluid dynamics to predict the coalescence of solder paste particles during the reflow process. These very challenging simulations, which also have the ability to predict the breakdown of oxide layers, are beginning to provide an insight into the manner in which solder particles break down and coalesce to form the required interconnect. Laser soldering is the theme of the third paper by Beckett and co-workers. The models used in this analysis accurately represent the energy from the laser and its effect on soldering temperatures. Comparisons are provided with thermocouple readings showing the close integration between experimentation and modelling. The fourth paper by Whalley and Hyslop details a modelling approach to predict temperatures across a printed circuit board as it travels through a reflow oven. Based on process measurements these simple models provide very accurate results and can be used quickly by analysts to ensure that specific oven conditions will result in appropriate board temperatures. The reflow process is again the topic in the fifth paper by Yu and Kivilathi. Very detailed computational fluid dynamics models are used to predict the temperature through the oven including the effects of airflow. Again good comparisons are provided between the model and measured data. In the sixth paper by Lee and Huang a model is presented for the shear test of a solder bump based on the JEDEC standard, JESD22-B117. Here the stress in the solder is predicted for different shear rates where the models are identifying optimal shear speeds. The final paper by Stoyanov and co-workers uses models to identify optimal conditions for solder joint reliability during thermal cycling. The modelling methodology used is based on integrating thermo-mechanical finite element analysis with optimisation tools, such as design of experiments. The papers in this issue clearly illustrate the breadth in which computational modelling analysis tools are being used by researchers investigating the soldering process. We should note that such tools do not replace experimentation. When possible these tools are used in harmony with experimentation and skilled judgement to provide a sound basis to scientific investigation. Dr Chris Bailey Guest Editor Computing and Mathematical Sciences University of Greenwich London SE10 9LS UK [email protected]

Correlation of solder paste rheology with computational simulations of the stencil printing process

Keywords Solder pastes, Rheology, Stencils Abstract Soldering technologies continue to evolve to meet the demands of the continuous miniaturisation of electronic products, particularly in the area of solder paste formulations used in the reflow soldering of surface mount devices. Stencil printing continues to be a leading process used for the deposition of solder paste onto printed circuit boards (PCBs) in the volume production of electronic assemblies, despite problems in achieving a consistent print quality at an ultra-fine pitch. In order to eliminate these defects a good understanding of the processes involved in printing is important. Computational simulations may complement experimental print trials and paste characterisation studies, and provide an extra dimension to the understanding of the process. The characteristics and flow properties of solder pastes depend primarily on their chemical and physical composition and good material property data is essential for meaningful results to be obtained by computational simulation.This paper describes paste characterisation and computational simulation studies that have been undertaken through the collaboration of the School of Aeronautical, Mechanical and Manufacturing Engineering at Salford University and the Centre for Numerical Modelling and Process Analysis at the University of Greenwich. The rheological profile of two different paste formulations (lead and lead-free) for sub 100 micron flip-chip devices are tested and applied to computational simulations of their flow behaviour during the printing process.

Soldering & Surface Mount Technology 14/1 [2002] 11–17 q MCB UP Limited [ISSN 0954-0911] DOI 10.1108/09540910210416422

R. Durairaj Electronics Manufacturing Engineering Research Group, School of Aeronautical, Mechanical and Manufacturing Engineering, University of Salford, Salford, Manchester, UK G.J. Jackson Electronics Manufacturing Engineering Research Group, School of Aeronautical, Mechanical and Manufacturing Engineering, University of Salford, Salford, Manchester, UK N.N. Ekere Electronics Manufacturing Engineering Research Group, School of Aeronautical, Mechanical and Manufacturing Engineering, University of Salford, Salford, Manchester, UK G. Glinski Centre for Numerical Modelling and Process Analysis, University of Greenwich, London, UK C. Bailey Centre for Numerical Modelling and Process Analysis, University of Greenwich, London, UK

Introduction Over the last few years, the electronics manufacturing industry has been under enormous pressure to meet the requirements for miniaturised products. Current technology advances are largely driven by the market for consumer products such as mobile phones, where there is an increasing demand for size and weight reductions whilst adding more features. The pressure for increased component density at ultra-fine pitch is also coupled with further pressures to reduce costs and to create more environmentally friendly products and assembly processes. Solder paste deposition is a critical stage in the mass production assembly of surface mount components. Solder pastes typically consist of solder alloy powder, flux, viscosity control agents and a solvent system. It is essential that a sufficient volume of solder paste be deposited consistently onto all joint pads without any failures to form good electrical connections. Achieving consistent solder paste deposits presents a considerable challenge in ultra-fine pitch applications such as sub 100-micron flip-chip assemblies. As PCB pad and component lead sizes continue to reduce, most paste suppliers have opted for smaller solder powder particle sizes and new rheology modifier formulations to meet ultra-fine pitch and flip-chip assembly requirements. One of the most widely used techniques for the deposition of solder paste is by stencil printing. Successful printing depends on several variables, such as the composition of the solder paste (e.g. particle size), stencil design and manufacture (e.g. smoothness of the aperture walls) and the print equipment (e.g. alignment capability between the stencil and substrate). The rheology and printing performance of solder pastes can be controlled by varying the size distribution and shape of the solder powder particles and the other particulate constituents. The correlation of paste characterisation studies with the results of experimental print quality trials offers a valuable means of identifying optimum material properties and print

machine settings. Analysis of the flow and properties within the bulk of paste during printing are extremely difficult if not impossible to measure experimentally. However, computational simulations can offer a visual insight to the internal flow behaviour, adding an extra dimension to the understanding of the factors influencing print quality.

Stencil printing In this process, solder paste rolls in front of a squeegee blade, filling apertures in the stencil some distance ahead of the squeegee (Figure 1). A 60 degree blade angle relative to the stencil surface is commonly used in most printing applications, though an angle of 45 degrees is also used in some instances. The apertures in a stencil are fully opened and do not obstruct the flow of solder paste. The paste flows through the apertures, which are aligned precisely with the printed circuit board (PCB) joint pads. The squeegee shears off the paste above the apertures as it move overhead. Once the print stroke is completed, the PCB is separated mechanically from the stencil leaving freestanding deposits on the joint pads.

Figure 1 Schematic representation of the stencil printing process

The current issue and full text archive of this journal is available at http: //www.emeraldinsight.com /0954-0911.htm [ 11 ]

R. Durairaj, G.J. Jackson, N.N. Ekera, G. Glinski and C. Bailey Correlation of solder paste rheology with computational simulations of the stencil printing process Soldering & Surface Mount Technology 14/1 [2002] 11–17

[ 12 ]

The stencil printing process accounts for about 60% of assembly defects, which include bridging between adjacent pads, slumping, poor shape definition or an insufficient deposited volume (skipping). These defects lead to either short or open connections, the presence of which leads to the failure of the component. Their reduction leads to lower rework costs and higher yield, therefore resulting in significant economical benefits for manufacturing. Though seemingly simple, the underlying process at a microscopic scale is very complex. During printing, the solder paste passes through a number of procedures including: pre-print paste treatment, paste roll in front of the squeegee, aperture filling, aperture emptying and finally free standing paste deposits prior to component placement. Several forces can be identified that are critical to the paste transfer process (Hwang, 1989). Hydrodynamic pressure is generated in the paste roll during the print stroke that helps inject the paste into the apertures. The reaction of the squeegee blade to this hydrodynamic pressure (the force exerted on the paste by the squeegee) can be separated into a downward vertical component and a horizontal component in the direction of the print stroke. The total downward vertical force exerted by the printing machine on the squeegee blade is usually an adjustable setting on most stencil printing machines. This force is opposed partly by the hydrodynamic pressure generated in the paste roll and partly by the reaction of the stencil surface to the force exerted by the blade tip. Both these reactions can cause deformation of the blade, the extent of which is determined by its material properties, e.g. metal or plastic. As the substrate moves away from the stencil, the paste experiences frictional forces at the aperture walls. In order to achieve good paste transfer, the adhesive force between the joint pad and solder paste must be greater than these frictional forces, which are dependent on the roughness of the aperture walls. The influence of gravitational force can be considered to be negligible in all these processes.

and performance of solder paste. The print quality, wetting and slumping properties of the solder paste are influenced by factors such as the shape, size, distribution and surface finish of the alloy particles. Spherically shaped particles with smooth surfaces are preferred as they minimise stencil clogging. The composition and formulation of the flux/vehicle system is also critical in defining the flow characteristics and printing performance of solder paste. The formulation of flux vehicle systems is complex and requires considerable experience and expertise. Consequently, these formulations are confidential and closely guarded by the manufacturers. The flux vehicle and solvent systems consists of rosin/resin, solvents, activators, thickeners, rheological agents and other additives. A flux vehicle must be able to perform a number of tasks. The activators provide the flux with the ability to clean contaminants such as oxides and organic layers from the surfaces to be soldered. The quantity of activators added depends on the required strength of the flux. The flux also coats the surfaces to prevent re-oxidation when close to reflow soldering temperatures and assists in heat transfer during reflow. The solvents are non-volatile and used to dissolve the flux and other active constituents in the solder paste. Gycols and esterified colophanes are most commonly used (Guillaume et al., 1989). Solvents are the largest constituents of the flux system, at a volume fraction of approximately 98 per cent. They help to prevent solder paste dry-out and also promote the tackiness of solder paste required to hold components in place prior to reflow. Rheological and thixotropic additives are added to the flux system to control the viscosity and flow properties of the solder paste during printing. Additives are also used to reduce the rate of chemical reactions at room temperature, thereby keeping the rheological behaviour of solder paste constant.

Solder paste composition

Computational modelling of the stencil printing process

Solder pastes are essentially composed of metal alloy powder particles dispersed in a thickened flux vehicle and solvent system. The rheological properties depend primarily on the chemical and physical composition, with all pastes exhibiting non-Newtonian, shear thickening and thixotropic behaviour. Solder paste consists of three main constituents (Nguty and Ekere, 1999). They include solder alloy powder; a flux system to promote metallic bonding by providing good wetting conditions and clean surfaces; and a vehicle carrier system binding the solder powder and the flux system together, and providing the desirable rheological properties. Solder paste characteristics can be analysed in two ways (Hwang, 1989). The paste can be considered to be a dense particulate suspension with a solder/vehicle volume fraction of approximately 50% with attributes such as solder particle size distribution, metal content, inter-particle forces and possible particle-flux interactions. Alternatively, the microscopic structure and individual constituents may be ignored and the paste can be considered as a homogenous and kinetically stable mixture, categorised by bulk macroscopic properties such as viscosity, and density. A standard solder paste will have a metal content of 88– 91 per cent by weight. Solder paste of powder particle size Type 3 (20–45 mm in diameter) is the industry standard for fine pitch components. Tin/lead (Sn/Pb) is the typical binary alloy used for the solder particles. Introduction of legislation to ban the use of lead in electronic products has put tremendous pressure to find a replacement for tin/lead solder. Tin/silver/copper metal alloys are considered to be a suitable replacement. Some electronic manufacturers are also considering metal alloys such as tin/silver/bismuth, tin/zinc/bismuth, tin/silver and tin/copper. A change in the alloy used has a strong impact on the rheology, printability

The features of the printing process that may be modelled include the macroscopic flow of the bulk paste roll across the stencil surface (Glinski et al., 2001) as well as the entry of the paste into the apertures and subsequent release during stencil separation at a microscopic scale. The microscopic structure and complex dynamic transport mechanisms of the paste particles result in observable bulk material properties at the macroscopic scale. The length scale of the paste bulk is much larger than that of the underlying microscopic structure so the individual particle properties and dynamics may be ignored and the paste may be modelled as a homogeneous continuum. The macroscopic properties may be measured and used to represent bulk density and shear rate/viscosity relationships. There is a very large difference in the spatial scale of the roll in comparison to the geometry of the stencil apertures. The small local disturbances to the flow field due to aperture filling at a microscopic scale may be considered to have a negligible effect on the macroscopic flow characteristics of the whole paste bulk. For the macroscopic study, the stencil is therefore considered to be a smooth flat surface with no apertures. Using the Computational Fluid Dynamics (CFD) method, the geometry of the paste roll ahead of the squeegee blade is discretised into a two dimensional grid shown in Figure 2. This two-dimensional slice represents a vertical plane through the paste roll at the midpoint of the squeegee blade. The computational modelling method involves the discretisation of the relevant partial differential equations representing fluid flow (Navier–Stokes Equations) over the defined grid. The material properties are considered to be uniform within the volume of each individual grid cell but may vary in a piece-wise manner from cell to cell. The shear

R. Durairaj, G.J. Jackson, N.N. Ekera, G. Glinski and C. Bailey Correlation of solder paste rheology with computational simulations of the stencil printing process Soldering & Surface Mount Technology 14/1 [2002] 11–17

rate/viscosity relationship of the paste may be represented by a suitable model, e.g. the Cross model or a power-law relationship. The models used for the simulations presented here are based on a Cross model relationship obtained by paste characterisation studies carried out at Salford University. The solution procedure starts with initially guessed pressure and velocity fields and proceeds thereafter in an iterative manner. The viscosity is determined and updated from the shear rate within each cell, calculated from the gradients of the velocity field results after each iteration of the solution procedure. The iterative process continues until the emerging solution converges, i.e. the results within each grid cell do not change significantly with further iterations and may be considered to be a sufficiently accurate approximation to the solution. The continuum approach of this modelling method presents difficulties when applied to the simulation of the solder paste entry and release from the stencil aperture. The reason for this is that the relative size of the solder spheres in relation to that of the stencil apertures is large. In the macroscopic model of the bulk paste roll, a single grid cell represents a region of paste containing a large number of solder spheres so that a homogeneous continuum can be assumed. However, at the microscopic scale, the discretisation of the aperture geometry using a continuum method would require cells much smaller than the solder particles if the flux flow is to be modelled with sufficient accuracy. Either frequent remeshing or other measures to overcome the deterioration of the grid structure is required, which would enable the grid to follow the complex motions of the solder particles moving around each other. Although, current research projects have overcome some of these problems, the methods tend to be very computationally expensive. New computational methods have emerged over the last decade that avoid the constraints of the Eulerian grids on established CFD methods. They are known as mesoscopic methods and are based on the collisions of fluid particles, similar to the concepts of molecular dynamics. The particles and their collisions are either constrained to occur on a fixed regular lattice as in the Lattice–Boltzmann method (Chen et al., 1992) or may be Lagrangian, i.e. unconstrained in space, as in dissipative particle dynamics (Hoogerbrugge and Koelman, 1993). The models are a course-grained representation of the motion of fluid molecules. However, each model particle represents a large number of the real molecules with collision properties set so that the transport properties of the real underlying fluid are mimicked at the macroscopic scale. The advantage of these methods is that they may be adapted to contain large suspended particles such as solder spheres in a flux system (Aidun et al., 1998).

Figure 2 Discretisation of the geometry into a computational grid

The motions of the suspended particles are calculated from the forces due to the collisions of the fluid particles. The material properties of the flux, therefore, need to be measured and defined without the solder particle content. The simulations may be applied to arbitrary complex geometries and offer a more computationally efficient method of establishing the behaviour of suspension flows than traditional CFD methods. The specification of the initial random arrangement of solder particles at the start of a simulation is quite a complex task. The arrangement needs to represent a realistic packing arrangement at the true volume fraction. Special algorithms developed at Salford University are used to achieve this. These have been used to study the packing arrangements of different solder particles size distributions inside the stencil aperture but are also ideal for the specification of initial a conditions of a microscopic simulation. The boundary conditions at the microscopic study boundary represent a small sub-region of the paste bulk close to the aperture and need to be specified from macroscopic results. These are based on the pressure, velocity and shear rate distributions obtained at the base of the paste roll at the stencil surface. Mesoscopic methods used for aperture filling and emptying simulations are however an approximation and must be treated with care. The emphasis is therefore on insight into the complexities of paste behaviour rather than a true representation of the particle motion.

Rheological testing A Reologica StressTech controlled stress and strain rheometer was used to investigate solder paste rheological properties. A sample thickness of 0.5 mm sandwiched between two smooth flat plates with a 40 mm diameter was used as shown in Figure 3. The deformation of the sample can be measured in either a controlled stress or controlled strain mode. The solder paste used had a metal content of 89% (by weight) and particle size and distribution of 5–15 mm. The solder powder alloys used where 63Sn/37Pb (Sample P1) and 95.5Sn/3.8Ag/0.7Cu (Sample P2), suspended in a noclean flux medium. Gentle stirring of the samples in their original jars for about 30 s before loading onto the rheometer is a necessary preparatory procedure. Once loaded, no preshear was applied. Loading of the sample is made difficult by the highly viscous properties of solder pastes. To facilitate sample loading, a mini-stencil (1 mm thick) was used to deposit the solder paste onto the bottom plate of the rheometer. The top plate was then lowered to the desired gap (0.5 mm), squeezing the extra paste out from between the plates. Excess paste was trimmed from the edges of the plates to minimise edge effects. Figure 3 Solder paste sandwiched between parallel plates

[ 13 ]

R. Durairaj, G.J. Jackson, N.N. Ekera, G. Glinski and C. Bailey Correlation of solder paste rheology with computational simulations of the stencil printing process

Figure 4 Shear rate/Viscosity curve for paste sample P1

Soldering & Surface Mount Technology 14/1 [2002] 11–17

Steady shear rheometry involves the measurement of viscosity at different shear conditions. The sensitivity or dependence of a sample to changing shear rates is evaluated (shear thinning, thixotropy, hysteresis, etc). Shear thinning properties (the reduction of viscosity of a material when sheared) are essential to assist the entry of paste into the stencil apertures. The degree of shear thinning is measured by comparing the viscosity at different shear rates. During the printing process, solder paste experiences a range of shear rates in the range of 0.01–1000 s 2 1 and more

Figure 5 Shear rate/Viscosity curve for paste sample P2

[ 14 ]

(Trease and Diets, 1972). Though the shear rates investigated here are low, they can be used to predict the flow properties during stencil aperture filling and emptying processes. The temperature was set at 258C and the relative humidity was also monitored during the exercise. A shear rate sweep procedure was conducted to measure viscosity from low-to-high shear rates of 0.1 – 20 s 2 1. The measured viscosity curves for samples P1 and P2 are plotted using logarithmic scales in Figures 4 and 5, respectively.

R. Durairaj, G.J. Jackson, N.N. Ekera, G. Glinski and C. Bailey Correlation of solder paste rheology with computational simulations of the stencil printing process Soldering & Surface Mount Technology 14/1 [2002] 11–17

The Cross model (Ferguson and Kemblowski, 1991), shown in equation 1 was applied to the measured viscosity/shear rate data.

h 2 h1 1 ¼ h0 2 h1 1 þ K g_m

ð1Þ

where h0 and h1 are the viscosity at zero and infinite shear rates and K and m are constants. The parameters obtained for the two paste samples are shown in Table I.

Computational modelling results A macroscopic analysis of the bulk paste roll was performed using the PHYSICA (PHYSICA Multi-Physics Computational Modelling Software) CFD software. The code solves the governing partial differential equations for fluid motion including the non-Newtonian constitutive model for viscosity/shear rate given by the Cross model above. The analysis was performed for the two paste samples P1 and P2 using data obtained in the rheological tests described above. Figure 6 shows a plot of the velocity vector field for the lead free sample P2. Pressure, velocity, shear rate and viscosity distributions were determined throughout the geometry of the paste roll for a blade angle of 60 degrees and velocity of 1, 2, 3 and 4 cm/s. The distributions obtained along the base of the paste roll are of particular interest as the aperture filling process depends on the paste behaviour and material properties encountered in this region adjacent to the stencil surface. Plots of the pressure, shear rate and viscosity distributions at the stencil surface up to 5 mm from the blade tip are presented in Figures 7 –9, respectively. The pressure and shear rate distributions are plotted using logarithmic scales. Particularly, large pressure gradients are observed in the region closest to the blade tip for both the lead (P1) and leadfree (P2) samples. The resolution of the grid must be sufficiently fine here to adequately represent these rapid

Table I Cross model parameters for paste samples P1 and P2 Sample P1 P2

h0 (K Pa s)

h1 (Pa s)

m

K (s)

21760 9589

11.66 45.35

0.87 0.95

152 10

changes. As would be expected the pressure generated increases with blade velocity. There is a considerable difference in the pressure generated between the two paste samples. Conversely, the shear rate distributions are shown to be remarkably similar (Figure 8) and increase rapidly towards the blade tip. The viscosity distributions in Figure 9 show a corresponding decrease in viscosity towards the blade tip due to the increased shear rate. The viscosity of the lead-free sample (P2) is considerably higher than that of the lead sample (P1) at the similar shear rates, which accounts for the difference in the pressures generated at the same blade velocity. The computational simulation results provide parameters based on the internal flow characteristics of the paste roll that can be correlated with machine settings and print quality. An insight is offered to the role of these parameters in the aperture filling process. For example, it could be argued that the lower pressure generated by the conventional lead solder paste (P1) is compensated by the lower viscosity which facilitates entry into the aperture. Conversely, the higher pressure generated in the lead-free paste (P2) compensates for the higher viscosity. The distributions also provide boundary conditions for the subgeometries of microscopic aperture scale studies currently under investigation.

Conclusions Solder paste characterisation and rheological studies coupled with computational simulations offer an insight into the mechanisms of the stencil printing process that would otherwise not be available. The comparisons between a lead and lead-free paste sample presented here are only one example of the vast characterisation studies possible. For example, the same procedures can be applied to the comparison of samples with varying rheological properties of the flux system whilst maintaining the sample particle material and size distribution or vice–versa. Further rheological testing and development of the computational models is required however to account for microscopic phenomena that may have a significant influence on the paste behaviour at a macroscopic scale. This includes the phenomenon of wall slip where flux rich regions adjacent to a solid surface form a lubricating layer, reducing or eliminating the shear rate of the solid particles. The application of these results at a macroscopic scale to the boundary conditions of aperture filling and emptying simulations at a microscopic scale will further enhance the

Figure 6 Velocity vector plots for the paste roll at 2 cm/s

[ 15 ]

R. Durairaj, G.J. Jackson, N.N. Ekera, G. Glinski and C. Bailey Correlation of solder paste rheology with computational simulations of the stencil printing process

Figure 7 Pressure distribution along the stencil surface

Soldering & Surface Mount Technology 14/1 [2002] 11–17

Figure 8 Shear rate distribution along the stencil surface

Figure 9 Viscosity distribution along the stencil surface

[ 16 ]

R. Durairaj, G.J. Jackson, N.N. Ekera, G. Glinski and C. Bailey Correlation of solder paste rheology with computational simulations of the stencil printing process Soldering & Surface Mount Technology 14/1 [2002] 11–17

understanding of the critical parameters and conditions that will be required to maintain stencil printing as a leading solder deposition process for the continued miniaturisation of electronic components.

References Aidun, C.K., Lu, Y. and Jiang Ding, E. (1998), “Direct analysis of particulate suspensions with inertia using the discrete Boltzmann equation”, J. Fluid Mech., Vol. 373, pp. 287-311. Chen, H., Chen, S. and Matthaeus, W.H. (1992), “Recovery of the Navier –Stokes equations using a lattice – gas Boltzmann method”, Physical Review A, Vol. 45 No. 8, pp. R5339-R5341. Ferguson, J. and Kemblowski, Z. (1991), “Applied Fluid Rheology”, Elsevier Applied Science, p. 206. Glinski, G.P., Bailey, C. and Pericleous, K.A. (2001), “A nonNewtonian CFD study of the stencil printing process”,

Proceedings of the Institution of Mechanical Engineers, Vol. 215 No. C4, pp. 437-46. Guillaume, B., Guinet, J. and Hubert, J. C. (1989), “Screen printable solder pastes for surface mounting technology”, 7th European Conference Proceedings, Hamburg, Hybrid Microelectronics, ISHM 1989, Session 3. Hoogerbrugge, P.J. and Koelman, J.M.V.A. (1993), “Dynamic simulations of hard sphere suspensions under steady shear”, Europhys. Lett., Vol. 21 No. 3, pp. 363-68. Hwang, J.S. (1989), Solder Paste in Electronics Packaging, Van Nostrand Reinhold, New York, NY. Nguty, T. and Ekere, N.N. (1999), “Correlating solder paste composition with stencil printing performance”, IEEE/CPMT Int. Electr. Manufact. Technol. Symp., pp. 304-12. Patankar, S.V. (1980), “Numerical heat transfer and fluid flow”, Hemisphere Publishing Corp. PHYSICA Multi-Physics Computational Modelling Software, available at http://physica.ac.uk Trease, R.E. and Diets, R.L. (1972), “Rheology of pastes in thickfilm printing”, Solid State Technology, pp. 39-43.

[ 17 ]

Solder paste reflow modeling S.H. Mannan King’s College London, London, UK

Keywords Solder paste, Computational fluid dynamics Abstract Solder paste printing and reflow are well established processes for producing solder joints in electronic assemblies. Solder paste consists of a dense suspension of solder particles in a liquid medium (vehicle) that acts as an oxide reducing agent (flux) during reflow, cleaning the metal surfaces of oxides. This paper reports on attempts to model the physical and chemical processes occurring during solder paste reflow using computational fluid dynamics (CRD). Axisymmetric, 2 dimensional and 3-dimensional models are described, and a method of reproducing oxide-like behaviour in these models in introduced.

This work was conducted under UK EPSRC grant AF/990541. Particular thanks go to Jim Watts at Nortel Networks, Harlow for supplying the 0.5 mm diameter solder spheres.

Soldering & Surface Mount Technology 14/1 [2002] 18–23 q MCB UP Limited [ISSN 0954-0911] DOI 10.1108/09540910210416431

[ 18 ]

Introduction Solder paste printing and reflow are standard processes used during surface mount assembly. Solder paste reflow involves the coalescence of solder spheres which are typically 10–40 microns or smaller for fine pitch applications, into a single solder mass which wets the surfaces to be soldered and then solidifies to form a soldered joint. Before the solder spheres can merge and wet solderable surfaces, oxide layers on the solder balls and the solderable surfaces (i.e. pads on the die and substrate) must be removed. If the oxide layers on the surfaces of the solder spheres are not completely removed then these spheres will only partially merge, or may even not merge at all. Due to the continuous trend towards miniaturisation in electronics packaging, the solder joints for flip chip devices, for example, have become smaller in volume, leading to smaller particle sized pastes for ease of deposition. This, together with the smaller number of solder particles required to make a joint, can lead to incomplete reflow of the solder particles and solder balling. The new generation of lead free solders can also exhibit reduced wetting characteristics, leading to defects during soldering. Hence there is an urgent need to a construct models which will help to understand the causes of soldering defects and how to eliminate them. Such models can also be used to predict the effects of changing process conditions or chemical activity in the paste. Computational fluid dynamics (CFD) has been applied successfully to many engineering problems over the last 30 years. The standard CFD technique involves splitting the domain into a number of discrete volume elements, and representing the governing partial differential equations for fluid flow as difference equations involving quantities on the nodes or faces of these discrete elements. This approach works well when the fluid is confined between rigid boundaries, and has more recently been extended to deal with situations when free fluid surfaces are present, e.g. fluid filling a void or two or more immiscible fluids in contact. The free-surface interface is tracked by the presence of an abstract scalar quantity that is advected with one of the fluids. Figure 1 shows how such a scheme might operate, with the quantity, f, having a value between 0 and 1, where the value 1 indicates a cell completely filled with fluid 2, and f ¼ 0 represents a cell completely filled with fluid 1. The advantage of this scheme is that the mesh used to split the fluid domain into individual elements remains fixed while the fluid can flow freely over the entire domain. The particular fluid tracking scheme implemented in the current work is known as Volume of Fluid, (Hirt et al., 1981), and a commercial CFD software package, FIDAP, is used to run the simulations.

