Laser Applications for Mechanical Industry

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Series E: Applied Sciences - Vol. 238

Laser Applications for Mechanical Industry edited by

S. Martellucci The Second University of Rome, Rome, Italy

A.N. Chester Hughes Research Laboratories, Malibu, California, U.S.A.

and

A.M. Scheggi IROE-CNR Florence, Italy

Springer-Science+Business Media, B.V.

Proceedings of the NATO Advanced Study Institute on Laser Applications for Mechanical Industry Erice, Trapani, Italy 4-16 April, 1992

Library of Congress Cataloging-in-Publication Data L a s e r a p p l i c a t i o n s f o r m e c h a n i c a l i n d u s t r y / e d i t e d by S. M a r t e l l u c c i ,

A.N. C h e s t e r , A.M. S c h e g g i .

p. cm. -- (NATO ASI s e r i e s . S e r i e s E, A p p l i e d s c i e n c e s ; v o l .

238) " P u b l i s h e d i n c o o p e r a t i o n w i t h NATO S c i e n t i f i c A f f a i r s D i v i s i o n . " I n c 1 udes i n d e x .

ISBN 978-94-010-4879-8 ISBN 978-94-011-1990-0 (eBook) DOI 10.1007/978-94-011-1990-0 1. M a n u f a c t u r i n g p r o c e s s e s — E q u i p m e n t and s u p p l i e s — C o n g r e s s e s .

2. L a s e r s - - I n d u s t r i a 1 a p p l i c a t i o n s — C o n g r e s s e s . I . M a r t e l l u c c i , S. I I . C h e s t e r , A. N. I I I . V e r g a S c h e g g i , A. M. (Anna M a r i a ) IV. N o r t h A t l a n t i c T r e a t y O r g a n i z a t i o n . S c i e n t i f i c A f f a i r s D i v i s i o n . V. S e r i e s : NATO ASI s e r i e s . S e r i e s E, A p p l i e d s c i e n c e s ; no. 238. TS183.L37 1993 670 . 4 2 - - d c 2 0 93-10592

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TABLE OF CONTENTS

Preface ............................................................................................................................................... ix

I. Laser Materials Processing: Tutorials

Topics in High Power Laser Processing .................................................... , ..... .................. 3 C. M. Banas

High Average Power Solid State Lasers for Materials Processing and Fiber Transmission ........................................ ........................................ 11

H. Weber

Laser Process Automation ................................................... ................ ........... ................... 31 W. M. Steen

H. Theoretical Foundations

Modeling of Laser Materials Processing ......... .................. ................... ....... ....................... 47 J.Mazumder and A. Kar

Analytical Model for Aberrated Diffraction in High Power C.W. Laser Beam Trains: Laser Cavity to Work Piece ........................................................................... 77

J. R. Palmer, W. M. Steen and S. Martellucci

Analytical Model for Evaluating Dual Co-Axial Repetitive Pulsed and C. W. Lasers in Drilling and Machining of Materials ................................ ...... 99

J. R. Palmer

HI. Laser Cutting and Welding Applications

Laser Machining of Composite Materials ......................................................................... 115 F. Jovane, A. Di Ilio, V. Tagliaferri and F. Veniali

Characteristics of Laser Heating of Materials .................... ,.............................................. 131 I. Yu. Smurov

vi

Laser Assisted Machining ................................................................................................ 151 I. Yu. Smurov and L. V. Okorokov

Heat Processes in Pulsed Laser Action ............................................................................. 165 I. Yu. Smurov and A. M. Lashin

Recent Development to Increase the Efficiency and Flexibility of Laser Materials Processing ............. ............ .................... .... ......... ..... ........ 207

H. Hugel and F. Dausinger

Laser Welding Technology for Joining Different Sheet Metals for One Piece Stamping .................... .............. ....... ............... ...... ....... ....... .... 219

Kazuo Azuma and Kimikazu Ikemoto

Laser Applications for 3·D Components: Beam Delivery Systems and Robotics .......... .............. ............ ... ......... ..... ...... 235

L. Pera, P. Perlo, E. Rabino and G. Marinoni

Laser Blank Welding and Stamping of Sheet Metal Parts .................................................. 263 F.A. Di Pietro

Robotic Laser Welding Systems in Automotive Operations ............................................... 271 F. A. Di Pietro

Robotic Laser Cutting Systems. ...... .... ................... .......... ........................... ...................... 277 F. A. Di Pietro

Development of Laser Material Processing in Romania ............. ........................................ 283 I. I. Far~as

IV. Other Applications: Stereolithography, Surface Hardening, 3-D Surface Patterning, Manufacturing and Repairing

The Role of the Laser in Rapid Prototyping .................. ............ ...... .......... ......... ............... 293 L. Pera and G. Marinsek

Laser Surface Melting of Bearing Steels .......................................................... ................. 305 R. Cola~o and R. Vilar

Laser Beam Lithography for 3-D Surface Patterning ......................................................... 315 C. Arnone, C. Giaconia, C. Pace, S. Bonura and M. Greco

Laser Techniques in Power Plant Component Manufacturing and Repairing .................... 321 W. Cerri, L. Garifo and D. D'Angelo

vii

V. Laser Measurement Techniques

Video Speclde Interferometry:. An Optical Measuring Tool for Industry ......................... 327 O. J. Lgkberg

Adaptive Profilometry for Industrial Applications ...... ............ ......................................... 351 G. Sansoni, F. Docchio, U. Minoni and L. Biancardi

Electrooptical Systems and Techniques for Dimensional Measurements for Industry .............................. ............................................ 365

F. Docchio, U. Minoni, G. Sansoni and E. Gelmini

Laser Velocimetry for Combustion ....... .......... ............ .............. .......... ............ ................ 381 D. F. G. Durao and M. V.Heitor

Industrial Applications of Holographic Interferometry ................. .................................. 403 N. H. Abramson

VI. Laser Safety

Laser Safety Devices for High Power C02 Laser Systems .............................................. 415 P.Gay

VII. INDEX

Index .................................................................................................................................................. 429

PREFACE

The NATO Advanced Study Institute "Laser Applications for Mechanical Industry" was

held April 4-16, 1992, in Erice, Sicily. This was the 17th conference organized by the

International School of Quantum Electronics, under the auspices of the "Ettore Majorana" Center

for Scientific Culture.

The present Proceedings give a tutorial introduction to today's most important industrial

applications of lasers, as well as a review of current results by leading researchers. We have

brought together some of the world's acknowledged experts in the field to summarize both the

present state of laser technologies and their background. Most of the lecturers attended all the

lectures and devoted their spare hours to stimulating discussions. We would like to thank them

all for their admirable contributions. The Institute also took advantage of a very active audience;

most of the participants were active researchers in the field and contributed with discussions and

seminars. Some of these seminars are also included in these Proceedings.

We did not modify the original manuscripts in editing this book, but we did group them

according to the following topics:

1) "Laser Material Processing: Tutorials" ;

2) "Theoretical Foundations" ;

3) "Laser Cutting and Welding Applications" ;

4) "Other Applications: Stereolithography, Surface Hardening, 3-D Surface Patterning, Manufacturing and Repairing" ;

5) "Laser Measurement Techniques" ; and,

6) "Laser Safety" .

ix

x

The six sections provide a broad review of laser applications in modern industry by taking

some relevant examples from the automotive field where lasers have shown important

potentialities in conjunction with the technologies of optical fibers, automation, computing and

flexible manufacturing. These chapters, and the further references therein, should form a useful

guide to today's industrial laser technology, and the basis for future advances in those areas

where flexibility, automation, CAD/CAM integration, precision, cost reduction and time to

market are important factors.

Before concluding, we acknowledge the invaluable help of Luciano Pera, from FIAT

Research Centre, co-director of the conference. We are also grateful to Eugenio Chiarati for

much of the computer processing work. We also wish to mention with sincere thanks Margaret

Kyoko Hayashi, Maria Teresa Petruzzi and Vanna Cammelli, secretaries to the AS.!. Directors

(A.N.C., S.M. and AM.S., respectively). Finally, we acknowledge the organizations who

sponsored the conference, especially the generous fmancial support of the NATO AS.!.

Programme Committee.