The current issue and full text archive of this journal is available at http: //www.emeraldinsight.com /0954-0911.htm

In order to accurately model the reflow behaviour of solder paste, it is important to model chemical reactions as well as the pressure and surface tension driven solder flows. CFD codes model chemical species concentrations by scalar quantities which are advected with the fluid motion in a similar manner to f, above. However, the equations governing chemical species behaviour include both diffusion and chemical reaction terms. For example, in the case of solder paste, we have a number of reactions at the metal surfaces. These include oxidation and reduction where metal oxide reacts with the flux. Diffusion of the oxygen and active flux reagents also take place through the vehicle, as well as a variety of chemical reactions within the flux itself. The reaction rate constants can be made dependent on temperature and on the concentration of other chemical species, but the detailed data for these reactions is mostly unknown at present, and also highly dependent on the exact chemistry used for each individual paste. This information is confidential and unlikely to be released by any paste manufacturer. However, as far as the flow of solder is concerned, the only significant reactions are those that control the oxidation and reduction at the metal surfaces. Therefore, in order to produce a model of solder paste reflow, only the overall effect of the flux on oxide layer thickness needs to be examined. This is shown schematically in Figure 2. Improved accuracy would be obtained by modelling all the chemical processes occurring as the rates of oxidation and reduction are governed by more factors than simply the concentrations of flux and oxygen, and local temperature. Ideally the oxide surfaces, which are of the order of 1–2 nm thick on solder balls (Kuhmann et al., 1988) would be modelled separately from the bulk solder reflow process because of the disparity in length scales. However even with a limited chemical model, and idealized solder surfaces, it is possible to model defects such as solder balling and incomplete solder reflow.

2-d models of solder paste coalescence without oxides In many CFD problems, it is possible to model the problem in 2-dimensions without losing a significant amount of accuracy. As an example, consider the flow of solder paste in the paste roll in front of the squeegee during stencil printing of solder paste. This situation can be accurately modelled by considering a cross section of the geometry as shown in Figure 3. The advantage of 2-d representations are much quicker solution times, and solutions in cases where 3-d problems might not converge at all. For solder paste reflow however, 2-d representations of solder spheres renders solder cylinders, which behave in a markedly different manner from spheres. Figure 4 shows an example, where a line of spheres coalesces into one central sphere and two satellites, while a similar array of cylinders coalesces into a single

S.H. Mannan Solder paste reflow modeling

Figure 1 Volume of fluid method for two fluid flow

Soldering & Surface Mount Technology 14/1 [2002] 18–23

Figure 2 Key modelling processes

cylinder. Given that one of the key modelling aims is to determine the circumstances that result in solder balling and incomplete particle coalescence, it is clear that only 3-d models will be able to give predictive quantitative results. In Figure 5 we see a line of 14 solder spheres coalescing to form two solder masses. In order to minimize the computational time, only 6.5 spheres are represented, with symmetrical boundary conditions at x ¼ 0 used to represent the rest of the spheres. Also, in this special case where the geometry is symmetric around the x-axis, an axisymmetric model has been used. In this type of model, the solution times and convergence properties are closer to that of a 2-d model, while a full 3-d geometry is actually solved (the 2-d geometry is rotated 3608 around the axis). Note that this axisymmetric approach cannot be used for more complex arrangements of spheres because any sphere displaced from the x-axis would appear as a distorted sphere or torus when rotated around the axis. These axisymmetric models can be used to probe the behaviour of single lines of spheres, or lines of spheres connected to a central solder pool. The computer simulation predicts that the particles merge into two or more masses if the number of particles is larger than 12. We can test the accuracy of the simulation by experiment and by agreement with analytical predictions. The solder viscosity is 0.0015 Pas (Beyer et al., 1972)

whereas the flux viscosity is expected to be $ 1 Pas. Thus the flow is dominated by the viscosity of the flux, with the solder viscosity negligible in comparison. The driving force for the solder reflow and coalescence is the surface tension, which has a value of 0.16–0.4 N/m (Murr, 1975) depending on temperature and whether the solder is in contact with flux or not. On dimensional grounds, the time-scale for a reflow process should be given by t ¼ ml=s

ð1Þ

where m is the flux (not solder) viscosity, l is the length scale (particle diameter for two spheres coalescing) and s is surface tension. Using a value of 100 Pas for m, s ¼ 0:4 N=m and 10 mm for l, we obtain t ¼ 2:5 ms which represents the time-scale for two solder spheres to coalesce. The axisymmetric models reproduce this result well (Mannan et al., 2000). If we now consider, a chain of n particles, we obtain a coalescence time-scale of 2.5n ms, which agrees well with the results in Figure 5 (time-scale = 34 ms). Experimental validation of these CFD predictions were attempted using 0.45–0.5 mm diameter sized SnCu solder spheres produced by solder jetting. The particles were placed in a 0.5 mm trench as shown in Figure 6:

Figure 3 2-d modelling of 3-d geometries

[ 19 ]

S.H. Mannan Solder paste reflow modeling

Figure 4 Coalescence patterns of spheres and cylinders

Soldering & Surface Mount Technology 14/1 [2002] 18–23

However, as the temperature of the solder was raised towards the reflow temperature of the solder spheres, the solvents in the flux bubbled away, disturbing the distribution of the particles and invalidating the experiments, since this effect has not yet been incorporated into the CFD models.

Figure 5 Merger of 14 spheres

[ 20 ]

Future experiments will use specifically formulated fluxes to overcome this problem, but this effect may play a role in actual paste coalescence. Similar problems were encountered when attempting to experimentally observe a single mono-layer of solder paste particles coalescing.

S.H. Mannan Solder paste reflow modeling

Figure 6 Sphere coalescence experiments

Soldering & Surface Mount Technology 14/1 [2002] 18–23

2-d models of paste coalescence with oxides The breakdown of the oxides on the solder particles and surfaces to be soldered involve chemical reactions between the constituents of the flux and the solder surfaces. Modelling the detailed breakdown of these oxides presents a formidable chemical modelling challenge and has not been attempted in full in the present work. A programme of work is however underway to quantify the major effects that affect solder paste coalescence. In this work, a simplified chemical analysis is undertaken as follows: Metal þ oxygen ! Oxide

ð2Þ

Oxide þ acid ! by-product

ð3Þ

These reactions proceed with rate constants k1 and k2 which are set independently of each other. The other parameters which affect the chemical reactions include a flux rate for oxygen diffusing into the system from the surrounding

Table I Diffusion coefficients of the various chemical species in each medium Region Diffusing species Acid Oxide Metal Oxygen

Flux

Solder/oxide

0.0002 10212 10212 0.004

1025 10212 1024 1025

atmosphere, diffusion coefficients of the chemical species in each region, and initial concentration of each species. These parameters must all be experimentally determined, but are currently given arbitrary values to test the model development. The oxide is represented in the CFD model by a chemical species advected along with the fluid movement. The oxide and metal chemical species do not diffuse in the flux. The viscosity in the fluid is set as a function of oxide concentration, which should prevent shear deformation of the oxide layer, while allowing translation and rotation of this layer, as the molten solder moves. This is necessary to model situations involving partially removed oxide layers where some solder movement is taking place, dragging partially coalesced solder spheres. Numerical limitations in the CFD code limit the maximum viscosity ratio of neighbouring fluids to , 104. Hence the solder: flux: oxide ratios are set at 1:100:10000. This results in a limited amount of unwanted coalescence to occur before the oxides have been removed in the model. Real oxide layers would also allow elastic modes of deformation which cannot be modelled using the high viscosity fluid approach. Furthermore the ratio of particle diameter: initial oxide thickness is known to be , 104 whereas in a CFD model, the mesh size limits the ratio to , 10. Table I shows the diffusion coefficients of the various chemical species in each medium. The initial distribution of the chemical species is shown in Figure 7. The mesh is shown in more detail in Figure 8. Figures 9 and 10 show intermediate stages of coalescence, showing the break-up of the oxide layer. Oxygen is diffusing in from the outer boundary. If this level of oxygen diffusion is increased sufficiently, then coalescence is prevented from taking place by the reaction of Equation (2).

Figure 7 Initial distribution of chemical species

[ 21 ]

S.H. Mannan Solder paste reflow modeling

Figure 8 Mesh and oxide boundaries

Soldering & Surface Mount Technology 14/1 [2002] 18–23

Figure 9 Oxide layer thinned down and start of break up

Figure 12 Section through coalescing mass

Figure 10 Remnants of oxide layer

3-d models of solder paste

Figure 11 Starting configuration of 3-d model

In the case of a long tube contracting under surface tension (see Figure 5), the tube can split into several masses if the aspect ratio is high enough. In contrast, for a thin sheet of material contracting under surface tension, the mass always coalesces into a single sphere in the centre. This has been tested for a square sheet, with aspect ratios of sheet thickness: width of 50. The cohesiveness of a truly 3-d structure such as a solder paste deposit, will be higher still, and defects will only be caused by incomplete removal of the oxide layers. A monolayer of particles, covered by an oxide layer is shown in Figure 11. One eighth of the model is shown, with symmetrical boundary conditions applied to the x ¼ 0; y ¼ 0 and z ¼ 0 planes. The parameters are set with initial conditions, oxygen flux, diffusion parameters and rate constants similar to the 2-d model, and a similar coalescence pattern emerges. Figure 12 shows the oxide layer broken up, and the paste coalescing into a single sphere.

Summary This paper reports on progress in work on modelling the solder paste reflow and coalescence process. Predictions on the coalescence patterns of strings of solder paste particles were made, and it is hoped to confirm these predictions in the near future. Previous 2-d models incorporating oxides have been extended to 3-d and further work is in progress to identify the reaction rates, diffusion coefficients, initial [ 22 ]

S.H. Mannan Solder paste reflow modeling Soldering & Surface Mount Technology 14/1 [2002] 18–23

chemical species concentrations and other parameters which have been identified as controlling oxide break-up phenomena.

References Beyer, R.T., et al. (1972) “The viscosity of liquid metals“, in Beer, S.Z. (Ed.) Liquid Metals: Chemistry and Physics, Marcel Dekker, New York, NY.

Hirt, C, et al., (1981) “Volume of Fluid (VoF) method for the dynamics of free boundaries”, J. Comp. Physics, Vol. 39 No. 1, pp. 201-25. Kuhmann, J.F., et al., (1998) “Oxidation and reduction of liquid SnPb (60/40) under ambient and vacuum conditions”, J. Electrochem. Soc., Vol. 145, pp. 2138-42. Mannan, S.H., et al. (2000), “Solder paste reflow modelling for flip chip assembly”, Proc. EPTC 2000, Singapore, pp. 103-109. Murr, L.E. (1975), Interfacial Phenomena in Metals and Alloys, Addison-Wesley, New York, NY.

[ 23 ]

Numerical modelling of scanned beam laser soldering of fine pitch packages P.M. Beckett University of Hull, UK A.R. Fleming University of Hull, UK J.M. Gilbert University of Hull, UK D.G. Whitehead University of Hull, UK

Keywords Soldering, Lasers, Electronics assembly Abstract Laser soldering provides a useful tool for the electronics manufacturer and has found a number of successful industrial applications. The laser provides highly controllable localised heating in a manner similar to hand soldering and has distinct advantages over other soldering methods. However, the heating processes that occur in laser soldering are complex making it difficult to predict whether a particular operation will be successful. Numerical modelling provides a valuable tool in answering such questions but the modelling process and the assumptions upon which it is based must be understood if the results are to be reliable.This paper explains the assumptions made in deriving a suitable model and describes the use of a commercial finite element modelling package to aid the understanding of laser soldering processes, with a particular emphasis on single and multi-pass scanned beam soldering operations.

Soldering & Surface Mount Technology 14/1 [2002] 24–29 q MCB UP Limited [ISSN 0954-0911] DOI 10.1108/09540910210416558

[ 24 ]

Introduction Laser soldering typically involves the application of a high power density beam to a small area of a potential joint for a short period. Selecting the appropriate pulse duration and power by trial and error is time consuming and can be unreliable. The manufacture of test devices and printed circuit boards (PCBs) can be expensive, particularly if a large range of trials are to be performed. Mathematical modelling can provide a valuable insight into which heat transfer mechanisms are important and how factors such as component dimensions, PCB characteristics and heating regime affect these. This in turn allows the laser parameters needed for successful soldering to be predicted. In this paper the laser soldering process is briefly described and the development of a numerical model, and its validation, discussed. The use of a range of models to give an understanding of single pass and multi-pass scanned beam soldering are discussed. Finally, conclusions about the uses and limitations of finite element modelling of scanned beam laser soldering processes are presented.

The laser soldering process The use of laser energy to form solder joints has been developed over a number of years (Bohman, 1974; Lea, 1987) and has found a number of successful industrial applications (Beckett et al., 1999). A typical laser soldering process involves the use of a focused beam directed onto the component leads, solder and lands. If the joint is irradiated for sufficient time and with sufficient power then both the solder and the local region of the component and PCB track reach the solder melting temperature, allowing the solder to melt and flow around the lead. Once the laser beam is removed, the solder cools and solidifies to form the required solid joint. This method has distinct advantages over conventional methods, such as Infra Red (IR) reflow, in which the entire assembly is passed through an oven (Beckett et al., 1999). In particular, the concentration of heat in the joint means that heat sensitive components can be soldered with reduced risk of damage. The rapid heating and cooling which are achieved in laser soldering, particularly point-to-point or single scan soldering results in reduced residual stress in the joints and improved metallurgy (Beckett et al., 1997). A number of different laser types have been used for soldering applications (Beckett et al., 1995; Greenstein, 1989) but it has been found that semiconductor diode lasers are advantageous due to the fact that the wavelength of the energy produced is highly absorbed by the metals used in solder but less readily absorbed by common PCB materials.

The current issue and full text archive of this journal is available at http: //www.emeraldinsight.com /0954-0911.htm

These lasers also offer relatively high efficiency, ease of control and fibre optic energy delivery (Beckett et al., 1999). The work described here is focussed on the use of semiconductor diode lasers for the soldering of fine pitch electronic components onto PCBs (Beckett et al., 1997). A typical arrangement for such applications involves the placement of 144 and 208 lead QFP packages with lead pitches of 0.3, 0.4 and 0.5 mm (12,16 and 20 thou) onto FR4 epoxy/fibreglass PCBs with either solid solder or solder paste deposits on the copper “lands”. The laser beam, having a typical diameter of 0.4 mm, is then applied to the joints either as individual pulses, typically of 3–5W for 100– 200 ms, as a single scan with a typical power of 5W scanning at 5 mm/s (as shown in Figure 1: a “still” taken from a video of the scanning of a 0.3 mm pitch device) or with repeated scanning. Selecting the laser pulse parameters (power, duration, scan rate, etc) must balance the need to supply sufficient energy to ensure complete melting while avoiding damage to the device lead, solder or the substrate. Unfortunately, the complexity of the joint geometry and interaction between processes makes this balance difficult to find. Appropriate modelling can provide a valuable insight into these processes and hence the balance of laser parameters. Although considerable research effort has been directed to the modelling of solder heating during conventional IR reflow (Whalley et al., 1992; Kim et al., 1998) and subsequent stress in joints (Bailey, 1999), there has been relatively little consideration given to the very different behaviour found in laser soldering (Beckett et al., 1997). Work carried out by the Authors on the modelling of pointto-point laser soldering and some of the major factors affecting heating is described in (Beckett et al., 2002).

Figure 1 Scan soldering of a 0.3 mm Pitch QFP Device

P.M. Beckett, A.R. Fleming, J.M. Gilbert and D.G. Whitehead Numerical modelling of scanned beam laser soldering of fine pitch packages Soldering & Surface Mount Technology 14/1 [2002] 24–29

Modelling development There are a large number of techniques available for the modelling of heat flow of which one of the most commonly used methods is finite element (FE) modelling. The FE method involves two major steps. The first step is to construct a mesh that divides the object to be modelled into a number of cubic, or wedge shaped elements. In general, the smaller the mesh elements the more accurate the solution but the greater the computation time. A compromise is generally reached by forming a fine mesh around areas of high temperature gradient, with a courser mesh in other areas. The process of constructing a mesh may be automated but is very time consuming for complex geometries. The second stage in the FE method is to solve the /governing the physical processes within each element. In this work a commercial FE modelling package, ANSYS (ANSYS) is used. The major factors affecting laser heating behaviour, which must be considered during model development may be classed as geometric, parametric and temporal. Geometric factors arise from the fact that a typical PCB is a highly complex three-dimensional structure with features several orders of magnitude smaller than the overall board dimensions. To accurately mesh the whole of such a structure would not be viable. Fortunately the localised heating, which is characteristic of laser soldering, means that only a relatively small region around the joint needs to be modelled. However, the extent of this region is not always obvious. The formation of a solder joint, especially if starting from solder paste, is highly complex, involving a change in shape, phase change, fluid flow and, in the case of solder paste, the evaporation of volatile flux and the coalescence of the solder spheres. It would be possible to model these processes individually but to integrate a complete model with the geometry of a real joint would not be viable. Although parameters such as the thermal capacity and conductivity of copper, component lead materials and solid solder are known, many of the other important parameters are less well quantified. The epoxy resin used in the FR4 PCB material has a thermal conductivity of approximately 0.18 Wm 2 1 k 2 1 while, when combined with the glass reinforcement, this increases to around 1.5 Wm 2 1 k 2 1. The varying orientation of the glass fibres within the FR4 makes conduction anisotropic and subject to board-to-board variation. The absorption of optical energy by the elements of a solder joint depends on both the material and the wavelength of the incident energy. The percentage absorption may readily be measured for stable, solid materials (Beckett et al., 1995) but during laser heating, the absorption changes. The absorption can be affected by an oxide coating on the solder and by it becoming molten. In addition, if the solder becomes overheated, its absorption increases significantly, leading to thermal runaway. For these reasons, it is usual, as far as possible, to focus the energy onto the component lead rather than directly on to the solder. However, given that the achievable laser spot size may be greater than the lead width, some energy is likely to impinge upon the solder, and even the substrate.

Model development must take account of temporal factors arising from the disparate time constants involved in laser soldering. Laser heating can cause extremely rapid temperature changes (of the order of 1058C/s) in the surface of a metal target while soldering operations can take several minutes. To efficiently determine both the initial temperature rise and the behaviour during the remainder of the pulse and subsequent cooling period automatic time stepping procedures are required. Given the factors outlined above it is clear that models aiming to represent laser soldering processes could become highly complex. In order to make models manageable, a number of assumptions must be made.

Modelling assumptions and validation The localised heating and relatively short heating duration which are characteristic of laser energy allows accurate results to be obtained by simulating a small portion of a board. The area of board simulated is determined by introducing perfectly insulating boundaries and monitoring the temperature rise at these boundaries. If it is found that there is a significant temperature rise then the boundaries are moved to a greater distance from the heat source and the process repeated. In addition, where symmetry exists in the structure, this fact is used to reduce the area modelled. For all of the models discussed here the material parameters given in Table I are used. Heat conduction through the FR4 PCB material is assumed to be homogeneous. It is assumed that a rectangular block of solid solder, rather than paste, is used and that the solder does not change shape as a result of melting. It is also assumed that no material flow occurs in the solder when it becomes liquid. The phase change in the solder at 1838C is represented using a plot of enthalpy against temperature. These assumptions have been tested by validation against two sets of experimental results. In the first case, individual elements in an array of simple solder/copper/FR4 structures were irradiated with a range of pulse powers and durations and the same structures modelled using ANSYS. The degree of melting, overheating and burning found experimentally was classified and compared with the simulated temperature rise and an excellent agreement found (Beckett et al., 2002). Further validation was provided by measuring the temperature, using a thermocouple formed inside a solder joint, while undergoing laser heating and cooling (Beckett et al., 2002). The resulting temperature profile was compared with the simulated temperature at the hottest point in the joint: on the upper surface of the solder directly below the centre of the laser spot (point A shown in Figure 2) and coldest point: at the furthest corner of the solder mass (point B).The temperature profiles, seen in Figure 3, show good agreement. The differences that are observed are a result of the manner in which the temperatures are measured and the way the simulated temperatures are logged. It may be noted that the phase change at 1838C results in a reduced rate of temperature rise as extra energy is absorbed to effect the phase change rather than causing further heating.

Table I Material parameters used Units 2 1

2 1

Specific heat Jkg 8C Thermal conductivity Wm 2 18C 2 1 Density kgm 2 3 Reflectance3 % Transmittance3 % 3 Absorbance % 1 Alloy used for component leads 2 1.6 mm thick FR4 without surface coating 3 Optical properties for wavelength of 805 nm

Solder

Invar1

Copper

FR42

230 50.2 8410 60 40

460 10.7 8137 60 40

383 380 8954 95.2 4.8

2000 1 2000 26.6 52.1 21.3

Scanned laser modelling A range of models have been developed based on the assumptions described to represent differing geometries, joint dimensions and heating scenarios. The simplest scenario to model is the heating of a single joint and although this is not the focus of this paper, it does provide a valuable insight into subsequent results. The details of this work are described in (Beckett, 2002). What this analysis shows is that the heating behaviour is significantly affected by factors such as the dimensions of the copper tracking [ 25 ]

P.M. Beckett, A.R. Fleming, J.M. Gilbert and D.G. Whitehead Numerical modelling of scanned beam laser soldering of fine pitch packages

Figure 2 Geometry of a two joint model

Soldering & Surface Mount Technology 14/1 [2002] 24–29

(where increasing the thickness from 0.04 mm to 0.12 mm results in a doubling of the time required to complete a typical joint) while the internal connectivity of the device has no significant effect on heating.

Single scan soldering Making a single scan of a series of joints provides some advantages over point-to-point heating. In particular, heat conduction through the substrate can reduce the time required to form a series of joints and the control of this operation is generally simpler, provided a constant laser power and scan speed are used.

Figure 3 Comparison of simulated and experimental temperature during laser oldering

In order to gain an insight into the single scan process, a set of four identical leads of a 0.4 mm pitch device is modelled as shown Figure 4. To represent the scanning, a series of discrete stages were simulated in which the laser spot irradiates the appropriate combination of lead, solder and FR4. To simplify the modelling of this process the 0.4 mm diameter circular laser spot was replaced by a 0.4 mm £ 0.4 mm square with equivalent total power. Two scanning procedures were investigated. In the first, the scan began with the laser centred on the first lead while the second scenario involved starting the scan 0.4 mm away from the centre of the first joint to give a “run-up”. The temperature distribution midway through a scan can be seen in Figure 4. The temperature at points A and B shown in Figure 2 were logged. These may be expected to represent the hottest and coolest points of the solder during the heating process. The temperature profiles at points A and B on each joint are shown in Figure 5 and Figure 6 for the two scanning scenarios described. Figure 4 shows that a relatively small area of FR4 around the laser spot undergoes significant heating due to the relatively low thermal conductivity of the FR4. During animation it can be seen that the temperature rise advances ahead of the laser spot once the spot impinges on the joint. This is due to the greater conductivity of the copper and solder. It can be seen from the temperature plots that in both cases the temperature attained in the second joint is significantly higher than in the first. The increase is due, in part, to conduction through the FR4 substrate but also to the fact that the edge of the laser spot impinges on the copper, solder and lead of a joint while most of the spot area is still irradiating the previous lead. Conduction into the remainder of the joint causes the temperature to rise well before the laser spot is centred on the joint. A further temperature rise occurs in subsequent joints and this is due to conduction through the FR4. The temperature profile indicates that a steady state will be reached well below the overheating temperature of 5008C. An important difference between the two scanning scenarios is that when no run-up is used (Figure 5) the maximum temperature in the first joint only reaches approximately 1808C while the minimum temperature reaches only approximately 1508C. Thus there is no melting in this first joint. With the run-up (Figure 6) the maximum temperature reaches 2258C and the minimum reaches approximately 1808C. Thus, although there is not complete melting, the majority of the solder would be molten. The fact that the peak in the minimum temperature profile is flattened is due to the significant energy being used to produce the phase change. This prediction of the need for a run-up has been found to be fulfilled in experimental trials (using the same device dimensions and laser parameters) in which the first joint is only successfully completed if a run-up is used.

Multi-pass scanning Although single scan soldering has proved viable, both through simulation and experimentation, repeated or multi pass scanning does offer some potential advantages. If all of the joints become molten simultaneously then surface tension will tend to cause the device to self-align. In addition, simultaneous melting of all joints allows removal of faulty devices for rework of the PCB. One difficulty found in this process is that as the number of joints scanned increases the average power, and hence temperature rise, of each joint decreases. Although this could be compensated for by increasing the laser power, the rapid heating which is characteristic of laser heating means that surface damage occurs rapidly as the power increases, even when applied for a short period. In practice it has been found that a 0.4 mm diameter spot with a power above around 10 W causes board damage even for very short pulse durations. [ 26 ]

P.M. Beckett, A.R. Fleming, J.M. Gilbert and D.G. Whitehead Numerical modelling of scanned beam laser soldering of fine pitch packages

Figure 4 Temperature distribution during scanned soldering simulation

Soldering & Surface Mount Technology 14/1 [2002] 24–29

Figure 5 Temperature profile without “run-up”

Figure 6 Temperature profile with “run-up”

[ 27 ]

P.M. Beckett, A.R. Fleming, J.M. Gilbert and D.G. Whitehead Numerical modelling of scanned beam laser soldering of fine pitch packages

To model the multi-pass laser scanning of a typical device with many leads would be extremely time consuming and memory intensive, both for meshing and subsequent simulation. However, useful results may be gained from greatly simplified models.

Soldering & Surface Mount Technology 14/1 [2002] 24–29

Pulsed laser model In multi-pass scanning an individual lead experiences a series of short laser pulses and will also receive heat energy through conduction from adjacent joints. This situation may be studied by simulating two adjacent joints one of which is irradiated by a series of short pulses. Using the model geometry shown in Figure 2 the first joint was irradiated

Figure 7 Temperature profile with pulsed laser

with 10 cycles of a 10 W laser for 10 ms followed by 90 ms cooling (to represent 10 cycles of a multi pass scanning of a 10 terminal device in 1s). The temperatures at points A and B in the joint being irradiated (Joint 1), representing the hottest and coolest points in the solder were logged. The corresponding points on the second joint were also monitored. In this model only Joint 1 is irradiated and so its temperature is not affected by thermal conduction from adjacent joints. This effect can however be approximated by adding the temperature rise found in Joint 2 onto that in Joint 1. This method for including the effect of conduction from adjacent joints is not intended to give precise figures but to indicate whether soldering or de-soldering is feasible. The resulting temperatures profiles are shown in Figure 7. It can be seen that the maximum temperature in Joint 1 rises very rapidly (200008C/s) during the pulse but falls again rapidly when the pulse ends. The minimum temperature in Joint 1 (point B) shows a smaller variation. The peaks in the temperature at point B in Joint 1 reach the solder melting point after around 0.75s (indicating complete melting of individual joints) while the troughs would exceed 1838C after a few more cycles (corresponding to simultaneous melting of all joints). Thus it can be inferred that the 10 W laser would be capable of both soldering and de-soldering a 10 terminal device although the precise times required to achieve this cannot be determined. It should be noted that the hottest part of the joint (directly below the laser) experiences repeated very large and rapid temperature excursions which may impact subsequent reliability.

Averaged power model

Figure 8 Approximate heating for various joint counts using 10 W scanned laser

[ 28 ]

Although the above approach to modelling multi pass scanning does give a useful insight into the soldering of devices with a small number of leads, it cannot readily be extended to higher pin count devices due to the greatly extended simulation times which would be involved. If we consider a 100 lead device then the individual pulse time for each lead is reduced to 1/100th of the cycle time (1 ms in this case) but the time for the temperature to reach equilibrium increases to around 500 s implying 500 000 iterations of the FE model. Even for the simplified joint model described here the computation time involved would be excessive. In order to understand the conditions under which multi-pass scan soldering would be viable it is possible to use a less computationally expensive approach in which the average temperature rise is assessed by applying the mean of the applied laser pulse power continuously to a joint. Thus a 10 W laser applied to a 100 lead device may be modelled by a power of 0.1 W applied continuously. Although this will not capture the peaks and troughs within each cycle, it will approximate the average behaviour. It does, in effect, represent the heating that would occur if the laser were scanned at infinite frequency. This approach does not take account of the fact that the cooling rate will increase during the temperature peaks above that in the average case. However, provided the scan rate is high, as is usually the case in practice, the temperature excursions are relatively small (around 208C for a 1ms pulse) and so this error is likely to be small. Since the same power is applied to all joints in a device it is only necessary to model a single joint with perfectly insulating boundaries at either side. Using this approach, the heating behaviour of 100, 166 and 200 lead devices has been studied for a 10W scanned laser (corresponding to average laser powers of 0.1 W, 0.06 W and 0.05 W, respectively) with an ambient temperature of 258C. The resulting heating curves for the coolest point in the solder are shown in Figure 8. It can be seen that the 100 joint device reaches the solder melting temperature of 1838C in approximately 190 s while it is apparent that neither the 166 or 200 joint devices will reach melting temperature.

P.M. Beckett, A.R. Fleming, J.M. Gilbert and D.G. Whitehead Numerical modelling of scanned beam laser soldering of fine pitch packages Soldering & Surface Mount Technology 14/1 [2002] 24–29

It is common practice in both laser soldering and other localised soldering processes to apply background heat to raise the ambient temperature of the device. The level of background heat must be set to balance the limitations imposed by heat sensitive devices and the rate at which the joint temperature can be raised with limited power input. Figure 8 shows the effect of applying 1258C background heat to the heating of a 200 joint device. It can be clearly seen that the joints reach melting temperature after approximately 90s. It may also be implied from Figure 8 that the 200 joint device would just reach melting temperature with a background temperature of approximately 908C while the 166 joint device would require a background temperature of at least 758C to achieve the simultaneous melting required for de-soldering. These values must be taken as approximate given the assumptions made in the model but they do give an indication of the feasibility or otherwise of certain important manufacturing processes which could not be assessed experimentally without considerable effort.

Conclusions The process of developing a simple model of laser soldering operations, and its validation, has been described. It has been shown that such a model can give a valuable insight into the underlying phenomena and can guide the selection of appropriate process parameters. Using a model of several leads undergoing scanned laser heating provides both a qualitative appreciation of the heat flow process and a quantitative measure of the effect of heat flow through the substrate. This emphasises the importance of using a run-up to ensure that the first joint scanned reaches the required temperature, a finding that is confirmed in experimental trials. Modelling of multi-pass scanning by the use of a model of two joints subject to pulsed laser energy provides a simple mechanism for predicting the approximate extent of short term temperature variations and the longer term temperature trends. This shows that although the multipass scan soldering of a 10 lead 0.5 mm pitch device is viable, the thermal stress imposed on the joint is considerable. By modelling the effect of the average power of a rapidly scanned laser it is shown that 0.5 mm pitch devices with up to around 100 leads could be soldered and de-soldered using a 10 W laser although taking a

considerable amount of time. However for packages with a greater number of leads significant background heat is required to ensure complete soldering or de-soldering. While the precise temperature profiles predicted using these models is liable to include errors due to the modelling simplifications, the effect of these is no greater than those arising from device to device or board to board variations and so do give valuable information in setting process parameters. The research on which this paper has been based has been supported by the Engineering and Physical Sciences Research Council under grant GR/H 89951 and by IMB (UK), Siemens AG and Diomed Ltd whose support is gratefully acknowledged.

References ANSYS, Inc., (n.d.), Canonsburg, PA 15317. Bailey, C, Wheeler, D and Cross, M (1999) “An integrated modelling approach to solder joint formation”, IEEE T Compon Pack T 22: (4) 497-502 December. Beckett, P.M., Fleming, A.R., Foster, R.J., Gilbert, J.M. and Whitehead, D.G. (1995) “The application of semiconductor diode lasers to the soldering of electronic components”, J. Optical & Quantum Electronics, Vol. 27. Beckett, P.M., Fleming, A.R., Gilbert, J.M. and Whitehead, D.G. (1997) Practical Application of Laser Soldering in Electronics Assembly IEEE International Symposium on Industrial Electronics, Portugal, pp. 59-64. Beckett, P.M., Fleming, A.R., Gilbert, J.M. and Whitehead, D.G. (1999) “The laser in manufacture- its use in the soldering of electronic assemblies”, Trans Institute of Measurement and Control, Vol. 21 No. 1, pp. 2-7. Beckett, P.M., Fleming, A.R., Gilbert, J.M. and Whitehead, D.G. (2002) “The finite element modelling of laser soldering for electronic assemblies”, Int. Journal of Numerical Modelling, Vol. 15 No. 3, May. Bohman, C.F. (1974) The Laser and Microsoldering, Society of Manufacturing Engineers: Technical Paper No. AD74-810, Mich. 48128, p. 19. Greenstien, M. (1989) Applied Optics, Vol. 28 No. 21, pp. 45-95. Kim, M.R., Choi, Y.K., Lee, G.B., Chung, I.Y., Kim, J.D. and Lee, J.H. (1998) “Thermal investigation of an infrared reflow oven with a convection fan”, KSME Int J Vol. 12 No. 5, pp. 972-79 October. Lea, C. (1987) “Laser soldering of surface mount assemblies”, Hybrid Circuits, Vol. 12, p. 36. Whalley, D.C., Ogunjimi, A.O., Conway, P.P. and Williams, D.J. (1992) “The process modelling of the infra- red reflow soldering of printed circuit board assemblies”, Journal of Electronics Manufacture, Vol. 2 No. 1, pp. 23-9.

[ 29 ]

A simplified model of the reflow soldering process David C.Whalley Loughborough University, Wolfson School of Mechanical and Manufacturing Engineering, Loughborough, UK Stuart M.Hyslop Loughborough University, Wolfson School of Mechanical and Manufacturing Engineering, Loughborough, UK

Keywords Soldering, Reflow

Nomenclature

Abstract Previous models of temperature development during the reflow soldering process have typically used commercially available, general purpose, finite difference/finite element modelling tools to create detailed three dimensional representations of both the product and of the reflow furnace. Such models have been shown to achieve a high degree of accuracy in predicting the temperatures a particular PCB design will achieve during the reflow process, but are complex to generate and analysis times are long, even when using modern high performance computer workstations.This paper will report on the development of a simplified model of the process, which uses less complex representations of both the product and the process, together with a simple numerical solver developed specifically for this application, whilst achieving an accuracy comparable with more detailed models. In the simplified model, the product is divided into elements, which are represented using a two-dimensional mesh of thermal conductances linking thermal masses. The values of these conductances and masses are calculated based on the averaged properties of the PCB material and attached components within the area of each of the elements. The representation of the specific reflow furnace is based on measurements of the temperature and level of thermal convection at each point along the length of the furnace, thereby avoiding the necessity of making detailed measurements of the furnace geometry and air flow velocities. The combination of these two simplification techniques allow the reduction of analysis time for a relatively simple PCB from in the order of an hour on a high performance Unix workstation to under a second on a Pentium class PC running Microsoft Windows.

A C EA FC FR FK

The authors gratefully acknowledge the financial support of Datapaq Limited.

Soldering & Surface Mount Technology 14/1 [2002] 30–37 q MCB UP Limited [ISSN 0954-0911] DOI 10.1108/09540910210416440

[ 30 ]

area heat capacity nodal plan area convective flux radiative flux resistive flux between a node and all adjacent nodes H(x ) convective constant for position x in furnace K thermal conductivity Q net heat flux into node R thermal resistance SA total surface area of a node t time T nodal temperature air/heater panel temperature in furnace TH x location along the furnace length X length of a node Y width of a node Z thickness of the PCB effective average emissivity of a node 1E 1H(x ) effective average heater panel emissivity at position x Dt time step s radiation absorption constant for node z Stephan Boltzmann constant

Introduction Infra-red and convection reflow are non-equilibrium processes in that the product temperature is never in thermal equilibrium with the heat source. The reflow furnace process settings must therefore be tailored to each individual product design in order to ensure the optimum time/temperature profile for that particular PCB assembly. In recent years, there has been a transition from earlier designs of reflow furnace, where radiation was the dominant heat transfer mechanism, to more modern designs where convective heat transfer is dominant. The higher level of heat transfer achievable with a convection dominant furnace has reduced variations in the temperature profile between different products and also the spread in the temperatures within an individual product and has consequently increased the process window. The accelerating transition to lead free soldering processes will however reduce the process window, thereby placing renewed emphasis on techniques for reflow profile optimisation. The standard approach to reflow profile set-up has been to attach a number of thermocouples to an example of the product to be assembled, with the choice of thermocouple location being based on a combination of engineering judgement and experience of similar products and component package types. The product is then passed

The current issue and full text archive of this journal is available at http: //www.emeraldinsight.com /0954-0911.htm

through the reflow furnace with a “first guess” set of process temperatures and a data logger records the temperatures measured by the thermocouples. The initial process settings are then modified until the required profile is obtained. This profile modification process can be assisted by very simplified models of the process, such as those within existing commercial reflow profiler software packages typified by those available from Datapaq and KIC, e.g. (Kazmierowicz, 2001). The use of such software reduces the number of iterations before the required profile is obtained, but does not eliminate the initial instrumented reflow process run using an example of the new product. A computational model of the process has the potential to entirely eliminate this on-line set-up procedure through the construction of a detailed model of the product using the PCB CAD data, and could even be used to ensure the compatibility of a PCB design with the reflow process before it is released to manufacture. Such a model can also eliminate the risk that the chosen thermocouple locations do not cover the full spread of reflow profiles within the product. The feasibility of such models was demonstrated by Whalley et al. (1990,1992) and Sarvar and Conway (1998) showed the high level of accuracy obtainable, provided accurate materials properties were available. Another approach to simulation of the process has also been reported by Eftychiou et al. (1993) who have developed a two-dimensional model of the process. Similar thermal processes in the ceramics industry have also been simulated by, for example, Hurst and Pulko (1993) but this type of product is composed of a single material and both product and furnace are geometrically less complex. There are however a number of disadvantages to the modelling techniques described by Whalley and Sarvar, the principal of which are the construction of a detailed physical description of the specific reflow furnace to be modelled and the long analysis time. The description of the reflow furnace has required detailed measurements of the geometry of each zone within the furnace and the measurement of the airflow velocity incident on the PCB throughout the furnace. The analysis time even on a high performance workstation has been found to be in the order of tens of minutes for even a moderately complex PCB design, which might be acceptable if the model only has to be run a small number of times for a specific product, but is not compatible with algorithms designed to automatically search for the optimum process settings, as these typically require a very large number of modelling iterations. This paper reports a modified approach to the modelling of the reflow process, which uses a simple twodimensional (2D) model of the PCB assembly and avoids the need to create a detailed physical description of the reflow furnace by using sensors to measure the furnace’s heat transfer performance. The model also uses a solver constructed specifically for the application and which is therefore highly efficient. Analysis times in the order of a second are consequently achievable using a typical PC running under the Microsoft Windows operating system.

David C. Whalley and Stuart M. Hyslop A simplified model of the reflow soldering process Soldering & Surface Mount Technology 14/1 [2002] 30–37

The simplified solver In representing the PCB assembly with a 2D model, it is assumed that there is no significant variation in temperature through the thickness of the PCB. This assumption is based on the low Biot number, Bi, associated with heat transfer through the PCB thickness. The Biot number is the ratio of the internal thermal resistance to the external thermal resistance and if significantly less than one indicates that internal temperature gradients will be small. For a uniform slab of material having a thickness Z and thermal conductivity K which is heated from both sides with a heat transfer coefficient of H, the Biot number is given by: Bi ¼

H£Z 2£K

ð1Þ

For a typical PCB with a thickness of 1.6 mm and a K of 0.31 W/m K, and for a typical average heat transfer coefficient for a convection reflow process of 50 W/m2 K, the Biot number is therefore only 0.129. It is also assumed in the model that there is no variation in temperature between the components and the underlying area of the PCB. This assumption is more difficult to justify, but is in general supported by the results reported by Sarvar and Conway (1998) and by the success of the approach being reported here. For components with a large thermal mass and which are poorly thermally connected to the PCB this assumption may break down and care will be necessary in developing the required library data for such components. General purpose finite element/finite volume analysis software such as the well known commercial software ANSYS and ABAQUS, are able to employ a variety of element shapes and sizes in order to mesh the region of interest, but the solver described here is based on a uniform recti-linear grid of elements in order to maintain simplicity, high execution speed and easy presentation of results. Each element contains a central node, to which all of the thermal mass of that area of the assembly is assigned. The nodes are then interconnected by thermal conductances. If the time steps chosen are small enough that the boundary conditions and materials properties can be assumed to remain constant over the time step then the basic formulae to be solved at each node can be based on an explicit time integration approach and for each time iteration is: T tþDt ¼ T t þ

assembly over the node, EA is the nodal plan area, and z is the Stephan Boltzmann constant 3:74 £ 10216 W m22 K24 : Provided the heights of the components are small in relation to the dimensions of the thermal nodes, then they will not significantly increase the effective area of the node absorbing radiation, but they will increase its effective emissivity although this effect is complex to calculate. Inter-node shadowing due to tall components is also complex to evaluate accurately and this approach assumes that it is negligible. The total conductive flux, Fk, is the sum of the four fluxes to/from the adjacent nodes: X ðT t ðx; yÞ 2 Y t Þ ð7Þ FK ¼ x¼^1;y¼^1 Rðx; yÞ

Q £ Dt C

ð2Þ

where: T is the nodal temperature (K), t is time (s), Dt is the time step (s), Q is the net heat flux into the node (W), C is the heat capacity of the node (J/K) The net heat flux into the node is the sum of the convective flux, FC, and radiative flux, FR, through the top and bottom surfaces, and the total conductive flux FK from the four adjacent nodes, i.e.: Q ¼ FC þ FR þ FK

ð3Þ

The convective flux, FC, is a function of the difference in temperature between the node and the air temperature at the current location, x, within the furnace, and also of the incident air velocity at that location: F C ¼ ðT H ðxÞ 2 T t Þ £ HðxÞ £ SA

ð4Þ

where: Tt(x,y ) is the nodal temperature for adjacent nodes, R(x,y ) is the thermal resistance between the node and adjacent node x,y (K/W) The thermal resistance from the centre of any element to its edge is calculated as follows (Figure 1): Rx ¼

X 2£K£Y £Z

ð8Þ

Ry ¼

Y 2£K£X£Z

ð9Þ

These node centre to edge thermal resistances are calculated for each node and then added to those for each adjacent node to give the node centre to centre resistances. A programme to iteratively solve the above equation has been implemented in C++. The executable is only 370 kB including an animated graphical display of the predicted PCB temperatures. Input and output files use a comma separated variable (CSV) format for ease of export/import to/from other software such as spreadsheet programs.

Generation of the product description As identified in the previous section the data required for each thermal node within the product model are its effective in-plane conductivity, both in the X and Y directions, its thermal mass, average emissivity, and the convection area. For an area of the PCB unpopulated by tracks or components this data is simple to calculate. For a node with a significant density of conductive tracks, but no components, the additional contribution of the tracks to the thermal conductivity of the node must be taken into account. This additional thermal conductivity will be principally in the direction of the tracks, with little effect perpendicular to them. The additional thermal mass of the tracks should also be included, although it will have a very small effect unless the board has a large number of layers. Provided the board has a solder mask, which substantially increases the emissivity of tracks, then the emissivity is unlikely to be significantly affected by the presence of tracks unless there

Figure 1 Thermal conductances to adjacent nodes

where: TH is the air/heater panel temperature (K), H(x ) is the convective constant for position x (W/m2 K), SA is the total surface area for the node (m2). The radiative flux, Fr, is given by: F R ¼ ðT H ðxÞ4 2 T 4t Þ £ 1H ðxÞ £ s

ð5Þ

where 1H(x ) is the effective average heater panel emissivity at position x and s is the radiation absorption constant for the node:

s ¼ 1E £ EA £ z

ð6Þ

where 1E is the effective average emissivity of the PCB [ 31 ]

David C. Whalley and Stuart M. Hyslop A simplified model of the reflow soldering process

is a large unbroken area of copper. The convection area will be the same as the nodal plan area. Where an area of the PCB is populated with components, calculation of the nodal properties is slightly more complex and the following modifications to the bare PCB properties must be made: Conductivity: Components much smaller than the element size, such as ceramic chip capacitors and resistors, will have little effect on conductivity and can be safely ignored. Larger components, particularly those with a metal lead-frame, will however have a significant effect on the local in-plane thermal conductivity. An effective thermal conductivity was therefore calculated for each of the IC packages, taking into account the relative thickness of the lead-frame and package body. The thermal conductivity of each node under an IC was then modified taking into account the proportion of its area covered by the IC. Thermal mass: The additional thermal mass of an individual component is the product of its volume, density and SHC and the total thermal mass of a node is therefore the sum of the PCB thermal mass and the individual component thermal masses. The components were however weighed so their volume and density did not have to be measured. Any component lying on the boundary of two or more elements was simply split between them in proportion to the component area within each element. Emissivity: As noted in the previous section an accurate calculation of the effective emissivity of a node is quite complex. In most modern reflow ovens only a small proportion of the heat transfer is by radiative heat transfer and obtaining a precise value for the emissivity is therefore less important. It was, therefore, decided to simply

Figure 2 The test PCB

Soldering & Surface Mount Technology 14/1 [2002] 30–37

Table I Materials properties for the modelling trial Material PCB laminate Chip capacitor Chip resistor Black Epoxy

Emissivity

Density (kg/m3)

S.H.C. (J/kg K)

Effective K (W/m K)

0.90 0.96 0.94 1.00

1,885 — — —

1,100 765 765 850

0.31 — — 2.1– 3.2

Figure 3 Comparison of an idealised furnace temperature profile with measurement

[ 32 ]

David C. Whalley and Stuart M. Hyslop A simplified model of the reflow soldering process

Figure 4 H sensor temperature measurements

Soldering & Surface Mount Technology 14/1 [2002] 30–37

calculate nodal emissivity based on the (plan) area weighted average of the emissivities of the materials present within an element. Convection area: The addition of components to a thermal node will increase the total surface area available to convective heat transfer. If the components sit close to the PCB, then there will be little airflow under the components and it can therefore be assumed that the bottom of the component and the area of PCB underneath it only play a small part in convective heat transfer. The additional convection area due to a component is therefore only the

area of its sides, i.e. the component height multiplied by the length of its perimeter. Where a component overlies an element border this additional surface area is split between the elements as for its thermal mass.

The test PCB Figure 2 shows the test board used in the modelling trials, which is about 20 cm by 15 cm and has a total of 37 components, including one 44 pin PLCC, and a mix of SO and

Figure 5 H values calculated from sensor data

[ 33 ]

David C. Whalley and Stuart M. Hyslop A simplified model of the reflow soldering process

Figure 6 Time temperature profile of the test board

Soldering & Surface Mount Technology 14/1 [2002] 30–37

chip components. In order to test the model a relatively coarse mesh of 20 by 15 elements was used, resulting in a total of 300 thermal nodes. The materials properties used in setting up the model of the test board are listed in Table I. Ideally the product description would be generated directly from the CAD data, but in order to test the new modelling approach this data was generated for an existing test PCB design using a spreadsheet. Data entry for the test board proved extremely time consuming, but testing of the significance of various parameters was then very straightforward.

Oven performance data A reflow furnace profile is often thought of as consisting of a number of zones of uniform temperature with step changes in temperature at the boundaries between zones as shown in Figure 3. Any real oven will however depart from this idealised view in a number of ways as is also shown in Figure 3. Firstly, there will be some degree of error between the oven set points and the actual heater panel temperature. There may also be some variation in heater panel temperature within a zone and the zones will interact to some extent, meaning that the transition from one zone to the next is not instantaneous. The sharpness of this transition between zones is dependant on various factors such as the tunnel height and the details of the air flow velocities. There are also areas of the furnace that are not under active temperature control, typically gaps between zones used for conveyor supports and the entrance and exit areas, which nevertheless may be significant to the overall reflow profile. The process temperatures throughout the length of the furnace (TH) were obtained using a low mass thermocouple and data logger for several combinations of process settings. It was then possible to use this data to predict the temperatures that would occur for any desired combination of zone set points. The details of this profile prediction process will be presented elsewhere. The variation of the convective heat transfer coefficient (H ) throughout the oven was then obtained using a second thermocouple attached to a piece of material of very low emissivity, and with a known thermal mass (CS) and surface area (AS). This sensor was supported only by the thermocouple wires, so there was negligible conductive heat transfer to/from it. As the [ 34 ]

emissivity was very low, heat transfer by radiation could also be assumed to be negligible so almost all of the heat flow to/from the sensor was via convective heat transfer. The values of H(x ) throughout the furnace length could therefore be calculated by solving what is effectively a combination of Equations (2) and (4) with the radiative and conductive terms removed: dT s ðT H 2 T S Þ £ H £ AS ¼ dt CS

ð10Þ

where TS is the temperature of the sensor of known thermal mass and surface area. For the initial trials of the model, a Quad QRS7 reflow furnace was used. Figure 4 shows the measured air temperatures and convection sensor temperatures along the furnace length for a particular reflow profile. It was found that the values of convective heat transfer coefficient calculated using Equation (10) appeared very noisy. This was initially believed to be primarily due to a combination of noise in the temperature measurements and their resolution. Averaging of the H values over several seconds

Figure 7 Predicted test PCB temperature distribution

David C. Whalley and Stuart M. Hyslop A simplified model of the reflow soldering process

Figure 8 Comparison of measured and predicted temperature profiles for the QFP components (a) Whole profile (b) Near reflow

Soldering & Surface Mount Technology 14/1 [2002] 30–37

reduced this noise, however comparison of the H values calculated from three runs through the furnace showed that the remaining “noise” was fairly repeatable from run to run and is therefore concluded to be due to real spatial fluctuations of the air flows within the furnace. Figure 5 shows calculated values for H including averaging.

a 28 pin PLCC and to both the corner and centre of the 44 pin PLCC. To reduce thermal degradation of the test board, and any consequent changes in its thermal properties, the reflow furnace was set to a slightly cooler profile than would typically be used in production. Figure 6 shows the resulting time temperature profiles for the five thermocouples.

Results

Modelling results

Experimental data acquisition

The model was run using the measured H data in Figure 5, together with a temperature/distance profile which was predicted for the furnace process settings used in the experimental run. This profile prediction was made using

A sample of the test board shown in Figure 2 had thermocouples attached to an 1812 capacitor, a 20 pin SOIC,

[ 35 ]

David C. Whalley and Stuart M. Hyslop A simplified model of the reflow soldering process

Figure 9 Comparison of measured and predicted temperature profiles for the 1812 capacitor and SOIC (a) Whole profile (b) Near reflow

Soldering & Surface Mount Technology 14/1 [2002] 30–37

data from measurements made on the furnace for different process settings. Figure 7 shows the predicted distribution of temperatures in the PCB at a particular instant in time during the reflow process and Figures 8 and 9 show a comparison of the predicted and measured time/temperature profiles for the five thermocouple locations. From Figures 8 and 9, it can be seen that there is excellent agreement between the model and experimental results throughout the entire reflow process. The average difference in peak temperature between model and experiment was 3.58C and the maximum difference was less than 58C. The analysis time was 0.44 s on a 300 MHz Intel Pentium processor with 64 MB of RAM. [ 36 ]

Conclusions It has been demonstrated that a simplified model of the reflow soldering process, in which the boundary conditions are inferred from in process measurements, can provide extremely accurate results. This modelling approach greatly reduces the time required to create and run the simulation compared with models where boundary conditions are established from detailed process equipment geometry and measurements or CFD predictions of the airflow velocities. Further work is required to test the limits of accuracy of the approach developed, both for other soldering furnaces, particularly those where heat transfer is IR dominant, and for

David C. Whalley and Stuart M. Hyslop A simplified model of the reflow soldering process Soldering & Surface Mount Technology 14/1 [2002] 30–37

more complex PCB assemblies. In addition to its use in process optimisation during the new product introduction process, the modelling approach is simple enough to use during the PCB design stage to ensure compatibility of the design with available process hardware. The process data acquisition and modelling approach described here is probably also applicable to other thermal processes where significant variations in product thermal mass require product specific process optimisation, such as in paint curing, ceramic kilning processes and in the food processing industry.

References Eftychiou, M.A., Bergman, T.L. and Masada, G.Y. (1993) “A detailed thermal model of the infrared reflow soldering process”, ASME Journal of Electronic Packaging, Vol. 115, pp. 55– 62.

Hurst, A.I., Pulko, S.H., (1993), “Modelling the transient thermal experience of large pieces of vitreous china ware during firing”, Proceedings of the IASTED International Conference, Applied Modelling and Simulation, Anaheim, CA, USA, pp. 150 – 3. Kazmierowicz, P.C. (2001) “Increasing reflow process efficiency and yield with automated reflow setup”, APEX Conference, San Diego, CA, January. Sarvar, F. and Conway, P.P. (1998) “Effective modelling of the reflow soldering process: basis, construction and operation of a process model”, IEEE Transactions on Components, Packaging and Manufacturing Technology Part C: Manufacture, Vol. 21 No. 2, pp. 126 – 33. Whalley, D.C., Ogunjimi, A.O., Conway, P.P. and Williams, D.J. (1992) “The process modelling of the infra-red reflow soldering of printed circuit board assemblies”, Journal of Electronics Manufacture, Vol. 21 No. 1, pp. 23 – 9. Whalley, D.C., Williams, D.J. and Conway, P.P., (1990), “Thermal modelling of temperature development during the reflow soldering of SMD Assemblies”, Proceedings of the 6th ISHM International Microelectronics Conference, Tokyo, May, pp. 385 – 94.