The Directors of the A.S'!. :

Sergio Martellucci

Professor of Physics

The University of Rome "Tor Vergata"

Rome (Italy)

Arthur N. Chester

Vice President and Director

Hughes Research Laboratories

Malibu, California (USA)

Anna Maria Scheggi

Director of Research

CNR Electromagnetic Waves Institute

Florence (Italy)

February 15, 1993

LASER MATERIALS PROCESSING: TUTORIALS

TOPICS IN mGH POWER LASER PROCESSING

C.M.BANAS Chief Scientist United Technologies Industrial Lasers 300 Pleasant Valley Road South Windsor, CT, USA 06074

ABSTRACT. Four topics relating to the materials processing performance of multikilowatt lasers are discussed. These include: 1) the effect of laser cavity and focusing optics on welding performance, 2) the effect of reduced pressure on laser welding performance, 3) the influence of shorter (than carbon dioxide) wavelength on welding behavior and 4) a brief review of some successful production applications.

1. Introduction

The first successful operation of a laser occurred more than thirty years ago. Since that

time both the available laser equipment and its applications have undergone extensive

development. Early solid state lasers were followed rapidly by gas lasing systems; the

workhorse of the continuous multikilowatt range, the carbon dioxide laser, was initially operated

in 1964. The first version of the latter was a slow flow, coaxial device in which output power

was, typically, 50 W per meter of discharge length.

fu the late 1960's convectively cooled fast axial flow and transverse flow carbon dioxide

lasers were developed. Although the specific design and means of excitation may differ, these

two basic designs represent the bulk of the systems used in production today.

Although the general features of multikilowatt carbon dioxide lasers are similar, it is

important to note that the processing perfonnance depends upon the characteristics of the output

beam itself and the environment in which processing is accomplished. It is to these latter factors

that the following is addressed.

3

S. Martellucci et ai. (eds.J, Laser Applicationsfor Mechanical Industry, 3-9. © 1993 [(luwer Academic Publishers.

4

2. Effect of cavity and focusing optics on welding performance

At power levels above five kilowatts, unstable resonator cavity optics are customarily used

for beam extraction. This occurs since appropriately-durable, partially-transparent window

materials are not available for carbon dioxide, and dimensions associated with Gausssian beams

necessitate multiple path cavities for attainment of good extraction efficiencies

Unstable oscillator cavity optics may be designed for positive branch (one concave and

one convex mirror) or for negative 'branch (two concave mirrors) operation [1]. Extraction

efficiency is the same for both, however, the negative branch yields a focal point within the

cavity which may cause internal cavity plasma foonation at very high power. For this reason

most of the high power systems in use today use the positive branch configuration.

Irrespective of the branch used, the focusability of the output beam is characterized by the

magnification, i.e., the ratio of the outer-to-inner diameters of the beam. Focusability increases

with magnification (M) approaching that of a plarie wave at an M of infinity. Unfortunately,

high values of M yield little power; maximum power output occurs in the range foon 1.5 to 2.0.

For this reason, M=2 optics, which represent a compromise between output power and

focusability, are often used. A higher M may be used ( with resultant decrease in maximum

output power) in circumstances requiring special weld properties.

2.1 MAGNIFICATION, INTENSITY AND WELD CHARACTERISTICS

It is well known that the onset of deep penetration, keyhole, welding occurs at

approximately a half million watts per square centimeter. This value can, in fact, be estimated

using a simple theonal diffusion model. This involves deteonination of the intensity required to

cause vaporization at the interaction point while the temperature at a distance one theonal

diffusion length into the material remains at its ambient value.

The influence of power intensity on keyhole welding characteristics has been well

documented in electron beam experiments. e.g. [2]. We refer to these results here because they

are bereft of possible influence of plasma breakdown which may occur with laser welding. It is

noted that the bead characteristics change from the roughly hemispherical profile of a typical

fusion weld at power densities below 5x106 W/cm2 to a high depth-to-width penetration at

high intensity. Between these limits the welds comprise a central zone dominated by keyhole

effects and an outer zone controlled by theonal diffusion resulting in a wineglass-shaped weld.

5

To explore this effect with the laser, comparative tests were performed with M=2, 3 and 4

cavity optics. With f/9 external focusing optics, this provided intensities ranging from 1.1 to 3.5

xlO 6 W/cm2 at the focal point with lO kW oflaser power. Weld penetrations in 9.5 mm thick

steel performed under these conditions led to formation of a wineglass-shaped weld at the lower

intensity and to a parallel sided keyhole weld at the higher end. It is underscored that this is the

same behavior as that obtained with electron beams and stems from the direct thermal response

of the material to the high intensity beam and not to plasma formation. This is all the more

evident because an increase in intensity would be expected to cause a more severe plasma

breakdown and result in a more pronounced wineglass; the converse was observed. A further

illustration is provided by observation of the influence of a plasma that has been "pushed into the

keyhole" [3]. An optically dense plasma within the workpiece inhibits penetration and results in

a local bulbous bead profile at the plasma location. The bottom line is that plasma breakdown

may sometimes be incorrectly blamed, or given credit for, many effects.

2.2. DEPTH OF FOCUS AND SPEED EFFECTS

Depth of focus of the external optics, as well as spot diameter, has a significant effect on

welding performance. Extremely small spot sizes, obtained with short f/number focusing optics,

provide excellent performance in very thin materials. As material thickness increases, however,

the performance degrades and improved performance is obtained at a higher f/number.

Comparison of results with fixed focus optics shows that best welding performance is

obtained when the depth of focus matches the thickness of the material provided that power

intensity can be maintained. The latter condition, of course, requires increase in power with

material thickness. Ideally, a zoom lens would be useful for a multipurpose laser welding

machine thereby permitting selection of optimum f/number for each application.

Increasing power results in increased processing speed; however, this increase can not

effectively be utilized due to the onset of a fluid dynamic instability of the weld pool. The result

is an irregular surface bead profile and erratic penetration which has been dubbed "humping" [4].

In parallel to behavior of high speed surface ships. it has been shown that the threshold speed for

humping can be delayed to higher levels if the weld pool length is increased.

One means for increasing weld pool length is use of a twin beam approach in which two

keyholes. one behind the other, occur within the same weld pool [5]. This has resulted in a

modest increase in the threshold for humping; additional effort is. however. required in order that

full advantage can be taken of the high speed welding capability of multikilowatt lasers.

6

3. Effect of atmosphere on welding performance

Laser welding in a reduced pressure environment has been shown to increase the

maximum penetration capability for a given laser [6, 7] The penetration obtained increases with

decreasing pressure until a plateau is reached at about 10-2 Torr. It is noted that this behavior

also characterizes electron beam welding. The similarity between the two processes is somewhat

surprising in view of the fact that beam scattering is responsible for the decrease in electron beam

performance as pressure increases and plasma formation results in the laser's performance

degradation.

The reasons underlying this behavior may be surmised by reference to a phase diagram of

the material being welded. It is found that the triple point for materials such as iron, nickel and

chromium falls at approximately 10-2 Torr. Because of this, material heated by a power beam in

vacuum reaches the vaporization point before the melting point. Were it not for a local increase

in pressure within the keyhole itself, welding would not be possible in vacuum since the liquid

phase is non existent at low pressure levels. The decrease in vaporization temperature with

environmental pressure eases the difficulty of providing the vapor keyhole essential to deep

penetration welding. For this reason the keyhole is stable at much lower speeds than it is in

atmosphere. In conformity with normal operating experience, which demonstrates an inverse

relationship between speed and penetration, decrease in keyhole welding speed leads to

significant increase in weld penetration. Once again we have a situation in which the basic nature

of the material being welded and the process characteristics determine the welding behavior.

4. Influence of short wavelength on welding performance

One of the drawbacks of the carbon dioxide laser for processing of metals is its relatively

long output wavelength. Shorter wavelength has a number of potential advantages:

1) Improved focus ability

2) Decreased tendency for plasma formation

3) Inexpensive, durable transmitting optics

4) Possibility of fiber optic transmission

5) Improved initial coupling with metals

7

In order to obtain some insight into the advantages offered by shorter wavelength in the

multikilowatt regime, a series of tests was conducted with a 3.8 micron wavelength, DF,

chemical laser system [8]. Power was varied from 8 to 20 kW and weld penetration was

determined as a function of speed for ten common metals. Although the DF laser is not, and will

most probably never be suitable for production use, the results of the tests provide a basis for

evaluation of the potential benefits of shorter wavelength.