[ 37 ]

CFD modelling of the flow field inside a reflow oven Hao Yu Laboratory of Electronics Production Technology, Helsinki University of Technology, Helsinki Finland Jorma Kivilathti Laboratory of Electronics Production Technology, Helsinki University of Technology, Helsinki Finland

Keywords Lead-free soldering, Reflow, Computational fluid dynamics Abstract Because of new requirements related to the employment of leadfree manufacturing and the diversity of components and metal finishes on high density printed circuit boards, better understanding and control of the reflow process is required in order to achieve acceptable yields and reliability of SMT assemblies. Accurate control of the temperature distributions within components and boards during the reflow process is one of the major requirements, especially in leadfree assembly. This paper outlines a scheme for reflow modelling and presents an oven-level model of the steady state flow-field inside a reflow oven, which will be needed in subsequent transient analysis and small-scale modelling. The model is constructed by utilising the advanced computational fluid dynamics (CFD) technology using commercial software. The computational results are discussed and compared with measured data.

We wish to thank Mr Toni Mattila for his help in operating the laboratory’s assembly line and collecting experimental data. We are also grateful to the National Technology Agency and Finnish electronics industry for financial support.

Soldering & Surface Mount Technology 14/1 [2002] 38–44 q MCB UP Limited [ISSN 0954-0911] DOI 10.1108/09540910210416459

[ 38 ]

Introduction The electronics manufacturing industry is facing new challenges, which tend to increase the complexity of materials selection and assembly processes. These challenges follow from continuing demands for higher performance and miniaturisation of consumer electronic products, which imply that a larger number of components must be accommodated onto higher density printed circuit boards. Components having large differences in size and/or in thermal masses such as BGA, QFP, CSP and flip chip as well as small passive components, have to be soldered simultaneously which makes it necessary to control more accurately the temperature distributions within reflow ovens. Other challenges follow from the more or less inevitable shift to Pb-free manufacturing processes. The most obvious requirements of Pb-free assembly are related to the high melting points of commercially available lead-free solder pastes. For example, eutectic Sn –Ag–Cu alloy has the eutectic temperature of 217 8C, which is notably higher than that of the conventional Sn–Pb solder. In most cases higher reflow temperatures are unavoidable, which increases the risk of thermal damage and narrows process windows significantly. Additionally, new metal finishes on printed wiring boards and components terminations should be developed, and new considerations of atmosphere and flux chemistry seem to be indispensable for Pb-free assembly, as well. Consequently, it becomes ever more important not only to develop reflow oven technology but also have a better understanding and control of the reflow process. Melting, interfacial reactions between lead-free solders and various metal finishes, fluxing, wetting and solidification during the soldering process have to be investigated in more detail. Temperature acts as a primary variable affecting these reactions and therefore controls the formation of many soldering defects such as voiding, solder balling, etc. Even though thermocouples are widely used to monitor temperatures inside a reflow oven, such measurement methods are also subjected to more restrictions and difficulties when components become smaller. The influence of the thermocouple is sufficient to disturb the temperature field around small components severely. For this reason, computational simulation is of great interest to visualise the temperature field on the circuit board. If conducted properly, a combination of theoretical modelling and experimental measurement may present a powerful tool to predict detailed temperature variation and distribution during reflow processes. Several researchers (Eftychiou et al., 1992; Eftychiou et al., 1993; Kim et al., 1996; Sarvar and Conway 1996; Sarvar and Conway 1998a, b; Kannapel and Przekwas 1997)

The current issue and full text archive of this journal is available at http: //www.emeraldinsight.com /0954-0911.htm

have contributed to reflow modelling since the early 1990’s and one of them (Eftychiou et al., 1992) seems to be successful despite the use of a 2D simplification. However, these works paid much more attention to radiation rather than convection because Infra-Red (IR) heating was widely used in 1990’s and only low speed flow appears during the IR reflow process. This is not the situation currently facing industry since forced convection has already replaced the role of IR heating. A more comprehensive model for forced convection heating is now indispensable, which is why further development of reflow modelling relies on the computational fluid dynamics (CFD) technique. Reflow modelling is difficult to handle and the ultimate objective can be only realised by hierarchical steps. At least two stages of model development are necessary: a largescale model or oven model, and a small-scale model or board model. The reason is that oven information can only be collected for the large-scale model, while details have to be presented by the small-scale model. The large-scale model is an intermediate step toward the ultimate objective and its output works as input for the small-scale model. Since the movement of the board disturbs the flow field, transient simulation may be required for the oven model, but the steady state field has to be modelled first as a preliminary requirement. In this paper, we attempt to model the heating gas inside a reflow oven theoretically. The temperature distribution and flow pattern inside our furnace are predicted, which will be used later in our following work. The commercial CFD software, CFD-ACE+ (http://www.cfdrc.com), is adopted to establish a 3D model for the oven.

Physical description The reflow oven to be modelled in this paper is an EWOS 5/1 offered by EPM (Niebling et al., 1998). It has six modules with forced convection, as shown in Figure 1. All of the modules have the same physical structure but working at different temperatures, five are used for the heating and one for the cooling. A small amount of outflow is generated through the exhaust pipe between the heating and cooling zones, which is balanced by the inflows through both ends of the oven. This flow rate mixes the fluids from different modules, but it is fairly small compared with large inner flow of the modules. The structure of the modules is shown in Figure 2, heating (or cooling) gas is impinged onto the PCB through an array of nozzles, then a tangential fan sucks the gas away through a space between the nozzle pipes, draws it over an electro-heating spiral (or water-cooling device) and conveys it back to the nozzles again. The recirculation fan and electro-heating spiral are not shown in the figure because they will not be modelled directly in this work. Instead, a circular outlet is used to represent the entrance of fan and this boundary condition will be explained later.

Hao Yu and Jorma Kivilathti CFD modelling of the flow field inside a reflow oven

Figure 1 The structure of the oven to be modelled

Soldering & Surface Mount Technology 14/1 [2002] 38–44

In the present model, the heating gas is air but other fluids can be used as well. Since the gas temperature varies from 333 to 573 K, temperature dependency of fluid properties has to be considered. Evaluation methods for these properties and their value ranges in this model are listed in Table. I. Based on these properties and the velocity range of 7.5– 11 m/s, obtained from experimental readings, the Reynold number for the nozzles takes a value from 632 to 2,335. Despite this relatively small value, turbulent vorticity develops easily near nozzle and the critical Reynold number is no more than 100. Hence, standard two-equation k-1 turbulent model is adopted in this paper. Partial differential equations for continuity, momentum, turbulence energy and its dissipating rate are solved numerically. This procedure is performed using the software CFD-ACE+. In order not to introduce too much mathematics, detailed description will not be presented in this paper and readers with interest should read materials about CFD technology (references). Conduction occurs through solid objects inside the oven. The conveyor is the only object considered in our model and its detailed structure has been neglected. The material is assumed to be aluminum alloy with the following properties: density r ¼ 2770 kg=m3 ; heat capacity C p ¼ 875 J=kg K; and heat conductivity k ¼ 177 W=m K:

Geometric model An important step for solving this problem using CFD is creating a mesh for the domain. Figure 3 shows the mesh structure for the modules. In order to restrict the grid size, five different parts are generated separately: (1) An unstructured mesh at the top of the module, which is introduced to present the complex geometry of this part; (2) A structured mesh for the space between the nozzles and the conveyor, which is unevenly distributed in x direction so as to get denser grids in the vicinity of nozzle; (3) A structured mesh for the space containing the conveyor, which is unevenly distributed in z direction so that more cells are allocated near the conveyor; (4) A structured mesh below the conveyor, which is uniform; (5) A structured mesh for space between the vertical slabs at the bottom of the oven. These five parts are connected with each other by “arbitrary interfaces”, which is a special function offered by CFDACE+ program. The mesh structure offers flexibility to model complex geometry as well as convenience for following transient analysis. In addition to the mesh for the modules, additional grids are also created for the exhaust pipe and for the oven’s ends. Finally a three dimensional geometric model is set up for the reflow oven, as shown in Figure 4. Nearly one million cells are employed altogether.

Figure 2 Structure of modules inside the oven

Table I Variable ranges for the flow field and determination of Reynold number for nozzles Property/variable

Evaluation method

Range of value

Temperature (T ) Velocity (V ) Width of nozzle (D ) Density (r )

– – – PM ; M ¼ 29 Ideal Gas Law r ¼  3=2 RT T T0 þ S Sutherland’s law m ¼ m0 m0 ¼ 1:716 £ 1025 ; T 0 ¼ 273; S ¼ 111 T0 T þS Second-order Polynomial Cp ¼ C0 þ C1 T þ C2 T 2 C0 ¼ 1019:53; C1 ¼ 20:1561; C2 ¼ 0:00035 Constant mC p Prandtl number k ¼ Pr rVD ReD ¼ m

333 –573 K < 7.5– 11 m/s 0.004 m 0.6168 – 1.0614 kg/m3

Viscosity (m ) Specific heat capacity (Cp) Prandtl number (Pr) Thermal conductivity (k ) Reynold number (ReD)

1:9993 – 2:9293 £ 1025 N s=m2 1006– 1045 J/kg K 0.70 0.02873– 0.04373 W/m K < 632– 2335

[ 39 ]

Hao Yu and Jorma Kivilathti CFD modelling of the flow field inside a reflow oven

Figure 3 Mesh structure used to model the flow field inside module

Soldering & Surface Mount Technology 14/1 [2002] 38–44

Boundary conditions The boundary conditions needed in the geometric model are presented in this section.

Nozzle The flow rate through the nozzles is controlled by recirculation fans and it is determined by the fan control curve as seen in Figure 5. It presents the relationship between the flow rate and the pressure difference over the fan. Unfortunately, the pressure drop inside the modules can not be obtained with sufficient accuracy enough to get the exact flow rate for the nozzles. A reasonable approximation therefore has to be made to define this boundary condition. According to the manufacturer of the oven, uniform heating can be realised inside the modules and the nozzle velocity does not vary significantly from point to point. Considering the average measured velocity to be 7.5 – 11 m/s, the flow rate is within the range of 0.195– 0.285 m3/s, which is marked in Figure 5. It can be seen that the flow rate has an almost linear dependence on the pressure drop. However, our preliminary modelling suggests that the pressure drop does not differ significantly at different module temperatures. Hence, it is assumed that all of the nozzles have the same velocity. This velocity is then selected to be 9.639 m/s (equal to a flow rate of 0.25 m3/s for a module) which could be modified later if necessary.

Figure 5 The fan control curve of the recirculation fan offered by the manufacturer and the corresponding working area of the fans

Figure 4 Three-dimensional geometric model for the reflow oven

[ 40 ]

Hao Yu and Jorma Kivilathti CFD modelling of the flow field inside a reflow oven Soldering & Surface Mount Technology 14/1 [2002] 38–44

Figure 6 Measured nozzle temperature and its relationship with setpoints for module 1, 3, 5, (Setpoints for other modules in the oven are listed in Table II)

The gas temperature is controlled according to the setpoint temperature for each module. Since the temperature is not regulated near the nozzle, however, the nozzle temperature is slightly different. A series of measurements were thus performed and parts of the results are shown in Figure 6. The nozzle temperature of neighbouring modules is also plotted in the figure. It is interesting that the nozzle temperature varies linearly with the setpoint temperature and as long as the setpoint does not change significantly, its neighbouring nozzle temperature is fairly stable. These facts offer some ideas on how to estimate nozzle temperature according to the setpoint. Since a turbulent model is applied, the turbulent energy k and it’s dissipating rate 1 should be defined for the nozzles as well. This can be done by assuming a turbulent percentage Tu ¼ 5% and a turbulent viscosity ratio mt =m ¼ 50 : 3 k ¼ Tu 2 V 2 2 1¼

ð1Þ

rCm k 2 mt

ð2Þ

Where Cm is a constant with a value of 0.09. These parameters for the nozzles in the model are listed in Table II.

Outlet Since the recirculation fan is located far enough away from the nozzles, it is safe to assume that its details have negligible effect on the flow field within the region in which we are interested. As a result, a subtle fan model is not thought to be necessary and was beyond the scope of this work. The circular outlet is therefore defined at the top of each module to represent the entrance of fan. The mass flow rate of this outlet should be equal to the sum of that for all of the nozzles in the module based on the principle of recirculation fan. Unfortunately, this can not be defined directly in the CFD program we used. This flow rate requirement is therefore realised approximately by defining a proper pressure value, which is obtained through trial and error.

Exhaust and end The outflow rate through the exhaust between the heating and cooling modules varies from 50 to 150 m3/h. The exact value depends on the conveyor width and is determined by the size of board passing through the reflow oven. Since a small test board is used in our project for lead-free soldering, the lower limit, 50 m3/h, is adopted to define the outlet boundary of the exhaust pipe. This outflow rate is balanced by inflows through both ends of the oven, where inlets are also defined. It is assumed that both ends of the oven have the same flow rate, and the inflow rate is determined accordingly. This inflow can be modelled as a laminar stream with low velocity.

Table II Setpoint temperature combination for modules and variables defining the boundary conditions

Module number 1 2 3 4 5 6

Setpoint temperature (K)

Nozzle temperature (K)

Velocity (m/s)

Turbulent energy (m2/s2)

Turbulent energy dissipation rate (m2/s3)

448 458 438 478 558 333

431 445 429 460 542 337

9.639 9.639 9.639 9.639 9.639 9.639

0.348414 0.348414 0.348414 0.348414 0.348414 0.348414

7.433995 7.040296 7.493189 6.655148 5.050462 11.36631

[ 41 ]

Hao Yu and Jorma Kivilathti CFD modelling of the flow field inside a reflow oven

Figure 7 Calculated temperature distribution inside the oven. (The displaying plane is the intersection through the middle-line of the conveyor, as indicated by the dashed line in Figure 2c)

Soldering & Surface Mount Technology 14/1 [2002] 38–44

Experimental measurements Temperatures at several points inside the reflow oven (EWOS 5/1) were measured in our production line for

comparison with the model. The thermocouples were placed inside the oven and fixed with aluminium tapes, exposing their tips to the heating gas at certain position. A thermometer was then used to read the temperatures from

Figure 8 Enlarged pictures for modules about the temperature presented in Figure 7 as well as the calculated flow field

[ 42 ]

Hao Yu and Jorma Kivilathti CFD modelling of the flow field inside a reflow oven Soldering & Surface Mount Technology 14/1 [2002] 38–44

outside the oven by leading the thermocouple wires/cables out. The nozzle temperatures mentioned above were also measured by the same procedure, while thermocouples were placed inside nozzles in this case.

Results and discussion The computational procedure was performed in SGI Origin 2,000 supercomputer using the CFD-ACE+ program. The calculated temperature fields are presented in Figure 7, which shows an intersection through the middle-line of the conveyor. Enlarged pictures for the modules are shown in Figure 8 together with the calculated flow field. All the modules were found to be working independently and the flows from different modules did not disturb each other significantly. After being impinged through the nozzles, the heating gas flows downward and expands horizontally. Where the flows from the adjacent meet each other, the heating gas rises up and flows out between the nozzle pipes. High velocity flow is found to be restricted within the space between the nozzles and the conveyor, while the flow speed under the conveyor is limited. According to the calculated turbulence energy presented in Figure 9, the heating gas has the highest turbulent energy inside this region as well, indicating that effective forced convection can be realised in this region during the reflow process. Figure 9(b) also shows that the most turbulent regions occur in the middle of each module and they should be accompanied with the strongest heating/cooling effect during the reflow process. In accordance with the flow pattern described above, the temperature inside the modules is uniform and a large temperature gradient only appears between the modules with a large temperature difference. The highest temperature gradient occurs between the heating and cooling zones, and the starting portion of the first module also has a significant temperature gradient. These regions extend into the corresponding modules and may have some influence on the reflow profile. Conduction through the conveyor rails causes smearing of the temperature difference between adjacent modules as indicated in Figure 8. This effect can be only roughly

evaluated in the present model since the detailed structure of the conveyor has not been considered. It is worth noting that reflow soldering is a process where many physical phenomena involving heat transfer take place at the same time. Despite strong motive for thermal analysis, computational simulation of reflow process has not been paid enough attention, partly because of difficulties on theoretical modelling. This is especially true for ovens with forced convection. Hence, although the capability of CFD technology has expanded rapidly over the last ten years, reflow modelling should be still conducted with care. It is critical for reflow modelling to be practical and realistic at the same time. A larger and complicated model is obviously more scientific but not always the best solution from the practical point of view if computing time and expense has to be considered. Many simplifications, for physical model, geometry as well as boundary conditions, have been used in our model and their influence on the calculated results should be carefully examined. In some senses, comparison with measured data can be used to judge the reality, although its effectiveness is sometimes limited by inadequate data. In order to examine how well the model agrees with the real flow field, the temperature was roughly measured at several points inside the oven using thermocouples. Generally they agree with the simulation results, as shown in Figure 10. Some difficulties in establishing the geometric model can be also noticed. Even with the capability of supercomputer, the calculation is still found to be timeconsuming, which imposes a restriction on the geometric model scale. On the other hand, unfortunately, the nozzles in the oven are small and a fine mesh is desired in order to present more information. As a result, a compromise between these two factors has to be made and the mesh is not refined enough to reveal many details around the nozzles. This is why small-scale modelling is needed later. In spite of these disadvantages, the success of the present model is still suggested based on two facts: (1) the calculated flow field is in accordance with previous analysis before modelling and (2) measured data are fitted. The original purpose of oven-level modelling has been achieved since only general features of the flow field are expected to be input for following modelling and then board-level

Figure 9 Calculated turbulence energy (k ) distribution inside the oven. (a) K distribution inside module 5 as an example; (b) Variation of k along x direction inside the oven. The plotting line locates in the middle of the upper surface of the conveyor, which is a little bit higher than the region where boards pass during reflow process. The starting point of module 1 is selected as the origin for x-axis. (the length of each module is 0.51 m)

[ 43 ]

Hao Yu and Jorma Kivilathti CFD modelling of the flow field inside a reflow oven

Figure 10 Comparison between calculated and measured temperature inside the oven. (The plotting line is the same as that in Figure 9b)

Soldering & Surface Mount Technology 14/1 [2002] 38–44

modelling would take the task of predicting detailed results.

Summary This paper presents a CFD model of the air heat flow inside a reflow oven and describes the geometric mesh and boundary conditions in detail. The results are found to be coincident with measured data. They will be used in following transient analysis as well as in a board-level model, which will be presented elsewhere.

References Conway, P.P., Whalley, D.C., Wilkinson, M. and Williams, D.J. (1997), “Automated adaptive control of the reflow soldering of electronic assemblies”, IEEE/CPMT Int’l Electronics Manufacturing Technology Symposium, pp. 229-36. Eftychiou, M.A., Bergman, T.L. and Masada, G.Y. (1992), “Thermal effects during infrared solder reflow—Part II: a model of the reflow process”, ASME J. Electronic Packaging, Vol. 114, pp. 48-54.

[ 44 ]

Eftychiou, M.A., Bergman, T.L. and Masada, G.Y. (1993), “A detailed thermal model of the infrared reflow soldering process”, ASME J. Electronic Packaging, Vol. 115, pp. 55-62. Kannapel, M.D. and Przekwas, A.J. (1997), “CFD model of the reflow soldering process and equipment”, report of CFD Research Corporation. Kim, M.R., Daewoo, I.A.E., Choi, Y.K., Lee, G.B., Chung, I.Y. and Kim, J.D. (1996), “Thermal investigation of an infrared reflow furnace with a convection fan”, Intersociety Conference on Thermal Phenomena, pp. 211-16. Niebling, R., Stehling, R., Nowottnick, M. and Wittke, K. (1998), “Combined IR and forced convection in reflow soldering”, SMT/August, 138-40. Sarvar, F. and Conway, P.P. (1996), “Effective transient process modeling of the reflow soldering of printed circuit assemblies”, Intersociety Conference on Thermal Phenomena, pp. 195-202. Sarvar, F. and Conway, P.P. (1998a), “Effective modeling of the reflow soldering process: basis, construction, and operation of a process model”, IEEE Transactions on Components, Packaging and Manufacturing Technology-Part C, Vol. 21 No. 2, pp. 126-33. Sarvar, F. and Conway, P.P. (1998b), “Effective modeling of the reflow soldering process: use of a modeling tool for product and process design”, IEEE Transactions on Components, Packaging and Manufacturing Technology-Part C, Vol. 21 No. 3, pp. 165-71.

Analysis on solder ball shear testing conditions with a simple computational model S.W. Ricky Lee Electronic Packaging Laboratory, Department of Mechanical Engineering, Hong Kong University Science & Technology, Kowloon, Hong Kong Xingjia Huang Electronic Packaging Laboratory, Department of Mechanical Engineering, Hong Kong University Science & Technology, Kowloon, Hong Kong

Keywords Plastic ball grid array, Solder joints, Finite element analysis, Reliability Abstract This paper introduces a simple computational model for the analysis on the solder ball shear testing conditions. Both twodimensional (2-D) and threedimensional (3-D) finite element models are used to investigate the effect of shear ram speed on the solder ball shear strength of plastic ball grid array (PBGA) packages. An effective thickness is identified for the 2-D finite element analysis. By using this effective thickness as a scale factor, it is shown that the 2D model is feasible for the study of 3-D problems. The computational model is validated by experimental data in terms of load-displacement curves. The results from both testing and modeling indicate that the shear ram speed has a substantial effect on the solder ball shear strength. In general, faster ram speed can result in higher ball shear strength. Therefore, the characterization of solder ball shear strength is loading rate-dependent.

The Research Grant Council of Hong Kong sponsored this study through the grant of HKUST6231/01E to the Hong Kong University of Science and Technology (HKUST). The authors wish to acknowledge this support.

Introduction In the past decade, BGA technologies have overwhelmed the whole IC packaging industry. This is because BGA packages have many advantages over conventional ones. Among them are larger number of I/Os, self-alignment capability, more robust assembly process and better thermal/electrical performance (Lau, 1996; 2000; Lau and Pao, 1997). Furthermore, BGA packages are compatible with surface mount technology (SMT), consequently, leading to high throughput and low cost assemblies for mass production. For surface mount components (SMCs), the solder joints are not only the passage of electrical signals, power and ground, but also act as the mechanical support to hold the module in position on the printed circuit board (PCB). Therefore, solder joint reliability is a major concern for BGA packages. Currently, the most popular method to evaluate the strength of solder ball attachment is the ball shear test. The typical value of solder ball shear strength for PBGA packages may be around 1000 gf (for 30-mil solder balls) (Coyle et al., 1999; Lee et al., 2000a; Levis and Mawer, 2000). Although such tests are simple and convenient to implement, there is not much mechanics justification to interpret the testing results. Most people still use the method of A-to-B comparison to determine the acceptance of packages. In July 2000, JEDEC published a new standard, JESD22-B117, for BGA solder ball shear tests (JEDEC, 2000). In this publication, the testing condition is specified in terms of the shear ram height only. The specification is that the gap between the shear ram tip and the ball mounting surface should be larger than 0.05 mm (2 mils) and smaller than (or equal to) 25 per cent of the ball height. However, another important testing parameter, namely, the shear ram speed (loading rate), is not addressed in this standard. The lack of specification in the shear speed may cause confusion for the comparison of solder ball shear strength characterized with different loading rates (Coyle and Solan, 2000; Hasegawa et al., 2000). Therefore, further research effort is required for understanding the rationale of solder ball shear tests and for investigating the effect of shear speed. The present study is aimed at establishing a finite element model for the analysis of ball shear tests in order to evaluate the solder ball attachment strength of PBGA packages. In particular, the emphasis is placed on investigating the effect of the shear speed. The results obtained from this computational modeling analysis will be validated by experimental data.

Solder ball shear tests Soldering & Surface Mount Technology 14/1 [2002] 45–48 q MCB UP Limited [ISSN 0954-0911] DOI 10.1108/09540910210416468

The purpose of solder ball shear tests in the present study was for model validation. Solder balls under evaluation were standard 0.76 mm (30 mils) 63Sn-37Pb spheres for PBGA

packages. The substrates for ball attachment were BT laminates with a thickness of 0.46 mm. The solder bond pads were solder-mask-defined with an opening of 0.6 mm in diameter. The metallization of bond pads was Cu with Ni/Au electro-plating. The solder balls were attached to the bond pads using a standard SMT reflow profile. After reflow, the average ball height and diameter were 0.64 mm and 0.78 mm, respectively. The ball shear tests were conducted using a Dage-4000S machine. In the present study, the shear ram height was fixed at 10% of the solder ball height (i.e. 64 mm), which complies with the new JEDEC standard, but the shear ram speed varied (20 mm/s, 50mm/s, 100 mm/s, 200 mm/s and 500 mm/s). For each testing condition, 25 solder balls were sheared. The peak shear forces were recorded and the calculated average was taken as the ball shear strength. The testing results of ball shear strength are presented in Table I.

Finite element model In the present study, both 2-D and 3-D finite element models were built and used in the analyses. The geometric models were based on the average dimensions measured from the cross-section of specimens for solder ball shear tests. A commercial finite element code, ANSYS Release 5.6, was used to perform the computational analysis. Figure 1 shows the 2-D model simulating the ball shear test. The finite element meshes consisted of the solder ball, the Cu bond pad, and the BT substrate, which were modeled by 8-node plane strain elements (PLANE183). The shear ram was considered a rigid body. A feature in ANSYS using the surface-to-surface target element (TARGE169) and the contact element (CONTA172) was employed to simulate the contact between the shear ram and the solder ball. The left, right, and bottom sides of BT substrate were fixed as rigid boundaries. In the present 2-D model, there were 1582 nodes, 490 PLANE183 elements, 8 TARGE169 elements and 40 CONTA172 elements. Since this was a timedependent non-linear analysis, both large deformation and transient options were enabled in the analysis. For the three-dimensional analysis, 20-node solid elements (SOLID186) were used to mesh the 3-D model. Due to the symmetry, only one half of the solder ball was modeled as shown in Figure 2. Similar to the 2-D modeling above, the surface-to-surface target element (TARGE170) and contact element (CONTA174) were employed to simulate the contact between the shear ram and the solder ball during the solder ball shear test. In the 3-D model, there were 11277 nodes, 2520 SOLID186 elements, 249 TARGE170 elements and 240 CONTA174 elements. The material properties and other modeling features for the 3-D model were the same as those assumed in the 2-D analysis.

Constitutive models for the materials The current issue and full text archive of this journal is available at http: //www.emeraldinsight.com /0954-0911.htm

In the present computational model, all materials, except the solder ball, are assumed to behave as linear elastic materials. The material properties used in modeling are listed in [ 45 ]

S.W. Ricky Lee and Xingjia Huang Analysis on solder ball shear testing conditions with a simple computational model

Table I Ball shear strength under various shear speed

Soldering & Surface Mount Technology 14/1 [2002] 45–48

Fixed ram height (64 mm) Testing Modeling

20

50

Shear speed (mm/s) 100

200

500

1131 (4%) 1143

1171 (5%) 1202

1244 (6%) 1244

1350 (4%) 1291

1418 (3%) 1350

Note: (*) – Standard deviation, unit – gf.