In the DF tests it was found that weld penetration and speed for the range well within the

keyhole regime (typically at speeds above 2 m/min) were comparable to those for carbon dioxide

in materials such as steel, stainless steel, titanium alloy and nickel alloys. Marked improvements

in weldability were noted, however, for the characteristically high reflectivity materials such as

copper and the aluminum alloys. It was also noted that the difficulty of plasma suppression was

significantly reduced at the shorter wavelength.

s. Production applications

5.1. AEROSPACE PRODUCTS

The evolution of production applications at power levels above 5 kW has been taking

place at a slow, but increasing, rate. One unusual application of a 6 kW system is piercing of

holes in a jet engine combustion chamber liner [9].

Cooling holes in this element are about 2 mm in diameter, about 3 mm deep and may

number as many as IO,OOO in a single unit. Low power pulsed laser trepanning produces a more

uniform hole but requires approximately 30 seconds per hole. This translates into a production

time of the order of 80-90 hours. Electric discharge machining with multiple electrode units

requires comparable times. The 6 kW laser system, however, pierces these holes at the rate of

two per second thereby reducing processing time to less than three hours. Hole quality, which is

determined by flow measurement, is well within specification. More than thirty million holes

have been pierced in this, one of the most cost-effective applications in production.

Another task on the same combustion chamber liner is assembly welding of the individual

rings which make up the chamber [9]. This requires a 1.8 mm deep weld between a 1.2 mm

thick segment of nickel superalloy sheet metal and a roll formed part. Constraints on the width

of the weld dictate use of a beam approach; the laser was found to be much more convenient

8

than electron beam for this application and has been in successful production use for more than

nine years.

5.2 AUTOMOTIVE COMPONENTS

The largest production use of lasers above 5 kW is welding of automotive components

(Ref.10). Such applications generally involve welds of approximately 4-5 mm depth with

requirements for minimal specific energy input. Because of high volume production

requirements, welds, which are of the order of 25 em in length, must be formed in less than ten

seconds.

Typical of the production applications in automotive components are air conditioning

elements and transmission parts. In manual transmissions, enhanced focusing optics are used in

order to provide narrow, parallel-sided welds with reduced thermal distortion.

A developing area for high power automotive applications is the remelting of camshafts to

yield a wear resistant surface. With appropriate preheat, a defect free layer of Ledeburite can be

formed to controlled depth on the wear surface of the cam. Individual cam processing times are

of the order of 5 seconds at 12 kW for a nominal I mm deep remelt layer. In this application, of

principal interest in Europe, the laser would replace gas tungsten arc processing.

6. Concluding remarks

High power lasers are finding an increasing role in high volume production in industry.

Significant increases in production use are anticipated within the next decade.

References

1. Siegman, AE. (1986): "Lasers". University Science Books, Mill Valley, CA. 2. Meier, lW. (1966): "Electron beam welding at various pressures". Metallurgical Society

Conference, 51(1): 505-533. 3. Banas, C. M. (1976): Paper presented at the SME Western Lasers in Manufacturing Conference. 4. Tsukamuto, S. et aI. (1983): "Effect of Focal Position on Humping Bead formation in Electron

Beam Welding". Transactions of National Reserch Institute for Metal, Vol. 25, No.2. 5. Banas, C. M. and Nuss, R. (1990): "Materials Processing with High Power CO? Lasers·,

Proceedings of the 3rd European Conference on Laser Tratment of Materials, Erlanger, Germany.

9

6. Brown, C. O. and Banas, C. M. (1986): "High Power Laser Beam Welding in Reduced Pressure Atmospheres", Invited Presentation-ICALEO 1983, (Published in AWS Journal, July 1986).

7. Arata, Y., Abe, N. and Oda, T. (1984): "Fundamental Phenomena During Vacuum Laser Beam Welding in Reduced Pressure Atmospheres", Paper presented at ICALEO 1984, Boston, MA.