Figure 1 2-D finite element model

ANSYS, equation (2) is called the implicit creep model and the equation is further modified as:
 d1 C4 ¼ C1 ½sinhðC 2 sÞC3 exp 2 ð3Þ dt T where s is the effective stress and de /dt is the steady-state effective creep strain rate. The values of C1, C2, C3, and C4 in equation (3) for the eutectic solder are given in Table III (Lee et al., 2000b).

Determination of effective thickness

Table II. In the present study, the solder is considered as a visco-plastic material, which follows the Garofalo-Arrhenius creep constitutive (Lau, 1995):
h
  t in dg G Q ¼C exp 2 sinh v ð1Þ kQ dt Q G where g is the steady-state creep shear strain; dg/dt is the steady-state creep shear strain rate; t is the time; C is a material constant; G is the temperature-dependent shear modulus; Q is the absolute temperature (8K); v defines the stress level at which the power law stress dependence breaks down; t is the shear stress, n is the stress exponent; Q is the activation energy for diffusion mechanism; and k is the Boltzmann constant ð8:617 £ 1025 eV=8KÞ: Darveaux and Banerji modified equation (1) to
 h  t in dg Q ð2Þ ¼ A sinh exp 2 kT dt B and have characterized the coefficients, A, B, n, and Q, for a few solder materials (Darveaux and Banerji, 1992). In

Figure 2 3-D finite element model

The direct result from the present computational analysis is the force-displacement curve simulating the ball shear test. It should be noted that, for the 2-D plane strain analysis, the force output from ANSYS was in the unit of N/mm (force per unit thickness). However, the actual diameter of solder was 0.78 mm ( , 1 mm). In order to obtain the meaningful shear force for comparison, an effective thickness (scale factor) should be defined. Based on the fact that the ball mounting pad has a diameter of 0.6 mm and the solder ball has an axisymmetric geometry, it may be concluded that the effective thickness of the 2-D model must be smaller than 0.6 mm. With such an observation, the scale factor for the effective thickness could be estimated as follows: In the present study, the shear test condition with 64 mm ram height and 100 mm/s shear speed was chosen as the base case. For this base case, the peak shear force obtained from ANSYS was 2969 gf/mm. Compared with the corresponding experimental data (1244 gf, see Table I), the effective thickness, t, can be determined as: t¼

1244 gf ¼ 0:42 mm 2969 gf =mm

ð4Þ

Figure 3 shows the original shear force-displacement curve from ANSYS and the modified curve with an effective thickness of 0.42 mm for the base case. Subsequently, the same value of effective thickness was applied to all other 2-D analyses. If the same effective thickness could yield to good agreement between testing and modeling for all cases (which will be shown in the next section), then it is reasonable to conclude that the selected effective thickness

Table II Elastic material properties for modeling Materials

E (MPa)

n

r (g/cm3)

63Sn-37Pb Cu pad BT Substrate

29,800 128,700 14,000

0.40 0.34 0.39

8.41 8.31 1.2

Table III Creep option input parameters for ANSYS C1 (s21) 339.0102

[ 46 ]

C2 (MPa21)

C3

C4 (8K)

0.062653

3.3

6,360

S.W. Ricky Lee and Xingjia Huang Analysis on solder ball shear testing conditions with a simple computational model

Figure 3 Modification of force-displacement curve

Soldering & Surface Mount Technology 14/1 [2002] 45–48

Model validation and discussion Figures 5 –8 present the results of 2-D analyses for the shear speeds of 20, 50, 200 and 500 mm/s, respectively. All corresponding experimental data are plotted as well for comparison. The peak shear forces, which are regarded as the ball shear strength, are also summarized in Table I. It is observed that the testing and modeling results are in good agreement for all cases. Since the scale factor was determined using the base case only, it is reasonable to conclude that the selected effective thickness was not a random coincidence. It should carry a certain physical meaning, which represents the scale factor between the 2-D

Figure 5 Force displacement curves for 204 mm/s

was not a random coincidence. It should carry a certain physical meaning, which represents the scale factor between the 2-D and 3-D analyses. Furthermore, the fact that the chosen effective thickness is smaller than 0.6 mm also enhances the confidence in the selected value.

Comparison between 2-D and 3-D analyses The force-displacement curve obtained from the 3-D analysis for the base case is presented in Figure 4 and compared to the corresponding results from the 2-D analysis and experimental data. It is observed that the 2-D modeling fits the experimental data better than the 3-D analysis. In general, it is considered that the 3-D model should give more accurate results than the 2-D analysis. However, due to the accumulation of numerical truncation errors, this may not be always true for the large scale computation in the non-linear time-dependent analyses. In the present study, the degrees of freedom for the 2-D and the 3-D analyses were 3164 (1582 nodes £ 2) and 33831 (11277 nodes £ 3), respectively. With one order of magnitude in the number of degrees of freedom, the accumulation of numerical errors might play a significant role making the 3-D analysis a worse case. In addition, the 3-D analysis is usually very time consuming. For the comparison shown in Figure 4, while the 3-D analysis took more than 3 days, the 2-D model only required 3 h using the same computing facility. Therefore, the 2-D analysis should be a better choice for the present study, as long as the proposed effective thickness model is validated. Figure 4 Comparison of force-displacement curves

Figure 6 Force displacement curves for 50 mm/s

Figure 7 Force displacement curves for 200 mm/s

[ 47 ]

.

S.W. Ricky Lee and Xingjia Huang Analysis on solder ball shear testing conditions with a simple computational model

Figure 8 Force displacement curves for 500 mm/s

Soldering & Surface Mount Technology 14/1 [2002] 45–48

.

and 3-D analyses. Based on the good agreement in comparison, the concept of effective thickness is validated. For the ease of summarizing the general trends of ball shear strength, the data given in Table I are plotted in Figure 9. From this figure, one can find that, the ball shear strength increases with the increase of shear speed. Another point to be noted is that, although in general the shear strengths from testing and modeling are in good agreement, there is a notable discrepancy for the cases with shear speed faster than 100 mm/s. From this point of view, it may be deduced that the ideal solder ball shear testing conditions are the cases with shear speed slower than 200 mm/s.

Summary and conclusions In the present study, both 2-D and 3-D finite element models were developed for the analysis of solder ball shear tests. The effect of shear speed on solder ball shear strength was studied not only by experimental investigation, but also by computational modeling. The results are summarized as follows. .

The shear speed has substantial effects on the solder ball shear strength. The data from both testing and modeling indicates that faster shear speed would results in higher ball shear strength.

Figure 9 Effect of shear speed on the ball shear strength

[ 48 ]

The results from testing and 2-D modeling were in good agreement. An effective thickness was identified for the 2-D plane strain analysis. With such a scale factor, it is feasible to study 3-D problem with a 2-D finite element model. The ideal solder ball shear test conditions were recommended to be the cases with shear speed slower than 200 mm/s.

It should be noted that although the present analyses were performed for solder balls of BGA package with 30 mil solder balls, the methodology may be applied to other cases such as BGA packages with smaller solder balls and the solder bumps of flip chips. With the assistance of computational modeling, it seems that it is possible to forecast the ball shear strength under various testing conditions as long as a base case is tested in advance. In addition, with further stress analyses, it may be possible to achieve more meaningful comparisons among different packages. The results obtained from this study should be very helpful for the electronics manufacturing industry to interpret their testing data and determine the acceptance criteria for the products with solder ball attachment.

References Coyle, R.J. and Solan, P.P. (2000), “The influence of test parameters and package design features on ball shear test requirements”, 26th IEEE/CPMT International Electronics Manufacturing Technology Symposium, pp. 168-77. Coyle, R.J., Holiday, A., Mescher, P., Solan, P.P., Gahr, S.A., Cyker, H.A., Dorey, J.K. and Ejim, T.I. (1999), “The influence of nickel/gold surface finish on the assembly quality and long term reliability of thermally enhanced BGA packages”, 24th IEEE/CPMT International Electronics Manufacturing Technology Symposium, pp. 23-35. Darveaux, R. and Banerji, K. (1992), “Constitutive relations for tinbased solder joints”, Proceedings of IEEE Electronic Components & Technology Conference, pp. 538-51. Hasegawa, K., Takahashi, A., Noudou, T., Nakajima, S., Takahashi, A., Nomoto, M. and Nakaso, A. (2000), “Electroless NiP/Pd/Au plating for semiconductor package substrate”, SMTA International Proceedings of the Technical Program, Edina, MN, pp. 225-31. JEDEC (2000), “BGA Ball Shear”, JEDEC Solid State Technology Association JESD22-B117, July. Lau, J.H. (1995), Ball Grid Technology, McGraw-Hill, New York, NY. Lau, J.H. (1996), Flip Chip Technologies, McGraw-Hill, New York, NY. Lau, J.H. (2000), Low Cost Flip Chip Technologies, McGraw-Hill, New York, NY. Lau, J.H. and Pao, Y.H. (1997), Solder Joint Reliability of BGA, CSP, Flip Chip, and Fine Pitch SMT Assemblies, McGraw-Hill, New York, NY. Lee, S-W.R., Newman, K. and Hu, L. (2000b), “Thermal fatigue analysis of PBGA solder joints with the consideration of damage evolution”, Packaging of Electronic and Photonic Devices, EEP-Vol. 28, November 5-10, FL, pp. 207-12. Lee, S-W.R., Yan, C.C., Karim, Z. and Huang, X. (2000a), “Assessment on the effect of Electroless nickel plating on the reliability of solder ball attachment to the bond pads of PBGA substrate”, Proc. 50th ECTC, May 21-24, NV, pp. 868-73. Levis, K.M. and Mawer, A. (2000), “Assembly and solder joint reliability of plastic ball grid array with lead-free versus lead-tin interconnect”, Proc. 50th ECTC, May 21-24, NV, pp. 1198-204.

Optimisation modelling for flip-chip solder joint reliability S. Stoyanov Computing and Mathematical Sciences, University of Greenwich, London, UK C. Bailey Computing and Mathematical Sciences, University of Greenwich, London, UK M. Cross Computing and Mathematical Sciences, University of Greenwich, London, UK

Keywords Optimization, Flip chip, Underfill, Reliability Abstract This paper details and demonstrates integrated optimisation-reliability modelling for predicting the performance of solder joints in electronic packaging. This integrated modelling approach is used to identify efficiently and quickly the most suitable design parameters for solder joint performance during thermal cycling and is demonstrated on flip-chip components using “no-flow” underfills. To implement “optimisation in reliability” approach, the finite element simulation tool – PHYSICA, is coupled with optimisation and statistical tools. This resulting framework is capable of performing design optimisation procedures in an entirely automated and systematic manner.

Soldering & Surface Mount Technology 14/1 [2002] 49–58 q MCB UP Limited [ISSN 0954-0911] DOI 10.1108/09540910210416477

Introduction The “face-down” attachment of silicon dies to organic substrates, known as flip-chip assembly, has a number of important advantages such as greater performance, low cost, high I/O number and distribution, and low weight, etc, (Lau, 1996). While this technology is continuously driven to satisfy the market demands for faster, smaller and cheaper products, the reliability of such components is becoming a major concern for the electronic manufacturing industry. During thermal cycling, high levels of stress and strain develop across the solder joints due to the extreme thermal miss-match (CTE difference) between the materials used in the package. As results of this, solder joints will degrade over time and are very likely to fail under thermal fatigue (Lau, 1996). The success of flip-chip relies in part on improving thermo-mechanical reliability offered by applying underfill encapsulation technology in electronic assemblies (Lau, 1996; Dai and Ho, 1997). The method commonly used to dispense underfill into the gap between the die and substrate is based on capillary flow (surface tension phenomenon) technology. At present, the assembly process (Figure 1a) is slowed down because the additional time consuming step of applying and curing the underfill after the reflow process. An innovative method of underfilling, called “no-flow” technology (Wong et al., 1998; Shi and Wong, 1999), has been developed and proposed as an alternative to conventional underfill. This “no-flow” alternative applies underfill material before the reflow process and uses the reflow process to cure the underfill and fuse the solder joints at the same time (Figure 1b). Despite the clear advantages over traditional assembly process, no-flow technology brings a number concerns for thermo-mechanical reliability. This is primarily due to the change in material properties of current no-flow underfill materials, and experimental results indicate that thermal fatigue lifetime decreases. Recent research, e.g. Wong’s group (Wong et al., 1998; Shi and Wong, 1999; Shi et al., 2000), is trying to achieve desirable changes in the underfill epoxy properties via inorganic silica fillers. Understanding and defining appropriate material and geometric parameters for flip-chip packages, assembled using this novel no-flow technology, is crucial for subsequent product reliability. Finite Element modelling techniques have been used extensively to predict thermal stress in solder joints. Coupled with suitable fatigue models these tools provide the ability to gain an insight into the effect that changes in material/geometric properties have on product reliability (Popelar, 1997; Michaelides and Sitaraman, 1998). Although such parametric studies provide critical data and insight on the effects of changes to design parameters such an approach can be very time consuming. Also, the

parametric study cannot investigate more than two parameters simultaneously, and very important interactions are very likely to be missed in such an analysis. Tools, based on numerical optimisation techniques, overcome the undesirable features of such parametric studies, giving a very logical, time efficient and automated approach in identifying the best model and/or process parameters for different design problems (Cimtalay et al., 1996; Li et al., 1996; Deshpande et al., 1997). Applying such optimisation strategies to microelectronics component design offers a fully automated, fast and systematic approach to achieve the most appropriate system characteristics. This paper demonstrates such a modelling framework that integrates

Figure 1 Traditional “flow” (a) and novel “no-flow” (b) underfill assembly process

The current issue and full text archive of this journal is available at http: //www.emeraldinsight.com /0954-0911.htm [ 49 ]

S. Stoyanov, C. Bailey and M. Cross Optimisation modelling for flip-chip solder joint reliability Soldering & Surface Mount Technology 14/1 [2002] 49–58

optimisation technology with thermo-mechanical fatigue modelling. This modelling methodology is demonstrated by using it to identify optimal design properties for flip-chip components that adopt no-flow underfill.

Figure 2 Top view of flip chip assembly and modelled part

FEM computational modeling Geometry and mesh model The flip-chip assembly consists of a silicon die attached to an FR-4 board and connected via solder joints. Underfill epoxy material fills the gap between the chip and substrate. Table I provides details on the dimensions of this package. The connection of the chip to the substrate is accomplished using 48 uniformly and peripherally arranged solder joints. The distance between any two adjacent solder joints, pitch size, is 457 mm, and the distance between the substrate and die (stand-off height) is 130 mm for the initial assembly. Figure 2 details the computer model used to represent this package. One-eight symmetry is assumed to lower the compute times. Once the geometry and all model components are defined, a mesh is generated and for the flip chip assembly under investigation the number of elements is 9212. These consist of linear wedges and bricks (Figure 3). All the solder interconnects inside the modelling one-eight part are taken into account and included into finite element model representation.

Figure 3 FE model – geometry and mesh

Material model Figure 4 details the materials present in the flip-chip package being investigated. Modelling the thermal fatigue failure of solder joints requires a constitutive law for the creep deformation of the solder. All of the other materials are assumed to be elastic. To ensure accurate material response in the analysis, all five materials used in the package are considered to have temperature dependent material properties similar to those used by Le Gall, Qu and McDowell (Le Gall et al., 1997) and Hong (Hong, 1998). With respect to coefficient of thermal expansion CTE, the organic composite materials (FR-4) made substrate is modelled as an orthotropic material. The detailed properties of all materials are summarised and reported in Table II. During temperature cycling the solder material experiences deformation due to time-dependent creep and stress relaxation. To model this deformation for eutectic solder, a sinh creep relationship, given in equation (1), has been adopted in the analysis:
 2Q 3 sij n ð1Þ 1_cr ij ¼ A½sinhðaseff Þ exp RT 2 seff

Figure 4 Assembly materials

where R is the gas constant, T is the temperature in Kelvin, sij is the deviatoric stress tensor, seff is the Von Mises stress, and all other symbols represent material properties. For 63Sn/37Pb solder alloy, they are: A ¼ 9:62 £ 104 s 21 , a ¼ 0:087 MPa 21 and Q ¼ 8058:37R (Darveaux and Banerji, 1995).

Table I Initial geometric parameters and variations

Geometry parameter Die dimensions Substrate thickness Stand-of height Pad Diameter Thickness Bump pitch size Number of bumps

[ 50 ]

Variations Range from Initial value

Initial value [mm]

Absolute range [mm]

6.3 £ 6.3 £ 0.6 1.5 130 mm

None 1.2– 1.8 104 mm – 156 mm

None ^ 20% ^ 20%

150 mm 35 mm 457 mm 48

None None None None

None None None None

S. Stoyanov, C. Bailey and M. Cross Optimisation modelling for flip-chip solder joint reliability Soldering & Surface Mount Technology 14/1 [2002] 49–58

Thermo-mechanical modelling procedure Thermo-mechanical simulations of flip chip assembly under thermal cycling predict system responses. Displacements, stress/strain and creep strains within the package are given by analysis using PHYSICA (PHYSICA). PHYSICA is a software environment for finite element modeling of multiphysics phenomena. This analysis tool is a collection of interacting solution procedures, including elasto-viscoplastic solid mechanics (Taylor et al., 1995). The incremental form of equilibrium equations is Dsij ¼ 0

i; j ¼ x; y; z

ð2Þ

stress in the die will be at its bottom surface, especially those areas at the center and corner of the die. The reason for this is that the low value of CTE for the die compared to the other materials being used in the assembly results in high compressive stresses at the die bottom side and tensile stresses at the die top surface. Figure 6 shows the die Von Mises stress for the initial design configuration reported in Tables I and II. The thermal solder fatigue failure is indicated by accumulated effective creep strain at the end of a thermal cycle (Figure 5, point b), 1creep sum :

In equation (2) sij are the components of Cauchy stress tensor, s. The incremental elastic strains, 1 el, relate to stresses via Ds ¼ ½DD1 el

1creep sum ¼

ð3Þ

where the elasticity matrix [D ] is defined in terms of the material properties, Young’s modulus and Poisson ratio. The elastic strain increment is given by D1 el ¼ D1 2 D1 th 2 D1 vp th

ð4Þ

vp

where D1, D1 and D1 are, respectively, the total, thermal and visco-plastic strain increments. In this paper the time dependent creep strain rate of eutectic solder is modeled using previously described sinh creep low (Equation 1), and the thermal strain rate is given by 1_th ij ¼ aCTE T_dij

ð5Þ

where aCTE is the linear coefficient of thermal expansion, T˙ is the rate of change the temperature, and dij is the Kronecker delta. Material properties, boundary conditions and temperature profile are the required analysis inputs. The thermal load that has been used consists of four stages: temperature decrease, dwell at low temperature extreme, temperature increase, and dwell at high temperature extreme. Each of these temperature phases is equal to 15 minutes and the temperature range is from 2 25 to 1258C (Figure 5). Time dependent non-linear stress/strain analyses has been undertaken to and evaluate the thermo-mechanical system response for a particular flip-chip design. A number of simplifications and assumptions are imposed. The stress free state for the entire assembly, including the die component, is at 1258C, which is the starting thermal cycle temperature. Modelling the temperature cycling test assumes isothermal loading throughout the package. In addition all initial stresses are neglected. Finally, perfect adhesion between all materials is considered. The potential for die cracking is estimated by monitoring the maximum levels of Von Mises stress in the die at the end of most critical cooling down stage in the thermal cycle (Figure 5, point a). It is known that the highest Von Mises

rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi X 2 creep creep d1sum d1sum 3 t

ð6Þ

where the sum is over all time steps completing a thermal cycle, and d1creep sum are the components of the incremental creep strain tensor during one time step. Since stress and strain are not uniformly distributed across the solder joints, the volume average calculated value of accumulated effective creep strain in solders over range of elements with maximum value of the quantity is used to define 1creep sum : Higher 1creep sum relates to lower reliability. In this study the Coffin-Manson relationship (Dieter, 1998) a N f ¼ uð1creep sum Þ

ð7Þ

is used to correlate 1creep sum to lifetime in terms of number of cycles to failure, Nf (reliability indicator). In equation (7), u and a are eutectic solder material related parameters, obtained experimentally ðu ¼ 0:38 and a ¼ 21:96Þ:

Figure 5 Thermal cycle specification

Table II Package material properties and variations Material Substrate: FR4 Card

Pad: Copper

Solder: Sn63/Pb37

Die: Silica

Underfill Underfill Variations

Temp. (8C)

CTE(1026/8C)

Young’s modulus E (GPa)

Poisson ratio n

30 95 125 225 0 50 125 0 50 100 0 50 100 – –

20.0 (xy), 86.5 (z) 20.0 (xy), 86.5 (z) 20.0 (xy), 86.5 (z) 16.0 16.0 16.0 16.0 25.2 26.1 27.3 2.8 2.8 2.8 80.0 50.0 – 90.0

22.4 20.68 19.3 128.0 127.0 124.88 121.45 26.45 12.52 6.91 162.1 150.0 140.0 3.0 2.0– 5.0

0.1425 0.1425 0.1425 0.34 0.34 0.34 0.34 0.360 0.365 0.378 0.23 0.23 0.23 0.35 None

[ 51 ]

S. Stoyanov, C. Bailey and M. Cross Optimisation modelling for flip-chip solder joint reliability

Figure 6 Die Von Mises stress distribution and deformation (initial design)

Soldering & Surface Mount Technology 14/1 [2002] 49–58

.

.

In the process of studying different assemblies it was detected that corner solder joint is always the one with highest accumulated effective creep strain. Figure 7 illustrates the creep strain distribution across corner solder joint for the initial assembly.

“Simulation-optimisation” methodology and strategy At present, finite element tools are used extensively to simulate and predict behaviour and different responses of systems, processes or phenomena using reliable material data and more precise modelling of the physics. Coupled numerical optimisation techniques and finite element analysis offers a design tool for virtual prototyping of best model/process parameters with respect to a specified design criteria. This strategy was implemented for designing reliability in electronic packages (Stoyanov et al., 2001). The finite element code, PHYSICA, is linked with the optimisation tool, VisualDOC (Vanderplaats, 2001). VisualDOC is a collection of software components provided with interfaces for coupling to simulation tools. Different numerical optimisation techniques and approaches are available inside the optimizer. The optimisation problem is specified in VisualDOC. To define the optimisation task in mathematical form, design variables, objective and possible constraints should be specified. These are: . Design variables are model/process parameters (i.e. underfill material properties) subject to change. Optimisation identifies the best values giving an optimal improvement in specified design aspect/ characteristics (represent by objective function).

Figure 7 Accumulated effective creep strain levels in corner joint (initial design)

Objective function represents one or more model features that we wish to optimize (e.g. maximize solder lifetime, or minimize creep strain, etc.). The objective is normally a function (in explicit or implicit form) of the design variables and system responses. Constraints are requirements (restrictions) imposed on the design that have to be satisfied. Similar to the objective function, constraints are also functions of the design variables and/or system responses. Lower and upper bounds on design variables are special types of constraints (e.g. if solder stand-off height is a design variable, it could not be negative, and must be restricted by a sensible upper limit).

VisualDOC interacts in iterative manner with PHYSICA using data files. To perform the optimal design procedure, the optimizer needs system responses for different sets of design variables that optimisation technique requires. Design variables are either geometry parameters (that reflect on geometry and mesh of finite element model) or analysis input data parameters (e.g. material properties, process conditions). In this paper only two optimal design strategies will be tested – Direct first-order search techniques based on gradient calculations (Vanderplaats, 1999) and Design of Experiments (DOE) combined with Response Surface (RS) approximations (Vanderplaats, 2001). These approaches use finite element system responses in different way. The first strategy, Direct Optimization, is a first order (gradient) search approach and it evaluates the objective function and constraints for a design based on system responses from analysis tool (PHYSICA ). Numerical optimisation search techniques rely on iterative improvement of design, starting with the initial and converge to the optimal design. If currently evaluated design is not optimal, then the numerical optimization search technique calculates new one. This new set of design variables is passed to the analysis program and simulation is performed for that design to obtain the corresponding system responses. In the second tested approach, based on DOE, analysis program is run for a set of different experimental designs at the beginning. At this initial stage analysis program PHYSICA is supplied by optimization tool with designs (sets of design variables) that have to be evaluated through finite element simulations and returns to the optimizer the corresponding system response data. This available response data is used then to construct approximations (Response Surfaces, RS) in explicit mathematical form for each of the responses being considered in optimization problem. The optimization procedure itself uses again numerical technique to solve the optimization problem, but no more analyses needs to be executed - all objective and constraint function evaluations for different designs in the iterative optimization process are based on the generated RS approximations. The optimal design procedure inside “PHYSICA-VisualDOC” design framework and interaction between these analysis and optimisation tools is given in Figure 8.

Simulations and results Previous studies into flip-chip assemblies have identified the importance of underfill material properties and some geometry parameters, such as substrate thickness, on solder joint fatigue reliability. With respect to “no-flow” underfilled assemblies, the underfill Young’s modulus (E ) and its coefficient of thermal expansion (CTE ) have been experimentally investigated (Wong et al., 1998) and it was observed that they had a major impact on reliability. In this modelling study we will investigate not only material properties of the underfill but also the geometric design parameters to predict their effect on reliability. 1 Young’s Modulus of underfill (E, GPa ) 2 Coefficient of Thermal Expansion of underfill (CTE, ppm ) [ 52 ]

S. Stoyanov, C. Bailey and M. Cross Optimisation modelling for flip-chip solder joint reliability Soldering & Surface Mount Technology 14/1 [2002] 49–58

Figure 8 Optimal design procedure

3 Stand Off Height (SOH, mm ) 4 Substrate Thickness (ST, mm ) The objective in the optimisation study is to maximize solder joint reliability in terms of number of cycles to failure, Nf. Within the optimisation procedure the above parameters will be allowed to vary and the limits on these design optimisation variables under investigation are given in Tables I and II. Another constraint used in this study is the maximum Von Mises stress allowed in the die, s¯, at the end of the cooling down period in the thermal load. Therefore the results from this study will ensure that another type of failure – die cracking – is avoided. The upper allowable limit value for Von Mises stress in the die, s max, is taken to be 130% of the maximum value obtained for the initial flipchip design. This value is entirely acceptable since it is significantly less than the allowable Von-Mises stress that may cause die cracking (Dieter, 1998). The constrained optimisation task can be expressed in the following form:

Find values for the design variables (E, CTE, SOH, ST ) that will a Maximize Lifetime, N f ¼ uð1creep sum Þ Subject to: s # s max , where s max ¼ 155:5 MPa and lower and upper limits on E, CTE, SOH, ST, as given in Tables I and II. For this study both the strategy of Direct Optimisation using finite difference gradient calculations and Design of Experiments (DOE) will be demonstrated.

Direct optimisation Direct Optimisation is based on gradient search techniques that are capable of finding the exact optimum design. This approach involves extensive calls to an analysis code, such as PHYSICA, to obtain system responses. Finite difference techniques are used to calculate gradients of the objective and constraints. These gradients are then used by to specify search directions in design space. Other calls to analysis programme are used to evaluate objective and constraints in the optimisation search process when the search direction is already available. A major drawback to Direct Optimisation strategy is that it will require generally many analysis calculations when there is a large number of design variables. In the case of time-consuming analysis this will slow down dramatically the optimal design process. Using this direct approach to optimisation the aim is to maximise the objective function (i.e. solder lifetime). Objective improvement in the optimisation process is illustrated in Figure 9. Starting with a lifetime of 136 cycles to failure for the initial design, an improvement to 589 cycles to failure is achieved by identifying the optimal values of the design variables (E, CTE, SOH, and ST ). To achieve these optimal values optimiser has made 15 analysis calls to the finite element programme PHYSICA. The changes in each design variable throughout the process of optimal design improvement are given in Figure 10 showing a decrease in CTE and substrate and an increase in Young’s modulus and stand-off height from the values at the initial design. Increased lifetime for the optimal design configuration is a result of significantly reduced levels of accumulated effective creep strain in solders at the end of a thermal cycle. Figure 11 illustrates the history of creep strain development across the most critical corner solder joint for both the initial and optimal design throughout a full thermal cycle.

Design of experiments (DOE) and response surface modelling DOE is exploited within the sensitivity and/or optimisation analysis by: 1 Identifying the experimental design points. 2 Undertaking a finite element analysis at each design point to obtain the system responses at these points 3 Using the response data to construct response surface (RS) approximations. 4 Perform sensitivity and optimisation analysis using the generated response surfaces. The starting point in a DOE analysis is to establish the set of experimental design points at which the finite element analysis will be undertaken to provide system response data (i.e. stress and strain values). In this study the Three-Level Taguchi Orthogonal Array L9 and the Factorial Notz design techniques are used (Vanderplaats, 2001) and these result in 19 design points. These two experimental design techniques are suitable to construct accurate full quadratic approximations to the system response data using a relative low number of design points. The central location of the design space is also added to the design points in the final set of design experiments. This extra point provides greater [ 53 ]

S. Stoyanov, C. Bailey and M. Cross Optimisation modelling for flip-chip solder joint reliability

Figure 9 Objective optimisation history

Figure 11 Maximum accumulated effective creep strain as a function of time during a complete thermal cycle

Soldering & Surface Mount Technology 14/1 [2002] 49–58

accuracy in the second order polynomial approximation to the system responses (i.e. the response surface, RS). Once the design points have been identified, the next step is to calculate the system responses at these points. This is achieved by running the finite element analysis and calculating the Von Mises stress in the die at the end of cooling down phase (constraint value) and the maximum accumulated effective creep strain in the solder at the end of thermal cycle (used to identify directly objective, Nf). Table III shows the results from this DOE study. After the system responses have been obtained at the experimental design points, the next stage is to construct a second order polynomial approximation to each of these responses. These responses are die Von Mises stress and accumulated effective creep strain in solder. The coefficients in both RS approximations are determined by fitting the response data using the least squares approach. If these RS approximations represent the responses accurately, then they can be used within the optimisation procedure to evaluate the objective function (Nf) and the constraint (s¯ ) thus avoiding repeated calls to the time consuming finite element analysis programme. With respect to maximum

Figure 10 Design variables optimisation history

[ 54 ]

accumulated effective creep strain in solders system response, the following full quadratic approximation is delivered: 1creep sum ðEðGPaÞ; CTEðppmÞ; SOHðmmÞ; STðmmÞÞ ¼ 3:267 2 1:013*E þ 0:063*CTE 2 0:027*SOH þ 0:923*ST þ 0:00008*E*CTE þ 0:003*E*SOH 2 0:108*E*ST 2 0:00003*CTE*SOH 2 0:0002*CTE*ST 2 0:004*SOH*ST þ 0:064*E ** 2 þ 0:00004*CTE ** 2 þ 0:00005*SOH ** 2 þ 0:185*ST ** 2

S. Stoyanov, C. Bailey and M. Cross Optimisation modelling for flip-chip solder joint reliability

Table III Design of experiments

Soldering & Surface Mount Technology 14/1 [2002] 49–58 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20

E (GPa)

CTE (ppm)

SOH (mm)

ST (mm)

Maximum die Von Mises stress (MPa)

Max. accumulated effective creep solder strain (per cent)

Cycles to failure: lifetime Nf

5 5 5 3.5 3.5 3.5 2 2 2 2 2 2 5 5 5 5 3.5 3.5 3.5 3.5

90 70 50 90 70 50 90 70 50 50 90 90 90 90 50 50 50 70 70 70

130 104 156 104 156 130 156 130 104 156 156 104 156 104 156 104 130 104 130 130

1.2 1.8 1.5 1.5 1.2 1.8 1.8 1.5 1.2 1.8 1.2 1.8 1.8 1.2 1.2 1.8 1.5 1.5 1.2 1.5

147.89155 127.05443 125.33030 128.68752 129.30603 113.37784 97.64297 108.06065 120.97282 96.85377 113.63030 106.59825 130.82172 148.09805 133.41422 122.07200 119.59298 125.39785 131.55486 122.82088

5.0527405 4.1510113 2.4542704 5.7681266 3.8058821 2.9955601 5.8739781 4.7173206 3.6309626 3.3120830 5.5149138 6.7686215 5.1244829 5.2085061 2.3565151 2.8720692 2.8529394 4.4147812 4.0000290 4.1535586

132 194 544 102 230 368 98 151 252 302 111 74 128 124 589 400 405 172 209 194

Similarly, an explicit full quadratic polynomial approximates the other investigated response, the maximum Von Mises stress in the die, s¯. Different statistical tools are available to judge the quality of the generated RS approximations. These are based on Residual Analysis, Analysis of Variance, and statistical efficiency measures (Vanderplaats, 2001). These techniques have shown that the coefficients predicted for each RS polynomial provide a good fit and are accurate for both die stress and creep strain responses. In Figure 12, the simulated DOE responses using finite element analysis are compared with those obtained using the RS polynomials. Clearly we can see that the polynomials provide acceptable accuracy to the response values at the design points. Apart from the benefits of avoiding actual analysis runs in the optimisation process, Response Surface Modeling has an additional advantage as it can be used to perform sensitivity analysis (something which is not explicitly available in the direct optimisation approach). This sensitivity analysis helps identify which design variables and their interactions are important with regards the system responses. These sensitivities are obtained from the coefficients in the RS approximation. To avoid the influence of different design variable units and the different lengths of design boundaries (i.e. Young’s modulus (E ): 3 to 5 GPa and CTE: 50 to 90 ppm), the sensitivity data is based on the

RS approximation where design space is normalised. In this normalised In this normalised design space lower design variable bounds correspond to 2 1, and the upper bounds to +1. As a result, one unit of change in normalised design space for E, CTE, SOH and ST corresponds, respectively, to change of 1.5 GPa, 20 ppm, 26 mm and 0.3 mm in their values. Figure 13 illustrates the sensitivity of creep strain response in the solder to the six most important terms in the RS approximation. These are represented by the coefficients in the RS approximation, generated in the described above normalised design space. For example, a positive unit change in the Young’s modulus (i.e. 1.5 GPa) of the underfill will result in a decrease of 0.445 in the creep through the RS approximation linear term (E ) for the underfill Young’s modulus. But at the same time the quadratic term for underfill E (E **2) will increase the creep strain by the 0.143. This sensitivity analysis also shows that the underfill CTE is most influential on changes in creep strain in the solder. This is followed by the underfill Young’s modulus E, then solder stand-off height SOH then substrate thickness ST and then the other terms in the response surface approximation. DOE based optimisation uses the constructed RS approximations explicitly to evaluate the constraint and objective values without running a finite element analysis. Different direct gradient based optimisation techniques,

Figure 12 Actual vs. predicted responses

[ 55 ]

S. Stoyanov, C. Bailey and M. Cross Optimisation modelling for flip-chip solder joint reliability

Figure 13 The influences of most important design variables and their interactions on the accumulated effective creep strain

Soldering & Surface Mount Technology 14/1 [2002] 49–58

such as the Modified Feasible Directions Method or Sequential Linear/Quadratic Programming (Vanderplaats, 1999) can then be used to obtain the optimal solution. Figure 14 illustrates the lifetime number of cycles to failure based on the creep strain RS. The first graph relates lifetime to both E and CTE and is generated by keeping the other two design variables, SOH and ST, at their optimal Figure 14 Lifetime as a function of design variables

[ 56 ]

values. Clearly we can see that taking underfill CTE towards its lower bound and Young’s modulus (E ) towards its upper bound results in an increase in lifetime. In the second graph CTE and E are kept at their optimal values and SOH and ST are allowed to vary. In this case we can see that decreasing substrate thickness (ST ) and increasing stand-off-height (SOH ) results in an increase in lifetime. Figure 15 shows

S. Stoyanov, C. Bailey and M. Cross Optimisation modelling for flip-chip solder joint reliability

is noted as the “constraint” in this table and the number of cycles to failure (lifetime) as the “objective”. For the observed design space, both tested optimisation strategies (Direct and DOE) locate the same optimal design: . Best material properties: underfill Young’s modulus E at its upper imposed bound (5GPa) and underfill CTE at its lower bound (50 ppm); . Best geometric parameters: stand-off height SOH at its upper bound (156 mm), and substrate thickness ST at its lower bound (1.2 mm).

Figure 15 Maximum die Von Mises stress as a function of design variables

Soldering & Surface Mount Technology 14/1 [2002] 49–58

similar graphs for maximum Von-Mises stress in the die where generated RS approximation for this response is used.

Comparison between direct optimisation and DOE The summarised data is given in Table IV. Maximum Von Misses stress in the die at the end of cooling down phase

In this table the objective, constraint and accumulated creep strain values given for the initial design are directly calculated using finite element analysis. For the optimal design the four optimal design variables are calculated using both optimisation approaches detailed above. Also, these two optimisation approaches are used to predict the objective, constraint and accumulated creep strain values for the calculated optimal design. Analysis based evaluations for both the initial and optimal designs show that a significant improvement in solder joint reliability is obtained. The calculated optimal design using direct optimisation suggests thermal fatigue reliability in terms of cycles to failure is 3.33 times higher than that estimated for the initial design, where the initial design gives 136 cycles to failure and the optimal design results in 589 cycles to failure. This optimal design satisfies all the imposed constraints. An acceptable increment of 11.53 per cent in maximum Von Mises stress in the die is observed when the optimal and initial designs are compared. It should be noted that the response values obtained using the direct optimisation approach are calculated using finite element analysis. Clearly we can see that the Design of Experiments Response Surface approach (DOE RS) locates the same optimal design in this study. Using the response surface approximation, as detailed above, we can see that the predicted values of the objective, constraint and accumulated creep strain are very close to those predicted using the Direct approach, above. The variation between the finite element (or direct optimisation) predictions for the objective function and the approximation (or DOE RS) is only 1.7 per cent. Since the RS approximation for both creep strain and die stress, 1creep sum and s¯, have been generated accurately, at optimal design the differences in their exact and approximate values are only 0.89 and 0.22 per cent, respectively. It should be noted that for this particular design problem, Direct Optimisation was able to find the optimal solution with relatively low number analysis calls (15 PHYSICA simulations compared with 20 calls for DOE). Although in this case the DOE analysis required more calls to the analysis code, generally, this will not be the case for most design problems where the optimum design point may lie

Table IV Summarised optimisation results

Design parameters

Lower bound

Upper bound

INITIAL design

1. E (GPa) 2.0 5.0 3.0 2. CTE (ppm) 50.0 90.0 80.0 3. Stand-off height (mm) 104.0 156.0 130.0 4. Substrate thickness (mm) 1.2 1.8 1.5 Constraint and objective values Constraint (die stress) 119.6173 Objective (cycles to failure) 136 Accumulated effective creep solder strain (per cent) 4.969925 Difference between FEM values and optimisation approach values at optimal design Constraint (die stress) Objective (cycles to failure) Accumulated effective creep solder strain (per cent) Number of FEM analyses for obtaining the solution

OPTIMAL design Direct DOE RS 5.0 50.0 156.0 1.2

5.0 50.0 156.0 1.2

133.414 589 2.356515

133.709 579 2.377451

0% 0% 0% 15

0.22% 1.70% 0.89% 20

[ 57 ]

S. Stoyanov, C. Bailey and M. Cross Optimisation modelling for flip-chip solder joint reliability Soldering & Surface Mount Technology 14/1 [2002] 49–58

inside the design space, not on the design space boundaries. Usually a significant number of simulations will be required to perform the gradient-based optimisation search process and therefore the preferred option is the DOE method with predictions for the responses provided by RS approximations. In addition, the DOE approach offers the opportunity to perform sensitivity analysis and subsequent estimation of the importance and influence of the design variables.

Conclusions Thermo-mechanical analysis to predict solder joint reliability has been coupled with optimisation techniques and demonstrated for a no-flow underfilled flip chip assembly. This integrated methodology can be used to identify optimal conditions for solder joint reliability. Solder joint lifetime has been improved dramatically in a very time efficient, intelligent and automated manner. Different optimisation strategies combined with response modelling and approximation techniques have been tested as part of the integrated design framework. It is shown that higher underfill Young’s modulus E and stand-off height combined with lower underfill CTE and substrate thickness have a positive effect on the thermo-mechanical reliability. In this analysis the variations of both geometric and material parameters are taken as continuous. In the manufacturing process, a suitable underfill can be chosen with the properties that are closest to the estimated optimal (taking into account all aspects of the sensitivity analysis). Results from this analysis have investigated both underfill material properties and geometric parameters. Experimental studies (Wong et al., 1998; Shi et al., 2000) investigated the influence of underfill material properties: different values of CTE and E. In this modelling study these have been taken as the bounds on allowable variations. Since in this paper we have demonstrated that the influence on solder reliability is dominated by these material properties rather than flip-chip geometry design we can compare our results with this experimental study. The results for reliability estimation and trends observed are comparable to published experimental results for no-flow underfills (Shi et al., 2000). This methodology will be extended in future to investigate different stages in the assembly process where finite-element modelling coupled with optimisation techniques can further help to identify key process conditions. This will require multi-physics modelling (e.g. coupled solidification and stress analysis) to simulate solder behaviour during the reflow process. Further studies on the reliability of assembled flip-chip packages in terms of other design parameters and different type of failures is also intended.

[ 58 ]

References Cimtalay, S., Peak, R.S. and Fulton, E.R. (1996), “Optimization of solder joint fatigue life using product model-based analysis models”, ASME International Mechanics Engineering Congress and Exposition, EEP, Vol. 18, pp. 47-53. Dai, X. and Ho, P.S. (1997), “Thermo-mechanical deformation of underfilled flip-chip packaging”, IEEE/CPMT International Electronics Manufacturing Technology Symposium, pp. 326-33. Darveaux, R. and Banerji, K (1995), Ball Grid Array Technology, McGraw-Hill, New York, NY, pp. 379-442. Deshpande, A.M., Subbarayan, G. and Mahajan, R.L. (1997), “Maximizing solder joint reliability through optimal shape design”, Journal of Electronic Packaging, Transactions of the ASME, Vol. 119 No. 3, pp. 149-55. Dieter, G.E. (1998), Mechanical Metallurgy, McGraw-Hill Book Company, New York, NY, pp. 241-46. Hong, B.Z. (1998), Thermal fatigue analysis of a CBGA package with lead-free solder fillets, Proceedings of InterSociety Conference on Thermal Phenomena, pp. 205-11. Lau, H.J. (1996), Flip Chip Technologies, McGraw-Hill, New York, NY. Le Gall, C.A., Qu, J. and McDowell, D.L. (1997), “Influence of die size on the magnitude of thermo-mechanical stresses in flip chips directly attached to printed wiring board”, ASME Advances in Electronic Packaging, Vol. EEP-Vol 19 No. 2, pp. 1663-70. Li, Y., Mahajan, R.L. and Nikmanesh, N. (1996), “Fine pitch stencil printing process modeling and optimization”, Journal of Electronic Packaging, Transactions of the ASME, Vol. 118 No. 1, pp. 1-6. Michaelides, S. and Sitaraman, S.K. (1998), “Effect of material and geometry parameters on the thermo-mechanical reliability of flip-chip assemblies”, Proceedings of InterSociety Conference on Thermal Phenomena, pp. 193-200. PHYSICA, http://www.gre.ac.uk/~physica. Popelar, S.F. (1997), “Parametric study of flip chip reliability based on solder fatigue modelling”, Proceedings of the IEEE/CPMT International Electronics Manufacturing Technology (IEMT) Symposium, pp. 299-307. Shi, S.H. and Wong, C.P. (1999), “Recent advances in the development of no-flow underfill encapsulants – a practical approach towards the actual manufacturing application”, IEEE Transactions on Electronic Packaging manufacturing, Vol. 22 No. 4, pp. 331-9. Shi, S.H., Yao, Q., Qu, J. and Wong, C.P. (2000), “Study on the correlation of flip chip reliability with mechanical properties of no-flow underfill materials”, Proceedings of International Symposium on Advanced Packaging Materials, pp. 271-7. Stoyanov, S., Bailey, C. and Cross, M. (2001), “Integrating computational mechanics and numerical optimization for the design of material properties in electronic packages”, TMS Proceedings of Computational Modeling of Materials, Minerals and Metals Processing, September 23– 26, San Diego, CA, pp. 551-61. Taylor, G., Bailey, C. and Cross, M. (1995) “Solution of the elastic/visco-plastic constitutive equations: a finite volume approach”, Appl. Math. Modelling, Vol. 19, pp. 746-60. Vanderplaats, G.N. (1999), Numerical optimisation techniques for engineering design: with applications, VR&D, Colorado Springs. Vanderplaats, G.N. (2001), VisualDOC 2.1: Theoretical Manual, http://www.vrand.com. Wong, C.P., Shi, S.H. and Jefferson, G. (1998) “High performance no-flow underfills for low-cost flip-chip applications: Material characterization”, IEEE Transactions on Components, Packaging and Manufacturing Technology, Part A, Vol. 21 No. 3, pp. 450-8.

Internet commentary In language, the ignorant have prescribed laws to the learned. (Richard Duppa, Maxims, 1830, No. 252) In this prologue, I shall be slightly more technical than usual. I wish to discuss mark-up languages used for generating Internet pages on the World Wide Web. The basic language which has been used up to now is called HTML or hypertext markup language. All the WYSIWYG translators change what the designer sees on the screen into HTML. The more savvy website designers may write a whole page or part of it directly in HTML, using an Ascii editor. But what is HTML? HTML is a subset of SGML or Standard Generalised Mark-up Language. This is a very complex mark-up language with very strict rules and is difficult to use. The subset was created to make a simple, easy-to-use, language specifically for Internet applications. It was originally intended, in version 1, to handle just text, without any possibility of making a Web page attractive to the eye or to include graphics. As the requirements increased, so did the complexity of HTML. Even this was not enough, so that it became necessary to add scripts, JavaScript being the most popular. These allowed simple programmes to be executed within a Web page. Unfortunately, the flexibility that these offered has allowed poor programmers to produce bloatware pages and increased all sorts of security issues. The basic concept of HTML is to surround each item in a page by two tags. For example, the tags at the beginning and end of a paragraph are , p . and , /p . respectively. However, lazy programmers (although not strictly accurate, I shall use this term to describe authors who write in HTML) soon found that some end tags are not necessary in the majority of browsers and the latter were increasingly tolerant of such poor programming, as a result. In other words, your Microsoft Internet Explorer or Netscape Navigator allowed countless sins. The number of permitted tags in the later versions of HTML run into the hundreds, although most programmers may use only 20 or 30 on a regular basis. Over the past couple of years, a new buzzword has appeared, XML or extensible mark-up language. The latest word processors all include means of writing in this language, even though very few persons are aware of what it is about or what it does. Even though it offers many advantages for Web page design, most programmers have found that it is far too difficult to understand and they simply do not have time to get down to the nitty gritty of it. The key advantage of XML over HTML, is that a programmer may create his own tags. As an example, if you wish to incorporate a database of fluxes in a web page, it may simplify your task by creating such tags as , flux_name . , , /flux_name . , , flux type . and so on. A single entry on a page may then appear as follows: , flux_database . , entry . , flux_name . Smith’s X123 , /flux_name . , flux_type . water-soluble , /flux_type . , flux_activity . high , /flux_activity .

, flux_cleaner . water , /flux_cleaner . , /entry . , /flux_database . I don’t think that I shall need to elaborate to show, from this simple example, that it is very much more powerful than HTML, even if it is so much more difficult to use. So the question is, should webmasters make the effort to write their pages in XML? The simple answer is No. This is because the current browsers are not able to take advantage of all that XML offers. Probably, in my opinion, the situation will not improve over the next few years, until all browsers are XML-compliant. One can therefore legitimately ask oneself what all the fuss is about. The answer is that there is a compromise available, called XHTML. This is beginning to come to life, since version 1.1 was published earlier this year. Programmers can easily make the transition from HTML to XHTML without busting their brain boxes. However, there are a number of things which must be remembered. Perhaps one of the most important is that the strict protocol is necessary – no more sloppily omitting end tags etc. Browsers can generally handle this easily, without modification. So what are the disadvantages? Generally speaking, HTML documents begin with something like: , !DOCTYPE HTML PUBLIC "-//IETF// DTD HTML//EN" . , html . This is to tell the browser what it will be contending with, the minimum being the second line above. Because XHTML is really a modified XML, restricted essentially to HTML tags, its starting lines are somewhat different: , ?xml version="1.1" standalone="yes"? . , !DOCTYPE html public "-//W3C//DTD XHTML 1.1 Transitional//EN" "http://www.w3.org/TR/xhtml1l/DTD/transitional.dtd" . , html xmlns="http://www.w3.org/2001/ xhtml" . There are two important things which spring to the eye. The most obvious one is the references to the URL http://www.w3.org, which refers to the World Wide Web Consortium (W3C), in turn the authoritative body recommending the use of XHTML and writing its rules. The other one is the increased reference to DTDs and the specific transitional one. A DTD is a Document Type Definition and, as can be seen from the first example, is not new but has been largely ignored. In XHTML, it is mandatory to use one of three DTDs to validate the document. These are: Strict: The W3C intends to phase out some of the weaker or superfluous HTML tags from XHTML as soon as would appear practical. A document type defined as Strict means that it will conform to future versions of XHTML and does not use any of the legacy tags likely to be eliminated. Transitional: This defines the document as containing legacy HTML tags. Most early use of XHTML, where current HTML pages are translated to XHTML, will use this DTD, for safety. Frameset: This is a slight departure from HTML. In HTML, a page which defines the size and contents of frames is defined uniquely as being between the tags , frameset . and , /frameset . instead of , body . and , /body . , as in other pages. In XHTML, this is the same but, in addition, the whole document type must be defined accordingly. I

suspect that this is because frames are also transitional (and a good thing, too: they have really outlived their usefulness). I mentioned before that XHTML will not tolerate sloppy syntax or programming, such as the omission of end tags. There are some other rules that change slightly. In HTML, the tag for a horizontal line is , hr . but it strictly should be , hr . , /hr . , although the second element was invariably omitted because there was nothing to go between the two tags and it did nothing, in any case. In XHTML, this is called an “empty” tag, because it did not have any attribute or text. Empty tags are written in XHTML by, for example, , /hr . , so that it is still permissible to use them without writing two whole tags. Tag names and their attributes and values must, in XHTML, be written exclusively in lowercase. It is no longer allowed to use, for example, ,p align="CENTER" . . This makes for a more consistent appearance and simplifies browser design. Also, all the attribute values must now be between inverted commas (quotation marks). The above gives a very short introduction into the intricacies of XML and XHTML. Of course, I don’t expect anybody to start using XHTML from what I have said. For further details, you can refer to the URL of the W3C website, referred to above. The object of this essay is to allow webmasters to decide whether to expand their knowledge at this stage or to wait until the maturity is such that it becomes more interesting. Just to give you a hint, my personal attitude will be to continue to use HTML for a limited time, while attempting not to use the tags which I have described as legacy in the Strict DTD, so that a transition to XHTML will be less painful when the time comes. It should perhaps be mentioned that most of the popular WYSIWYG HTML code generators have not yet made the transition. As for XML, I am not even going to bother to look at it, at least for the time being. For the Review section of this Commentary, I shall revisit the lead-free debate. For those of you who subscribe to the IPC lead-free net list, you will realise that this subject is becoming more and more emotional rather than technical. I’ll be using the same technique as I did last time, by just reviewing pages rather than websites. My Google search words are “lead free solder OR solders OR soldering”.

http://www.leadfree.org Almost inevitably, the first reference to come up is the Home Page of the ITRI’s (formerly known as the International Tin Research Institute) subsidiary site on lead-free soldering. The page itself is uninteresting but it has a menu leading to many other, much more useful, pages. The site is worth exploring but it is a bit of a sheep in wolf’s clothing because, despite the.ORG domain, it is undoubtedly commercial, promoting the sale of tin.

http://www.solderworld.com/ If you don’t believe that the last reference has commercial overtones, you must visit this site which is sponsored by the same organisation. Again, this opens into a Home Page which, through a menu, can lead you to lists of suppliers of products for lead-free soldering. Interestingly, in the case of products containing alloys, the indexing is done according to the alloy itself. Actually, I believe that this is a very useful directory. [ 59 ]

http://www.elvesofester.com/ leadfree.html This is a very interesting reference, although it has nothing whatsoever to do with the electronics industry. It shows that the artisans making stained glass windows have exactly the same problems as we do in our industry. This page is very well worth reading if only because it shows that other sectors have come to similar conclusions as we have.

http://www.ivf.se/elektronik/Ep/ Publications/ LeadfreeSoSanAbstr.htm This page offers an abstract of a paper (for sale) published by the famous Swedish organisation, IVF, describing the addition of a small quantity of zirconium to lead free alloys, in order to reduce the grain sizes. At the same time, it would seem that the hardness of the alloy also increases. Before the question arises, the incorporation of 0.2 per cent of zirconium will not significantly increase the price of the alloy in itself, but the method of incorporation (cold attrition milling) may do so. An interesting idea but too early to state whether it has any commercial value.

http://www.boulder.nist.gov/div853/ lead%20free/props01.html The US National Institute of Standards and Technology and the Colorado School of Mines have combined to produce a very useful database. This collects together in one document nearly 50 tables of properties of lead free solders of many different alloys. For anybody wanting reliable data on an alloy, then this is the place to go to.

http://www.tms.org/pubs/journals/ JOM/0106/Frear0106.html This is a full article entitled “Pb-Free Solders for Flip-Chip Interconnects”, re-produced from the Journal of Manufacturing, complete with all the references and figures. It is long to download. Nothing further need be said because the title is selfexplanatory.

http://www.alphametals.com/ lead_free Of course, the solder manufacturers also have their websites. This one, from Alpha Metals, appears to have quite a lot of information accessible from this page. Unfortunately, I could not investigate this in detail because the Alpha server was either functioning very poorly or had an inadequate connection to the Internet backbone, at the time of writing this article.