8. Banas, C. M., Alholm, H. and Olihan, W. (1983): "Multikilowatt Laser Welding at 3.8 Microns", paper presented at Golden Gate Welding Conference, San Francisco.

9. Banas, C. M. (1991): "High Power Lasers in Manufactoring". Paper Presented at ASME International Gas Turbine & Aerospace Congress, Orlando, FL.

10. Ogle, M. and Gustaferri, D. (1989): "Why Laser". International Conference on Laser Processing, Welding Institute, Cambridge, England.

HIGH AVERAGE POWER SOLID STATE LASERS FOR MATERIALS PROCESSING AND FIBER TRANSMISSION

H.WEBER Festkorper-Laser-lnstitut Berlin GmbH Optisches Institut, Technische Universitiit Berlin, Germany.

ABSTRACT. The different high average power solid state lasers in the kW-range are compared with respect to their suitability for material processing. The problems of fiber transmission are discussed.

1. Introduction

Two laser systems are most often used in material processing, the Nd-Y AG (Glass)-Laser

(A = 1.05 - 1.06 J.IlIl) and the C02-gas laser (= 10.6 J.IlIl); see Table 1.

Table 1. Comparison of C02 and Nd-YAG lasers.

CO2

power,kW 15 - 25 (40)

beam quality up to 500 W beam quality hip;h power ~ood DOwer dependent variation of beam quality low efficiency, % 10 -20 beam handlinp; ri~id mirror systems pulsed mode limited wavelenp;th 10.6 11m lifetime high

runninl! cost costs of investment costs of beam handling

(a) higher absorption in metals

11

S. Martellucci et al. (eds.), Laser Applications/or Mechanical Industry, 11-29. © 1993 Kluwer Academic Publishers.

Nd-YAG

2-3 (5)

comparable low strong <5 fiber hil!h enerl!V short pulses 1.06 11m (a)

"" 400 h (lamps) comparable

more expensive I cheaper

12

Fig. 1. Increasing output power by using several lasers in parallel, but beam parameter product increases proportional to the number of lasers (IK Lumonics). Total power P=N PO' Irradiance single system Io= PoI02Ao ' and, Irradiance N systems I~ N PoI02N Ao ' where: N systems; Po output power of single laser; and, AD emitting area of single laser.

From the broad variety of solid state lasers, only two are of interest in the field of material

processing, the Nd-Glass and the Nd-Y AG laser. The Nd-Glass can be used as a pulsed laser,

mainly for spot welding. However, if we consider high average power systems in the kW-range,

only Nd-YAG presently qualifies. Therefore the discussions in the following sections will refer

to Nd-YAG.1t is no problem to produce several kW with a Nd-YAG laser: one may take three

systems with I kW each, and use them in parallel as shown in Fig. 1. This looks quite simple,

but we must remember the second law of thennodynamics, which in this case for uncorrelated

lasers implies: "The i"adiance of a light source, that is the power per area and solid angle,

cannot be increased by any passive optical system". In brief: the beam quality of this high power

laser becomes three times lower than that of the single system, and it can be only used for very

special applications. That is the main problem of solid state lasers: to realize a kW-system with

high beam qUality. To understand the technical background of this problem, we have to discuss

"beam quality" in more detail.

ZR

Fig. 2. Beam quality is defined by the beam parameter product wo·eh = 1/4 do e where d is the waist diameter and e the full farfield angle. zR is the Rayleigh length or depth of focus.

13

2. Beam quality

Assume a laser system with a plane outcoupling mirror as shown in Fig. 2. The beam

parameter product is defined by the product of the half far field divergence 8tt and the waist

radius Wo or by the full angle e and the diameter do:

Beam parameter product BP = woo 8tt = do·e/4

The International Organisation for Standardization ISO has proposed to measure both parameters

by the 86,5 % power content area [1]. This beam parameter product has its minimum value if the

laser operates in the fundamental mode TEMOO (Gaussian beam): ( 1/4 do·e)oo = AJrc with A. the

laser wavelength. For all other beams (higher order modes, distorted beams) the product is larger.

If the beam parameter product is normalized relative to its ideal value, a dimensionless number is

obtained:

(1)

Often the parameter M2 is used, which is defined by the normalized beam parameter product of

the second moments [2] :

(2)

It is approximately true that:

(3)

In Table 2 typical values of the beam quality factor K are compiled. Note that K is equal or

smaller than one, and that in passive ideal optical systems K is a constant. Beam quality can be

improved (increasing K) outside the laser only by limiting apertures, which reduces the power.

14

Table 2. Beam parameter products and K-values of different laser systems.

Laser Power (wo'f)h)oo wO'f>b K W

mm 'mrad mm'mrad HeINe 5x10-3 0.2 0.2 1 Nd-YAG 15 0.34 0.36 0.94

90 2.25 0.15 1000 30 0.012

NdGlass 15 0.34 28 0.012 directly coated 10 0.034 CO2 500 3.2 4.2 0.76

1500 8 0.4

The beam quality or beam parameter product is of major importance for several reasons.

a.) ~.dQl!!l zR. The focal depth or Rayleigh length is the distance behind a focus (waist) at

which the beam area has doubled, or the beam radius has increased by "2. From laser textbooks

[3] we can take the relation:

(4)

with df the focus diameter and K the above mentioned quality factor. Low beam quality means a

small value of K and a short Rayleigh length, which is disadvantageous for material processing

(conical shaped beam, positioning). An example is given in Fig. 3.

LASER

HIGH ORDER IDlE

TELESCOPE

W. 'h~ [)-: ; : JC-W'

2R

Fig. 3. Focussing of a Gaussian beam (K = 1 ) and a high order mode ( K < < 1 ).

15

FigA. A fiber is characterized by its core diameter 13 and its numerical aperture NA.

b.) fiber transmission. A fiber is characterized by its core diameter ~ and the numerical aperture

NA, which depends on the refractive indices of core and cladding (Fig. 4). Low loss coupling of

a laser beam into a fiber requires approximately:

BP <cpNA/2 orK> 21../ 1tcpNA (5)

A fiber used for high average power transmission has typically a numerical aperture of NA = 0.2

and a core diameter of ~ = 0.6 mm. Eq. 5 implies for the beam parameter product of the laser:

BP < 60 mm mrad, which is an upper limit. To avoid fiber damage, the beam parameter product

of the laser beam should be smaller by a factor of two than the fiber product.

3. Requirement for industrial applications

Table 3 is a rough summary of the requirements for material processing. These data

depend strongly on the special process, material, precision, and so on.

Table 3. Requirements for material processing. The beam quality of the fundamental mode Nd­Y AG laser is 0.3 mm mrad.

power, kW 0.1 - 5

beam quality, mm mrad

efficiency, % lifetime, h fiber transmission core diameter, 11m Dulse width repetition rate

0.5 microprocessing < 10 cutting < 20 fiber transmission >5 > 1000 <600

cw, 10 ms, 100 ns 10 Hz - 100 kHz

16

The processing velocity is detennined by the power, the beam quality and the mode of

operation (pulsed, cw). Several kW are required for high cutting speed (cw) and for surface

treatment (pulsed).

The beam quality depends on the application. Microprocessing requires a very high beam

quality, fiber transmission a lower value. The laser efficiency should be about 4 %. Lower

efficiency means more electrical power and large cooling units. The life time is limited by the

lamps. 400 h life is possible now with selected lamps,lOOO h is within reach. Laser beams of low

beam quality can be transmitted by fibers with larger core diameters (1-2 mm). However, to

permit small bending radii and small focii with nonnal cutting optics, a core diameter of 200 -

400 JlIll would be desirable. Low transient effects mean that the laser output power can be varied

between 100 W and 2 kW without variation of beam quality and focus diameter. This is

important for on-line control and flexible manufacturing.

'''''' __ ~''*'-__ -t--+-_-TEMOO -': , # ~----~~----~~----~

#

o « lid

TEMOO §TEMmn

0= lid

TEMOO

O.2/d

Fig. 5. A pumped rod becomes a thick lens and affects the beam quality.