http://www.npl.co.uk/npl/cmmt/ programmes/mpm/lfs.html The UK National Physical Laboratory has been a leader in the move towards lead-free solders and has done valuable work towards this end. However, I find a very telling sentence on this page, “Legislation on the use of lead, along with commercial pressures, will force the industry to use lead-free solder within a few years.”. You will note that there is no mention of scientific proof that lead in solder is harmful to humans or to the environment. It is known that the European legislation is not based on a scientific risk assessment and has been pushed through on false premisses. Notwithstanding, useful information is obtainable here for the inevitable. [ 60 ]

http://www.jgpp.com/projects/ lead_free_soldering/lead_free_ soldering.html This page appears to be a description of a project to demonstrate and validate lead-free solders, run by an organisation, the Joint Group on Pollution Prevention. The page itself and its sister pages seem to be rather weak on what they are trying to achieve. However, they appear to hold regular meetings and this may be of interest to our American friends.

http://it.jeita.or.jp/jhistory/english/ information/pbfree/roadmap2.html It is well known than the Japanese have been pioneering the use of lead-free solder in some lowcost consumer goods. The motivation behind this has been purely commercial, in an attempt to show consumers that the companies “care” for the environment. It would appear, from the latest information that, although it was successful initially, the move has backfired on the manufacturers. This page, published by the Japan Electronic Industry Development Association, is a very good question and answer document, entitled “Challenges and Efforts towards Commercialisation of Lead-free Solder”.

http://www.btu.com/html/Solutions/ ARTICLES/leadfree.htm This would appear to be an article “reprinted” from Circuits Assembly, last year. It is entitled, “Optimising Solder Joint Quality - Lead-free”. It describes a designed experiment, using two alloys and three flux types. A statistical analysis of the results is given. Unfortunately, the reader cannot benefit directly from this as the flux types were not revealed but contact with the authors may possibly remedy this situation.

http://www.kester.com/leadfree/ leadfree_update.htm Obviously, Kester Solders have an interest in the lead-free debate. This page is somewhat unexpected from a solder manufacturer who may be expected to push towards the higher cost lead-free solders, impartially weighing up some of the pros and cons. The title of the page is, “A Concise Snapshot of the Lead-Free Soldering Issue Today”. What is good is that the company still considers there is an issue and not just a fait accompli, as many others have done. In fact, this is the first page of my review with sound references to papers presented by opponents to lead-free soldering on the grounds of the environment. Yes, there is good reason to believe that the environment will suffer from the introduction of lead-free soldering.

http://epims.gsfc.nasa.gov/ctre/act/ techdocs/solder/leadfree.htm This is another paper written by an expert from the Goddard Space Flight Center. It is best summarised by the last part of the abstract, “The purpose of this report is to provide background information on environmental concerns of lead-based solders, pros and cons of continuing to use tin/lead solders, properties and preliminary technical assessment data from potential alternative lead-free solder alloys.”. It is a lengthy document discussing various alloys, including some quite exotic ones.

http://www.ate.agilent.com/emt/ industry/leadfreesolder/index.shtml To remind you, Agilent is an offshoot company from Hewlett-Packard, incorporating that part of the company not involved with IT equipment. This page is interesting in that it discusses the X-ray inspectability of lead-free solder joints, which are obviously less opaque than those made with a tinlead alloy. This is a matter which is often ignored but is very important to people using BGA and, above all, micro-BGA components. There is a full PDF paper downloadable from this page, as well as an abstract describing the problem.

http://www.wartonmetals.co.uk/ lead_free_reports.htm This page is a summary of much of the data available in the UK on the subject of lead-free soldering. It is edited by a manufacturer of such solder. There is nothing outstanding in the page, other than quotations from many sources. As such, it may be considered as useful, provided that one can support the flashing link to the commercial pages (I would advise those persons suffering from epilepsy and who are sensitive to flashing lights to avoid this page).

http://www.ti.com/sc/docs/ products/leadfree For our final word, let us enter into the world of components. Texas Instruments has been famous for supplying integrated circuit packages with a nickel/palladium finish on the solderable leads for many years. It has not been without problems, of course, but it is lead-free and should present no greater difficulties using these alloys than with lead-containing solder. This would appear to be a fairly formidable range of valuable pages to visit. However this is only a paltry fraction of those that came up in the search engine (about 67,000 results). On the whole, it would appear that there is an enormous amount of technical information available on the Internet on the subject, of varying degrees of quality. The one thing that has become apparent in the hundred or so pages that I looked at to write this article and that is the paucity of scientific information on the hazards that lead in solder may present. This tends to confirm my opinion that the whole movement towards lead-free soldering in electronics is based on the dirty buzzword “lead”, rather than a reasoned and reasonable action. As I said at the beginning of this section, this is a highly emotional subject, as well as having severe negative impacts on the economy of the electronics industry and the reliability of its products. Additionally, it is probable that the environmental damage done as a result of introducing lead-free electronics is considerably greater than it would have been if the status quo had been allowed to continue - not to mention the pollution caused by all the hot air that comes out of the mouths of the proponents and opponents! In any case, the holistic answer to the problem is recycling the solder which is relatively easy and economical, once the infrastructure is in place (in Europe, it will have to be put in place within a few years, as well as going mandatorily lead-free, a paradox). This would allow lead to be continued in use without any of the pretended risk. This easy solution to the whole problem hardly came up, at all, in my search. Why? Brian Ellis Cyprus [email protected]

Book review Environment-friendly Electronics: Lead-free Technology Jennie S. Hwang Electrochemical Publications Limited, IOM Pages 879 +xxxii ISBN 0 901150 40 1 Price £ 119; US$ 238 + p&p URL: http://www.elchempub.com The biggest negative criticism I must make about this book is its title, which is somewhat misleading. There is no in-depth treatment of the environmental consequences surrounding the use of lead in electronics, only a repetition of the eco-political regulations, which are often divorced from the real issues of lead in the environment. In addition, there is no discussion on the use of lead, or its elimination, in any part of electronics other than the small proportion used in solder. Nor is the notion of recycling discussed in relation to either lead-containing or lead-free solders. This criticism apart, this book is an excellent and monumental dissertation, essentially on the metallurgy of leadfree solders, and a worthy complement, in its own rights, to Klein-Wassink’s Soldering in Electronics. Physically, this book is the second fattest in the Electrochemical Publications’ range, beaten only by a short head by the second volume of William Macleod Ross’ A Comprehensive Guide to the Design and Manufacture of Printed Board Assemblies. It is produced in the same style and quality to which we are accustomed. The typesetting seems, subjectively, to be slightly less consistent than usual, in that the line- and letterspacings appear to be variable from chapter to chapter. The majority of the chapters are written by the author, but there are also a few from other sources. Perhaps this provides a good compromise between the single-author work and the edited compiled work. It has not, however, totally eliminated the repetitions and contradictions that sometimes occur in the latter style. For example, Figures 21.8, 23.6 and 28.7 are all an identical microsection of solder fillet lifting. On the assumption that lead-free solder in electronics, if not yet a fait accompli, will become so within the decade, the majority of us within the industry are at the start of a difficult learning curve, especially with respect to choosing the right alloy for a given application. Where near-eutectic tin/lead is an ideal compromise for most soldering work, there is no single substitute that can be dropped in to replace it for all jobs. This is where this book is both ideal and timely. It discusses most, if not all, of the candidates from binary to pentanary (sic! Wouldn’t quinternary be better and more logical?) and higher alloys and all their technical characteristics that may have an influence on their efficacy as a chosen substitute. Their secondary characteristics, such as the cost and availability of the components of each alloy, are not discussed in much detail, just a little bit at the end. This is important when we come to alloys containing elements with market prices in the order of hundreds of dollars per kilogram, such as gallium or indium. Dr Hwang puts quite a lot of

emphasis on alloys containing bismuth, paradoxically ignoring the fact that much of this metal is produced as a by-product of lead mining and smelting. The chemical side of fluxing and cleaning is not discussed in any great detail. The first three chapters constitute an introduction to the book, including such matters as industry trends and legislation. The fourth chapter is devoted to the basics of the technology. The fifth chapter is remarkable as listing all the existing patents on lead-free solder alloys, covering 51 different ternary to sexternary systems. Chapter 6 covers generally all the major binary systems. In this chapter, and the following ones, all the tables of characteristics of the alloys include a comparison row of 63Sn37Pb, so it is easy to discern the major differences between the “new” alloys and our old faithful one. Chapters 7 to 14 each discuss a single ternary or quaternary system, covering the most common candidates, whereas the two following chapters discuss the more exotic ternary and quaternary ones, respectively. Chapter 17 is devoted to higher systems. Up to now, the emphasis has been on substitutes for ordinary solder, whereas chapter 18 discourses on lead-free low- and high-temperature alloys. The nineteenth chapter touches a more ‘handson’ approach, because its subject matter is the finishes required for component leads and printed circuits to maintain both lead-free technology and the ability to solder the components onto the boards to produce a reliable joint. Closely related to this subject is the formation of intermetallic compounds, which is handled in the following chapter, written by experts from the Sandia National Laboratory. Two authors from the Indium Corporation of America discuss the compatibility of leadfree solders with the reflow process in chapter 21. They developed an empirical eqsfor determining the compatibility based on wetting, solder balling, tack time, shelf life and solder appearance. The weightings may be somewhat subjective, but the results are quite interesting. Tables 21.9 to 21.11 summarise the results for 2 profiles, 11 alloys and 10 flux types (developed for tin/lead solders), a remarkable experimental matrix. This may be useful for the beginner to shortlist a number of candidates or, at least, to eliminate the less likely ones, in conjunction with the earlier chapters. Chapter 22 discusses how the structure of tinsilver alloys evolves, bearing in mind the copper of the intermetallic layer. This is quite a long chapter covering both the theoretical and practical aspects of the subject. The following chapter, written by three US authors from diverse organisations, covers the reliability of lead-free solders. Chapter 24 is perhaps a sop to the title of the book, as it treats a non-solder interconnection material, conductive adhesives, and is written by a representative of Cookson Electronics. Although it is a long chapter (74 pages), it obviously cannot be as comprehensive as Liu’s Conductive Adhesives for Electronics Packaging. Notwithstanding, this chapter is a very good introduction to the subject and is perhaps complementary to Liu’s book as it has a more ‘hands-on’ approach. The 25th chapter must not be ignored as it will give invaluable help to anyone wishing to choose a soldering process for his own requirements, by defining the methodology of the experiments necessary in the selection process. As there were so many variables in the experiment design, I was rather surprised that the author, from the University of Massachusetts Lowell, did not suggest a Taguchi matrix for the first selection. Chapter 26

bears the title Application Considerations. This title does not convey much, but the contents are important. Amongst other things, the formation of whiskers, tin pest and the roˆle of gold are discussed. Chapter 27 is also important, bearing in mind that some components cannot be obtained with a surface finish of the connections other than tin/lead. What effect does this have on joints made with lead-free solders? With some alloys, its effect proves somewhat negative. Dr Hwang, in chapter 28, discusses in detail the part that bismuth plays in lead-free alloys. This metal has some unusual properties, on its own, but imparts some significant changes to alloys containing it, some of which may be considered as good and some less so. Like so many things, choosing a bismuth alloy is a matter of compromise. Chapter 29, written by Bob Willis, is a useful English contribution to the book and is a well illustrated guide to defects, which one hopes will never be seen in the practical use of lead-free alloys. Some of these are quite banal, not to be taken in a pejorative sense, but many are less so. Chapter 30 is unusual: it takes the form of a number of views, of the converted (at least, in spirit), of course. Some of them do, nevertheless, express some practical concerns, such as long-term service of assemblies or recycling issues. The last chapter is a series of diverse recommendations and data. The chapters are concluded by many references and most of them have an additional bibliography. The book itself ends, of course, with an index in the usual Electrochemical style, but this is preceded by lists of the many figures and tables. Now, for the nitty-gritty: who needs to buy this book? Well, just about everyone who will be concerned with phasing out lead from PCB fabrication and assembly within the next decade. That means just about every factory in our industry should have at least one copy. Even though I, personally, remain to be convinced that there is any need for eliminating lead in solder on environmental grounds, my recommendation, as a pragmatist realising that it will happen, is buy it.

SMART Group Fifth SMART group european conference 2002 announced The SMART Group has announced the fifth SMART Group European Conference to be held in Brighton between 19–21 November 2002. The conference action committee recently met at NPL to start inviting speakers and to plan the full programme of events for the Fifth SMART Group European Conference, Dr. Colin Lea chaired the meeting with Bob Willis, SMART Technical Director, Mike Judd, SMART PR Director. Peter Swanson, Financial Director will be coordinating the budget and sponsorship opportunities. The conference has, in the past, featured speakers from all over the world, from leading manufacturing and supplier companies. Most popular at recent events has been the three hour workshops and separate technical sessions during the three days of the conference. The conference will again feature a tabletop exhibition, gala dinner and the opportunity to [ 61 ]

.

Plate 1 Delegates at the environmental seminar

.

Topics in Fiberoptics Assembly, and Critical Issues in Hybrid Microcircuit Technology

Please provide 200 –300 word abstract to the SMTA Boston 2002 Conference Coordinator, Kristin Stromberg. Abstracts can be submitted via e-mail as a Word document attachment, to the address: mailto: [email protected]. Or, the abstract can be sent to Ms. Stromberg through the US Mail at the following address: Attention: Kristin Stromberg, Conference Coordinator, Surface Mount Technology Association, 5200 Willson Road, Suite 215, Edina, MN 55424, USA.

New products New convertible head flip chip bonder

network with delegates from all over Europe, in a truly World Class Event. It is also a fun event! Further details on the conference will be circulated on smart-e-link and the web site www.smartgroup.org

The SMART group environmental electronics seminar The SMART Group are again planning what has become an annual seminar and table top exhibition, at Wycombe FC Conference facility in February. These have been predominantly lead-free events. The suggestion this time is to widen the subject matter to include environmental issues. If you would like to offer a paper or suggest a topic/speaker, then please let us know. Keep the date clear in your diary- Thursday February 14–2002 (Plate 1).

Association news SMTA Pan pacific microelectronics symposium preliminary program announced The Seventh Annual Pan Pacific Microelectronics Symposium (February 5– 7, 2002, at the Westin Maui Hotel in Maui, HI) promotes international technical interchange and provides a premier forum for extensive networking among microelectronics professionals and business leaders throughout the Pacific Basin. Sponsored by the SMTA, the Pan Pacific Symposium focuses on the critical business markets and technologies of microelectronic packaging, interconnection, microsystems and assembly. The 2002 technical program consists of 15 sessions with 63 papers on topics such as Flip Chip, [ 62 ]

DCA, Underfills, Eutectic & Lead Free Solder Materials, HDI, High Performance Embedded Passives, Product Management, Thermal Management, Inspection, Emerging Technologies, 3-D Packaging, and a new focus area on MEMS & Optoelectronics. Also featured during the Pan Pacific Symposium are the Keynote Reception and the Plenary Session. The Keynote Reception will feature Martin Goetz, Alpine Micro, with a presentation entitled “SoC or SiP? Alternative Paths to HDI (SLI)!”, and the Plenary Session will focus on Markets and Business, including a presentation from Samsung on Memory Market. Speaker for the Symposium come from the United States, Canada, Italy, Belgium, Germany, Poland, Denmark, Japan, Taiwan, Singapore, and Korea. The program has been arranged by General Chair Iwao Tachikawa, NAMICS, Japan; North American Coordinator Charles E. Bauer, TechLead Corporation, USA; and Technical Program Chair Mahadevan Iyer, IME, Singapore. Visit /pan_pac or contact SMTA executive administrator JoAnn Stromberg for more info on SMTA Dallas: 952-920-7682 or [email protected]

Third technology symposium to be held at SMTA Boston 2002 The SMTA will present the Third Annual Advanced Technology Symposium as part of SMTA Boston 2002 Conference. The Boston 2002 event is being held June 11–12, 2002 at the Bayside Expo Center, Boston, MA. The SMTA Vice-President for Technical Programs, Dr. Ken Gilleo and Conference Chairman, Dr. Paul T.Vianco invite you to submit an abstract to participate in next year’s exciting program. The 2002 symposium will focus on four topics: . .

Future Trends in Surface Mount Technology Case Studies of Failure Analysis in Electronics

A new development of the PESL. DB5-TSV vertical head ultrasonic flip chip bonder flow allows for rapid conversion to horizontal ultrasonics and with standard head it enables use as a highly accurate pick and place machine (Plate 1). The vertical ultrasonic head enables improved planar attachment of very fragile chips such as galium arsonide and for some of the very small microwave chips presently in development. Conventional gold bump chips can then be bonded using cither horizontal or vertical ultrasonics, involving only a very quick and simple change of the active head unit. Further information regarding PFSL products may be found at: www.pesl.win-uk.net.

Dexter introduces 2608C reflow silver-filled conductive adhesive Dexter Electronic Materials has introduced QMI 526, a silver-filled thermally conductive adhesive used to attach integrated circuits and components to advanced substrates including PBGAs, CSPs, stacked die, and array packages based on flexible tape and organic laminates. This low stress material is designed for use with large die ( . 500 mil) or for applications where the substrates’ co-efficients of thermal expansion (CTE) differ substantialy. This product is qualified to perform in lead-free, 2608C reflow applications. Thermally conductive, hydrophobic, and stable at high temperatures, the new adhesive delivers void-free bond lines with excellent adhesive strength on a wide variety of organic and metal surfaces including solder mask, BT FR, polyimide,

Plate 1

gold, Kaptone, and Mylare. QMI 526 offers a robust adhesion mechanism that can compensate for the presence of fluorine or other backside residues, for example on wafers. The adhesive also offers excellent electrical properties. Designed to achieve UPHs substantially higher than conventional oven cures, it is formulated to cure in-line either on the diebonder using a postdiebond heater or on the wirebonder preheater. QMI 526 can be cured in-line using Skip-Curee processing in just 8 seconds at 1508C, or using a variety of times and temperatures in conventional ovens or snap cure stations. For more information visit: www.loctite.com/electronics.

Loctite multicore introduces antitombstoning solder pastes formulated with 63S4 alloy Loctite Multicore has introduced new family of solder paste products formulated with the unique 63S4 alloy that eliminates costly tombstone defects and reduces component misalignment that occurs when reflowing boards with very small components. Multicorew 63S4 solder pastes produce safe residues and are effective over a wide range of reflow profiles in air or nitrogen. Suitable for fine pitch printing to 0.4 mm and high speed stencil printing up to 150 mm/sec these products contain a high activity no-clean flux that is highly effective on difficult to solder substrates such as HASL, OSP copper, gold over nickel, and silver immersion. All three 63S4 products offer excellent abandon time, long open time, exceptional slump resistance, and superior tack life. For more information visit: www.loctite.com.

EKRA announces the release of the E5-36 large format in-line stencil printer With a print area of ”36 £ 24” the E5-36 offers the largest board size capability available today in an in-line fully automatic printer, while offering features like the patented EKRA vision alignment system with software controlled variable lighting technology, 2 1/2D post print inspection system, and vacuum stencil wiping. These features combined with a heavy-duty transport system allow for the same repeatability and robustness performance levels, while processing large PCB’s, as when printing smaller substrates. www.ekra.com

IC-SCAN DE tracks MSDs for double side reflow boards Cogiscan, Inc. announces the IC-scan DE for tracking and controlling Moisture Sensitive Devices (MSDs) for double side reflow applications. Utilizing barcode readers integrated into SMT line conveyors, the IC- Scan DE system provides complete material visibility throughout the factory, including between the first and second reflow.

The IC-scan DE system improves product quality by automating challenging material control requirements. The system tracks the remaining floor life of all MSDs on the first side of each Printed Circuit Board (PCB) and continues to track the partially assembled board and remaining floor life until the second side reflow is completed. Using the floor life and location information in its database, the IC-Scan DE system calculates exposure times of the MSDs and generates warnings and alarms when operator intervention is required. An optional SMEMA interrupt can be used to stop SMT line conveyors from transferring a board front the placement machine before the system has verified that all components on the first and second side are still within their specified floor life limit. Existing IC-scan systems may be upgraded to IC-scan DE as needed. www.cogiscan.com

appropriate kit of parts and its operation verified before it is placed on the machine. The test jig incorporates a camera and crosshair generator that allows the user to correctly calibrate the pickup position, as well as visually monitoring the performance of the set position while the feeder is indexed at machine speed. The jig is fully programmable for all functions, including speed and number of indices and selection of single or multiple indices. The system comes complete with a feeder test position for Fuji CP4, CPO or CP7 feeders in widths from 8 to 24 mm. It can cycle each feeder up to 500 times at speeds of up to 30,000 components per hour. It also ensures that the top film of the component reel is peeled back and collected on the take-up reel. For further information visit: www.altenativesmt.com

Selective soldering New thin bondline, system from Vitronics thermal interface material Soltec Intertronics has released Thermoset Geleasee MG-121, which combines the extremely low thermal resistance properties of a grease with the integrity of a gel. Gelease-MC-121 can be used in cross-sectional thickness as low as 1–2 thou in order to minimise the thermal pathway and maximise heat flow. It offers the same thermal performance as a high performance grease while minimising the possibility of bleed, separation and ‘pump-out’. Gelease-MC-121 is designed for applications such as flip chip to thermal plate, BGA to heatsink and PGA to heatsink. The new product exhibits excellent flow properties and dispenses easier than most thermal greases. A short, low temperature over cure or the operating temperature of the processor itself will allow the material to gel to its final integral state. The material’s thermal resistance is 0.06– 0.078C in2/W. It has a low Modulus of Elasticity ð3:5 £ 105 PaÞ and a very low Glass Transition Temperature ( 2 1218C). Full details at www.intertronics.co.uk

Test jig boosts productivity for users of surface-mount feeders AlternativeSMT has developed a Fuji feeder test jig to check the performance and functionality of feeders for the Fuji CP4, CP6 and CR7 surfacemount assembly machines. The test jig was developed in response to the problems faced by Fuji users in effectively carrying out feeder maintenance and ensuring the performance of feeders once they are fitted to the machine. Often, serviced feeders are returned several times before they function normally. These problems result in increased machine downtime during changeovers and production. The response from alternativeSMT was to develop a functional feeder test jig coupled with a comprehensive feeder maintenance program. The new jig allows users to test the feeders off line in order to minimise machine downtime caused by adjustments after the feeder is mounted on the machine. The test jig carries out a full functional test on the feeder, including checks on the covertape removal process and positioning accuracy and repeatability. A feeder can be loaded with the

Vitronics Soltec’s new Selective Soldering System is a highly flexible, fast and reproducible automated soldering system. The system delivers high throughput with highest quality in batch or inline production. It offers fully modular set-up, allowing for easy and quick upgrades in the field, and its robust, flexible platform offers a wide variety of modules that can be easily added and interchanged based on a “plug-and-play” principle. Selective Soldering s X-Y robotic system offers full process flexibility and high throughput via its ‘smart’ sub-system set-up. Flexible SecureGrip systems in combination with a programmable MultiFlux minimize the need for customized tooling and applications. Multiple preheat zones address the need for higher preheat requirements without impacting cycle times. Combined MultiWave soldering and SelectWave soldering provides optimum balance between high throughput and full flexibility. Selective Soldering consists of two soldering stations – SelectWave and MultiWave – that can be configured four different ways. The combination of MultiWave, for reducing cycle time, and SelectWave, for highest soldering accuracy and flexibility. For further information visit: www.vitronicssoltec.com.

High-performance stencils for vectore and tetrae frames Design engineers at TSL, have developed Apex, a remountable stencil for use with the widely adopted Vectore and Tetrae frames – offering fast, easy and accurate fitting with superior tension characteristics from patented stress-relief technology. The Apex stencil allows users to increase yields and decrease rework. Self-aligning trapezium shaped fixing slots, around the four sides of the stencil, provide effective and easy location of the stencil onto the frame’s interlocking comb mechanism. Only two simple operations are now required to mount a stencil on the frame overcoming complicated alternative procedures. Improved printed results are readily achieved when mounted to Vectore and Tetrae frames, [ 63 ]

Plate 2

fully programmable PCB/stencil separation parameters. An under stencil cleaner option is available with both wet and dry modes plus vacuum assist. For more information visit: www.reprintservices.co.uk

Plate 4

Pyramax lowers cost of ownership

using existing printing parameters. Individual foils can be produced from the company’s comprehensive range of in-house manufacturing techniques, including electroformed, laser cut, precisionetched and multi-level stencil combinations. This ensures a total technology choice, delivering optimum performance for the most demanding applications – from ultra-fine pitch to glue printing and pin-in-hole-reflow (PIHR) stencils. Proven stress-relief tensioning for optimum stencil flatness, easy mounting and a wide choice of manufacturing options are combined to ensure maximum performance from a new generation of remountable stencil (Plate 2).

New printers from Reprint Featured at Productronica by Reprint Services based near Dorchester, Dorset, were the R29 and Compact 29 range of screen printers each designed to handle screen sizes up to 29”. The all-new R29 series of screen printers bring major advances in the technology and functionality of mid-range stand-alone solder paste printers. Providing features and user benefits previously only found in more expensive SMT printers, the R29 series delivers unrivalled performance and cost advantages compared to similar systems. Easy-to-use Windowse based interface combined with state of the art micro-controller based servo motion control, provide precise repeatable control of all printer functions including fast product change over. High accuracy, very fine pitch printing is achievable with the easy-to-use vision system that along with the low running costs and outstanding reliability make the R29 series of printers the ideal choice for a wide range of production environments. All functions of the printer can be controlled via the custom written Windowse based interface. Operators need no special expertise with the software providing a set-up wizard for fast, reliable product configuration. Using the unique button-driven cursor, this userfriendly program allows easy access to the extensive feature set. The R29 has a maximum board size of 800 £ 560 mm (min. 50 £ 25 mm) with print speeds up to 200 mm/sec and print pressures up to 20 Kg with [ 64 ]

BTU Europe recently launched the Pyramax 150 and 98, in addition to the established and cost effective VIP 70 system. The BTU Pyramax range offers two systems – the Pyramax 150 incorporating a 150” heated length and 10 controlled heated zones and the Pyramax 98 incorporating a 98” heated length and seven controlled heated zones. Developed to criteria set by the world’s largest EMS producers, the Pyramax provides a higher operating temperature for lead-free reflow, increased flexibility, reduced gas consumption and maintenance together with improved ergonomics. Radical changes in design and manufacturing technologies enable BTU to provide the fastest order to shipment time of any oven in the market. The Pyramax’s maximum operating temperature of 35O8C makes it ideal for the higher temperatures required for lead-free reflow, whilst a 24-inch processing capability and easier configuration provides greater flexibility. As with aIl BTU ovens, the Pyramax features the proprietary Flux Management System to remove flux residue from the process chamber which now incorporates a new electronic maintenance procedure assistant, easier access to filtration and trap arrangement together with hot zone recirculation to minimize condensation (Plate 3). For further information, please contact: BTU Europe, Tel: +44(0)1252 660010 Fax: +44(0)1252 660011

Plate 3

Focus is on the Hawk X-Tek, has recently launched their new Hawk System which boasts impressive magnification abilities and rapid real-time X-ray inspection. A significant increase in magnification capability of up to 6000 £ is due to a new, advanced xray target and the latest X-Tek in-house suite of software ensures that the Hawk user-friendly in all areas of its function. X-Tek’s ixs control and analysis software, which runs on Windowse NT platform, allows the user to operate the system manually or program for automatic inspection. Other features include a unique x-ray source which incorporates the vacuum assembly, electron gun and HT generator in one housing to offer greater reliability. This design has reduced the cost and physical size of the unit to a space-saving footprint of only 975 mm £ 1195 mm (Plate 4). www.xtek.co.uk

Multicore solders introduces new lead-free no-clean solder cream for SMT assembly Multicore Solders, Inc, has introduced new Multicorew CR32, a lead-free no-clean solder cream suitable for printing and reflow in air where process yield is critical. CR32 is a medium residue product suited to meet the demands of high volume production processes that use components and boards with less than the desired level of solderability. Multicore CR32 produces safe, non-corrosive, clear post-soldering residues that eliminate the need for cleaning and will not pose a hazard to long-term reliability. This lead-free cream offers excellent open time and tack performance, superior slump resistance, and good soldering activity for greater tolerance to process variations and minimal solder balling. CR32 contains an IPC Type L flux formulation and meets ROLO to ANSI/J-STD-004 standards. For more information visit: www.loctite.com.

New electronics coating lowers cost of moisture protection for PCBs 3 M has launched the new 3 Me Novece Electronics Coating EGC-1700, to provide a fast

Plate 5

and low cost way of protecting sensitive electronic circuits and assemblies against moisture and corrosive chemical environments. Novec EGC1700 dries quickly to form an extremely thin, yet remarkably repellent, light-duty coating. This new product now makes it viable to apply moisture protection to inexpensive circuits and assemblies, where the cost of most conformal coatings is prohibitive. It is a clear, low viscosity solution of a fluorochemical acrylate polymer coating, diluted in a 3 M Novec hydrofluoroether solvent. The solvent is nonflammable and has low toxicity, zero ozonedepletion and low global warming potential, to provide acceptable workplace safety and environmental properties. When applied to clean and moisture-free surfaces such as copper, aluminium, ceramics, steel, tin or glass, Novec EGC-1700 dries quickly to create a thin, transparent, protective film with excellent anti-wetting, antistiction, antimigration and anti-corrosion properties (Plate 5). For further information visit www.3M.com

Industry news Oxford Instruments Analytical extends product range to include coating measurement instruments Oxford Instruments Analytical has acquired the assets of CMI International (CMI). The new Oxford Instruments business unit is based at the CMI premises in Chicago with Bradley Boyer, President of Oxford Instruments X-ray Technology Group, as its new President. The initial priorities are to increase the depth of management strengthen internal processes and create a sound base from which to grow the business. This will be achieved with the help of appropriate complementary engineering, manufacturing and distribution strengths in Oxford Instruments. Many of CMI’s products are based on X-ray technology, an area where Oxford Instruments is already a leader.

Plate 1 New aaliance in the SMT arena

ems – ALLIANCE the new concept for electronics manufacturing ems-ALLIANCE consists of six companies in six different countries. Together they form a world wide electronic manufacturing services network of independent and equal providers, co-operating to be a competitive alternative in the electronics business. The initiative to form ems-ALLIANCE was taken by the Swedish EMS company NOTE. The basic idea behind the alliance is to offer local electronics production – no matter where in the world your market is situated. The thought is that you can stay with your local producer and still reach the whole world. “All companies in ems-ALLIANCE are strong players in their respective home market and by working together we learn from each other and can further increase our knowledge and expertise,” says Marina Filipsson, Chairman of ems-ALLIANCE. Current members of the ems-ALLIANCE are situated in Brazil, China, India, Italy, Sweden and in the USA, represented by the following companies: Brazil – VRI – Industria Eletronica Ltda China – LeeMAH Holdings, Ltd. India – Rangsons Electronics Italy – Elem S.p.A Sweden – NOTE AB USA – Xetel Corporation For further information, please visit: www.emsalliance.com

The National Physical Laboratory celebrates the strength of UK measurement through national awards The UK’s highest industrial and academic achievers in the field of metrology (the science of measurement) were honoured last week at the Metrology for World-Class Manufacturing Awards 2001.

Accurate measurement contributes around £6.5 billion pounds to the UK economy and is one of the key technologies for manufacturing. These important awards reflect the entire metrology application process, from laboratory to factory floor, and the National Physical Laboratory (NPL) has been a main sponsor since 1995.

New alliance in the SMT arena Assemble´on and Cookson Electronics’ Speedline Technologies have joined forces to provide SMT manufacturers with best-of-breed equipment and integrated line solutions (Plate 1). Assemble´on and Speedline announce a groundbreaking strategic alliance. The Alliance combines the wide-ranging expertise and resources of these industry-leading organisations, merging Assemble´on’s strengths in pick-and-place systems and processes, and Speedline’s expertise in equipment, systems and processes. For further information contact Wenny Hageman, Assemble´on, Regional Centre Europe, Building HVM, PC Box 216, 5600 MD Eindhoven. Telephone: +31 4027 66444. Fax: +31 402766541, Email: [email protected]

A range of new chemicals Nairobi, September 14 2001 - A range of new chemicals, used in everything from fire extinguishers to cleaning fluids, are appearing on the market to the concern of scientists studying the ozone layer. The new substances, with names such as npropyl bromide and halon-1202, are not controlled by the Montreal Protocol which lists ozone-depleting substances that are to be phased out. Studies indicate that some of the new substances, which are being used as replacements for banned ones, may have the potential to damage the ozone layer. This protective layer is located in the stratosphere where it shields life on Earth from high levels of cancer-causing ultra violet light. [ 65 ]

The quantities being manufactured are at the moment believed to be small. But scientists at universities and institutes around the globe, along with researchers at the United Nations Environment Programme (UNEP), are concerned that over the coming years they may be produced in ever increasing quantities. UNEP has estimated that the ozone layer and the ozone hole over Antarctica, which so far this September is extending over 24 million square kilometers or an area about the size of North America, will recover by 2050. This is as a result of the banning and phasing out of existing, long lived, ozone depleting chemicals such as chloroflurocarbons (CFCs). These were widely used in products such as hair sprays until they were outlawed by the Montreal Protocol. But the emergence of these new chemicals has triggered concern that this recovery date may be significantly delayed. Klaus Toepfer, the Executive Director of UNEP, said: "The Montreal Protocol has been a success story of which we can all be proud. Ninety-six ozone damaging chemicals have been banned or are being phased out and 1.3 billion US dollars have so far been contributed by developed countries to the Multilateral Fund to help developing countries implement the Protocol.". The fund was set up in the early 1990s to provide support to developing countries to help them phase out listed substances. "But we must remain vigilant if our success story is to ultimately have a happy ending. Some of these new, replacement chemicals, may prove to be no threat at all to the ozone layer although they may pose threats to human health, wildlife and the environment generally. Others however may have the potential to cause significant damage to stratospheric ozone, undermining our efforts to date. I would urge countries to carry out immediate scientific assessments of these new chemicals and to ban those that are shown to have real ozonedepleting potential. Finally governments, industry and organizations like UNEP must, based on sound science, work together to devise a long term strategy so that we know in advance the ozone depleting potential of future chemicals before they appear on the market," he said. Such a strategy might involve a public/private partnership where companies voluntarily agree to test and assess a new chemical for its ozone damaging potential, presenting their findings to governments before manufacturing and marketing the new substance. Mr Toepfer’s views were echoed by Professor Mario Molina, a joint winner of the 1995 Nobel Prize for Chemistry for his work on the chemistry of stratospheric ozone, and Institute Professor at the Massachusetts Institute of Technology in the United States. "It is important to keep the issue of new, emerging, chemicals in perspective. Firstly we must be worried about those substances with a large potential for depleting the ozone layer like the old CFCs. But even these new, short lived, substances need to be watched, "he said. Until recently it was thought that these new substances could not damage the ozone layer. It was thought they did not live long enough to reach the stratosphere. However new research is changing this view. There is evidence that under the right conditions such substances and their breakdown products can travel far enough to reach the Earth’s protective shield. [ 66 ]

"At the moment I believe we do not have a big problem with these new substances. But we cannot be complacent. If enough of them are manufactured and emitted, we will delay the recovery of the ozone layer quite significantly, " said Professor Molina, also a winner of the UNEP Sasakawa Environment Prize. He urged for more scientific studies to fully assess the ozone damaging potential of the new chemicals: "We need to know which of them are safe and which may be a worry in the future. We have enough experience to know that we cannot take the ozone layer, which shields life on Earth from levels of ultra violet light which can cause cancers and eye cataracts, for granted". Professor John Pyle of the Centre for Atmospheric Science at Cambridge University in England and another leading ozone layer expert said: "Relatively small amounts of these new substances are being produced. But new measurements indicate that the levels of some of them are growing rapidly in the atmosphere. In the past, for example with the CFCs, we only took action once we had a problem, when we found damage to the ozone layer was already under way We do not want a repeat of this so we must not take our eye off the ball". At least four new substances with a potential to damage the ozone layer have been identified. Nelson Sabogal of UNEP’s Ozone Secretariat based in Nairobi, Kenya, said: "These are however only the ones we know about so far. There may be many more. One of our difficulties is that each of these new substances can have up to fifteen different names which can make monitoring and tracking of them in international trade a tough task for governments". Hexachlorobutadiene has an ozone depleting potential (ODP) higher than some ozone damaging substances which are already banned under the Montreal Protocol. (See notes to editors on the meaning of ODP). It is used as a solvent and is produced as a by-product of chlorinated chemical production such as the manufacturing of vinyl chloride. It has been classified as a high-volume production chemical by the Organization for Economic Cooperation and Development with one country reporting that its factories were now producing more than 10,000 tonnes annually. A Canadian government report has concluded that Hexachlorobutadiene has "the potential to contribute somewhat to the depletion of stratospheric ozone". n-Propyl Bromide is being used as a solvent in applications including degreasing, vapour cleaning and the cold cleaning of metal parts. Among the new substances of concern, this one has been the most intensively studied by the Technology and Economic Assessment Panel of the Montreal Protocol. It has concluded that npropyl bromide is being aggressively marketed .as a solvent, a feedstock and as a carrier and intermediate for pharmaceutical and other industries". The Panel estimates that around 5,000 to 10,000 tonnes of this substance are being used and emitted annually. It estimates that by 2010, emissions are likely to rise to between 20,000 and 60,000 tonnes as industries switch from banned substances to this one. Professor Pyle, a member of the Scientific Assessment Panel of the Montreal Protocol, said that many of these new chemicals, like n-propyl bromide, were short-lived. Their ability to reach

the ozone layer depends on the time of year and geographical location where they are used, new studies show. This has made it difficult to precisely assess its ozone-depleting potential. However some preliminary estimates indicate it may be as damaging as substances which are already banned or subject to phase-outs. 6-bromo-2-methoxyl-naphthalene. Little is known about the scale of manufacture or the ozone depleting potential. It is currently being used in the manufacture of methyl bromide which is a fumigant that is subject to a world-wide phase out under the Montreal Protocol. Halon-1202 is extremely effective for putting out electrical fires and is being used by the armed forces in some countries. Scientists have yet to estimate the quantities being made and emitted. But new research indicates that it may be long lived. Its potential to deplete the ozone layer has been tentatively estimated to be higher than some banned substances. The Scientific Assessment Panel of the Protocol have been asked to study this issue. Their findings are expected to be disclosed at the Thirteenth Meeting of the Parties to the Montreal Protocol who are scheduled to meet in Colombo, Sri Lanka, in mid October. Apart from the new substances entering the market, the phase out of existing ones which have been definitely identified as damaging the ozone layer, is far from complete. Hydrochloroflourocarbons (HCFCs) will not be fully phased out until the year 2030 and methyl bromide, which as mentioned before, is used in agriculture as an insecticide, is not due to be fully phased out until 2015. Meanwhile smuggling of banned substances is continuing. A report in the Japan Times last month said more than 100,000 bottles of CFC-12, have been circulating in Japan. The chemical, which was banned in industrialized countries in 1996, is used in car air conditioning units. It is believed that the chemical was illegally imported from countries where production of CFC-12 is still permitted. Notes to Editors The Ozone Depletion Potential or ODP of a substance is based on a comparison with the CFC-11. So CFC-11 has an ODP of one and all other substances are either bigger than one or lower than one. Hexachlorobutadiene has an ODP of 0.07 npropyl bromide, which has a variety of other names including 1-bromopropane or nPB, may have an ODP of between 0.0033 and 0.111. If the ODP is found to be at the higher end, then it is potentially more damaging than HCFC-21 which is already controlled by the Montreal Protocol 6bromo-2-methoxy-naphthalene has other aliases including BMN. There is no information of its likely ODP Halon-1202, which has alternative names including dibromodifluoromethane, Freon 12-B2, R12B2 and UN1941, may have an ODP of between 0.3 and 1.7 A report by the Scientific Assessment Panel, Assessing the Impacts of Short-Lived Compounds on Stratospheric Ozone, was published by the UNEP Ozone Secretariat in May 2000. A document by the Technology and Economic Assessment Panel, Montreal Protocol on Substances that Deplete the Ozone Layer, was published by UNEP’s Ozone Secretariat in April 2001. It carries a Report on the Geographical Market Potential and Estimated Emissions of n-Propyl Bromide. www.unep.org/ozone or www.teap.org

International Ozone Day is an annual event that takes place on September 16. This year’s slogan is Save O3ur Sky:Protect Yourself Protect the Ozone Layer. Web site for downloading posters, copyright UNEP, is www.unep.org/ozone/ozone_day2001/ The site also has a graphic showing the decline of CFCs in developed countries and a small increase in use in developing ones. For more information on this news release and regional events and activities linked with International Ozone Day please contact: Nick Nuttall, Head of Media Services, UNEP, P.O. Box 30552, Nairobi, Kenya. Tel: 254 2 623084, Mobile: 254 (0) 733 632755, E-mail: [email protected] or Tore Brevik, UNEP Spokesman/Director, Division of Communications and Public Information, on Tel: 254 2 623292, Fax: 254 2 623927, E-mail: [email protected] or Nelson Sabogal, Ozone Secretariat, UNEP, on Tel: 254 2 623856, E-mail [email protected] or Rajendra Shende, Chief, Energy and OzonAction, Division of Technology, Industry and Economics, UNEP, on Tel: 33 1 44371459, Fax: 33 1 44371474, E-mail: [email protected] Scientific Contacts: Professor Mario Molina of MIT on Tel: 617 253 5081, E-mail: [email protected] or Professor John Pyle, Cambridge University, on Tel: 44 (0)1223 311772, E-mail: [email protected]

Appointments New man at the helm for X-Tek X-TEK has announced the appointment of a new Managing Director. Peter Marsh joined the company at the beginning of September, having spent the last four years as Sales, Marketing and New Product Development Director with Truvox International who manufacture industrial floor care machines and equipment (Plate 1). Peter brings with him a wealth of experience in the manufacturing industry, gained from employment with companies such as Northern Telecom and Rank Xerox, from which he developed his Plate 1 Peter Marsh

interest and enthusiasm tot the High Tech environment in which X-Tek is so active. He started his career as a Production Engineer before moving up through the ranks. This range of experience has helped with Peter’s desire to understand the whole process from design and manufacture through to the final product.

Plate 1 Hand soldering for through-hole components

A new CEO for Instem Technologies Instem Technologies has announced the appointment of Peter Robertson as Group Chief Executive. As Group Chief Executive, Peter Robertson will be responsible for the strategic and operational direction of the Company, together with the implementation of a five-year development plan. Peter Robertson joins Instem technologies from Bran + Luebbe, the international pumps and processing technologies company, where he has been Managing Director of its UK business for the past nine years.

. .

IPC-VT-43 Hand Soldering for ThroughHole Components This video aimed at Through-Hole Soldering and application: . .

Tecnomatix-Unicam expands executive team Tecnomatix Technologies has appointed Sean Roche to the key position of new vice president of product marketing for Tecnomatix-Unicam, its subsidiary specialising in manufacturing process management software for the electronics assembly industry. Mr Roche brings a wide range of product marketing and management skills to the company, and promises to focus specifically on customer and partner relations as Tecnomatix-Unicam strengthens its market leading position. Mr Roche has over ten years’ experience in the enterprise data storage industry, most recently as marketing manager for Quantum Corporation, where he was responsible for all product marketing efforts in the DLTtape division. Prior to Quantum, be held various marketing and engineering positions at Kao Infosystems Company, a subsidiary of Kao Corporation. He has a BSc in business administration from the University of Massachusetts, USA.

Publications IPC-VT-42/43 hand soldering for through-hole components IPC has launched two new hand soldering training videos IPC-VT-42/43. These new hand soldering videos demonstrate hand soldering theory and procedures under highly magnified/ideal conditions to visually imprint the proper procedures and practices (Plate 1). IPC-VT-42 Hand Soldering for ThroughHole Components This video covers tools, materials and other workstation practices such as: . . . . . .

Personal safety considerations ESO prevention Soldering iron types Temperature selection and heat transfer principles How to avoid thermal damage Tip and solder types! sizes and applications

Flux theory and low residue fluxes Wetting principles

. . . .

Soldering through-hole components on to printed circuit boards and thick boards Tip tinning, tip and tool maintenance Re-heating solder Heat sinks/thermal shunts Cleaning flux residues Solder joint workmanship requirements to IPCA-610C including: – fillet shapes and contact angles – cold and disturbed solder – Fractures, thermal damage and lifted lands – Solder bridging, projections and balls – Nonwetting and dewetting – Blowholes and pinholes – Product classes per J-STD-OO1 C

IPC-VT–42/43, when working in conjunction with IPC-VT-49 The Seven Sins of Hand Soldering and IPC – DRM-40 Thru-Hole Solder Joint Evaluation. Preview all the PC videos at /train.htm

New standard published in the BS-7000 series Obsolescence affects all equipment, software, tools, processes, support products and specifications. This British Standard gives guidance for establishing a framework for obsolescence management and for planning a cost-effective obsolescence management process that is applicable through all phases of the product life cycle. It thus aids in minimizing both costs and any detrimental impact on consumers and manufacturers. BS 7000-5 is applicable to all products that include electronic, electrical and electro – mechanical components, although its use can be extended to include other products and processes. The BS 7000 Design Management Systems Series currently comprises the following parts: – Part 1: Guide to managing innovation; – Part 2: Guide to managing the design of manufactured products; – Part 3: Guide to managing service design; – Part 4: Guide to managing design in construction; – Part 5: Guide to managing obsolescence (this part; – Part 10: Glossary of terms used in design management [ 67 ]

BS 7000-5 costs £50 (£25 to members) and is available from: BSI Customer Services, 389 Chiswick High Road, London W4 4AL, England Phone:-+44 (0)20 8996 9001, fax: + 44 (0)20 8996 7001, Email

International diary IPC Expo 2002 24 – 30 March Long Beach Convention Center, CA, USA IPC WWW: www.ipcprintedcircuitexpo.org

Nepcon Shangai 9– 12 April Shanghai Everbright Convention & Exhibition Centre, Shanghai, China Reed Exhibition Companies (REC) CCPIT Electronics & Information-Industry Sub Council Tel: +(852) 28240330 Fax: +(852) 28240178 E-mail: [email protected] WWW: www.nepconchina.com

Electronic Week 2002

Globaltronics

24 – 26 April

3– 6 September

COEX, Seoul, Korea

Singapore Expo

Kyungyon Exhibition Corporation Tel: +82 2 785 4771 Fax: +82 2 785 6117 WWW: www.electronicweek.org

Reed Exhibition Companies Tel: +65 434 3615 Fax: +65 334 2748 Email: [email protected] WWW: www.globaltronics.com.sg

FiberComm 4 – 6 June Munich Messe, Munich, Germany Messe Mu¨nchen GmbH Tel. (+ 49 89)949-20315 Fax (+ 4989)949-20319 E-mail: [email protected]

14 – 17 April Paris, France WWW: www.eurosime.com

IMAPS Tel: +1 202 548 4001 E-mail: [email protected] Website: www.imaps.org

Nepcon UK

Bayside Expo & Convention Center, Boston, MA, USA

1– 3 October 2002

Tel: +1 203-840-5344 Fax: +1 203-840-9344 WWW: www.nepcon.com/nepconeast/suggest.asp

IMAPS Israel 2002 Daniel Hotel, Herzelia – On – Sea, Israel Website: www.imaps.org

European Microelectronics Packaging & Interconnection Symposium 16 – 18 June Selim Achmatowicz Tel.: +48 22 835 30 41 ext. 457 Fax: +48 22 834 90 03 E-mail: [email protected] http://www.itme.edu.pl/imaps.cracow2002/index. html

17 – 19 April

SMT/Hybrid/Packaging

Tokyo Ryutsu Center, Tokyo, Japan

18 – 20 June

JIEP Tel: +81-3-5310-2010 Fax: +81-3-5310-2011 E Mail: [email protected] Web: http://www.jiep.or.jp/jiep/icep/index.html

Messezentrum, Nu¨rnberg, Germany

[ 68 ]

Colorado Convention Center, Denver, CO, USA

10 – 12 June

Sofitel Hotel, Cracow, Poland

2002 International Conference on Electronics Packaging (ICEP) and 16th Microelectronics Show

4– 6 September

Nepcon East

13 June

EuroSimE 2002 3rd International Conference on Benefiting from Thermal and Mechanical Simulation in Microelectronics

IMAPS USA

Mesago GmbH Tel: +49-711-61946-79 Fax: +49-711-61946-93 WWW: www.mesago.de

NEC Birmingham, Halls 1 & 2 Reed Exhibition Companies Tel: + 44 (0)20 8910 7706 Fax: + 44 (0)20 8334 0608 WWW: www.nepcon.co.uk

EPC2002 – European PCB Convention & ECWC9 World Conference October 8– 10 Koln Messe, Cologne, Germany EIPC Services GmbH Tel: +41 61 482 3900 Fax: +41 61 482 3910 E-mail: [email protected] Website: http://www.eipc.org

Electronica 2002 – The 20th International Trade Fair for Components & Assemblies in Electronics November 12 –15 New Munich Trade Fair Centre, Munich, Germany Press Officer: Angela Prag Tel: +49 89 949-20670 Fax: +49 89 949-20679 [email protected] Website: www.electronica.de

NEPCON WEST

SMT/Hybrid/Packaging

December 4– 6

6 – 8 May

2004 Nepcon UK

McCenery Convention Center, San Jose CA, USA

Messezentrum, Nu¨rnberg, Germany

October 2004

Phone: 203-840-5689 Fax: 203-840-9689 Website: www.nepconwest.com

2003 Nepcon UK South 2– 3 April 2003 Brighton Metropole, England Reed Exhibition Companies Tel: + 44 (0)20 8910 7706 Fax: + 44 (0)20 8334 0608 WWW: www.nepcon.co.uk

Mesago GmbH Tel: +49-711-61946-79 Fax: +49-711-61946-93 WWW: www.mesago.de

NEC Birmingham, England Reed Exhibition Companies Tel: + 44 (0)20 8910 7706 Fax: + 44 (0)20 8334 0608 WWW: www.nepcon.co.uk

Nepcon UK North 16 – 17 September 2003 Harrogate International Centre, Harrogate, England Reed Exhibition Companies Tel: + 44 (0)20 8910 7706 Fax: + 44 (0)20 8334 0608 WWW: www.nepcon.co.uk

[ 69 ]

Note from the publisher

How to get the best from your Emerald subscription If you are reading this then the chances are that you or your organisation/institution are a subscriber to Soldering & Surface Mount Technology. But, are you using your electronic entitlement? Whatever the level of subscription, Emerald delivers the following access features:

.

Academic Institution wide license offering unrestricted, concurrent, multi-user access Seamless IP-based authentication

.

Additional username and passwords for remote access

.

ATHENS authentication for remote access

.

With Emerald you have content that is second to none and a search engine that is tailor-made to enable time-pressed practitioners, researchers and students to locate the articles that best meet their requirements. You can search through 8 years of content using a whole range of search criteria such as keywords, author and article title. If you prefer browsing, you can do this on an issue by issue basis. In addition to the articles including archive, the nonarticle content of our journals (industry round-up, new products and materials etc) is available online, so that it can be accessed by users simultaneously. It is possible to browse or search by keyword, allowing users to keep up to date with everything that is happening in their specific subject areas. Visit www.emeraldinsight.com/ssmt.htm to access the electronic element of your subscription. Emerald Alert provides personal e-mail notification of new articles published online, ensuring that users are kept up to date with the latest research. Emerald Alert includes: .

TOC (Table of Contents) of any chosen journal

.

Weekly/Monthly Digests

.

Saved Search Alert options

For further information and to register please log onto www.emeraldinsight.com/news/email While Emerald is designed to be very easy to use, we do, of course, offer full on-screen help and the services of a technical helpdesk — in fact everything you need to get started and to use your electronic access to best effect. For more information visit the Emerald web site at http://www.emeraldinsight.com or contact Emerald Customer Support at [email protected] Literati Club – the benefits of writing for Soldering & Surface Mount Technology In addition to achieving wide dissemination of your work by publishing with Emerald, our author relations service – the Literati Club — provides services and support for those authors who publish with an Emerald journal. It is a tangible expression of commitment by a publisher to its authors and Editors — we offer support and resources for all Emerald contributors worldwide.

Soldering and Surface Mount Technology 14/1 [2002] 70 q MCB UP Limited [ISSN 0954-0911]

[ 70 ]

The Literati Club offers the following benefits: † A ‘thank-you’ to outstanding Authors and Editors The Literati Club hosts an annual Awards for Excellence, which celebrates outstanding contributors among Club Members in the following categories: .

Best Paper Award for each participating journal

.

Editor of the Year

.

Leading Editor Awards

.

Research Fund Awards

.

Regular Literati Club Newsline

All major developments occurring within the Literati Club and Emerald as a whole are reported and guidance is given on publishing, writing and editing. — Liberal copyright principles Emerald seeks to retain copyright of the articles it publishes, without the author giving up their rights to use their own material. This enables users to .

Work uncumbered in our development of electronic publications and their delivery to meet customer and end user needs and create maximum dissemination of author’s work.

.

Protect our authors’ moral rights and their work from plagiarism and other infringement

.

Recoup 100% of copyright fees from Reproduction Rights Organisations to reinvest in new initiatives (dependent upon 90% or more of copyrights being assigned to Emerald in 90% of journals)

.

Provide an efficient ‘one-stop-shop’ for permissions

— Authors’ Charter A unique Charter detailing your rights as a contributor to an Emerald journal. † Privileged personal subscriptions to home address Any Literati Club Member may at any time take out a personal subscription to an Emerald journal at one-third of the published price for delivery to a home address. † Complimentary personal subscription You will personally receive a complimentary subscription to any Emerald journal in your area of interest if your library takes out a full-price subscription. † Calls for papers All Editors’ calls for papers in your areas of interest and previous article keywords will be mailed or e-mailed to you personally. † Photocopying rights You are authorized to make up to 25 copies of any single article published by Emerald, without seeking prior permission, provided they are not for re-sale. For further information about the Literati Club, please seewww.emeraldinsight.com/authors/index.htm. For further information about how to submit material to Soldering & Surface Mount Technology, please see www.emeraldinsight.com/smmt.htm