17

4. kW-Solid state lasers

4.1 BEAM QUALITY OF SOLID ST ATE ROD LASERS

Solid state lasers in the kW-range are pumped by gas discharge lamps. The overall

efficiency is about 3 - 3.5% (plug), and about 50% of the electrical power is dissipated as heat

inside the laser rod. This heat has to be removed by water cooling and results in a temperature

gradient inside the rod. This parabolic gradient from the rod center to the outer region produces a

thermal lens, due to the temperature dependent refractive index and the internal stress. The latter

moreover causes circular birefringence. For a Nd-YAG rod of 10 mm diameter, the following

relations hold approximately for the two refractive powers (radial and azimuthal polarisation):

Dr=0.3m- I/kW

Dd = 0.27 m -1/kW

where kW refers to the electrical input power of the lamp [4]. A rod of 20 cm length with an

output power of 400 W becomes a thick lens with a focal length of about fr = IlDr = 33 cm, and

influences the beam quality strongly.

As shown schematically in Fig. 5, at a low pumping level the rod-lens stabilizes the plane­

plane laser resonator. The system operates near fundamental mode. With increasing pumping

power the fundamental mode is reduced in diameter and focused on the mirrors. Higher order

modes start oscillating, and the beam parameter product grows. If the pumping power is still

increased, the spots on the mirrors become small, in such a way that by diffraction the beam­

diameter inside the rod is enlarged. The mode order is reduced as well as the beam parameter

product. Finally, the system should oscillate in fundamental mode again and will then become

unstable. In the stability diagram the resonator moves with increasing refractive power on a

straight line as plotted in Fig.6.

The most used system, the symmetrical plane-plane resonator can be described by rather

simple relations [4, 5]. The resonator is stable, if the refractive power D is in the range

~ -2 DL r==z 1 - 1- F ~ - ~ 1 + V1-F -, 2

(6)

where L is the resonator length and R the radius of the limiting aperture. In this range the beam

parameter product reads:

18

/~?////'t! /, " :///92 2 , 'unstable ,./ / . / '/ ~, '

/

/ /'

conuntric .

////~, , . unstable . ' // /;: " .'

/ ,/ ,'-' / '/ .'..'.,

/%-//"/

.'

Fig. 6. With increasing pumping power the spherical rod resonator moves on straight line in the stability diagram.

2 BP=~{DL[l-DL/ 4]}1/2

L

The largest product is obtained for DL = 2 with:

The range of stable operation is [8]:

dD=4/L

(7)

(8)

(9)

In most cases the refractive power range is proportional to the pumping range d Pp , and this is

proportional to the output power range. This leads finally to a very useful relation:

(BP)max/dPout = const.

This relation depends on the specific crystal type only and holds for all spherical resonators, If

the BP is low, which means high beam quality, the stable range of output power is small and

vice versa. For Nd-Y AG:

(BP) / Llli = 100 mm· mrad max out kW

19

3

bJ IJ· .. C2 ... ,-~ ) ..fdrl-d2~

fie 2

~

...........

i1~

0

O/lm-1)

Fig. 7. Beam divergence 8m versus refractive power D of the rod in the above configuration. Rod: LG

706,6 x 100; dl = 33 mm; PI ::: 00 ; P2 = 00 (D), 5 m (0), 2m (.Cl); d2 = 150 mm (a), 300 mm (b) [5].

A 500 W plane-plane Nd-Y AG laser will have a beam parameter product of about 50 mm mrad.

This holds for ideal systems only. For several reasons fundamental mode operation at high

pumping power level is impossible: spherical abberation, inhomogeneities, non homogeneous

pumping light distribution and so on. A typical dependence of the beam parameter product on

pumping power is shown in Fig. 7 [5]. This effect is unavoidable in rod systems without any

passive or active compensation, and up to now no reliable compensation exists.

Another approach is the slab laser as shown in Fig. 8. The slab with Brewster end faces is

side-pumped and side cooled. Again a temperature profile appears. However, the laser beam

inside the slab propagates in a zig-zag fashion, and thennal effects are compensated to first order.

Other problems arise, which will be discussed later. Anyway, the slab produces a laser beam of

high quality, which is very little influenced by varying pumping power.

- .------ --- -:~ - . .. . . ..

-- - - -_ .. ---- ---

Fig. 8. The slab-laser.

20

........ ---101 ~d I. d .f h-Id dZ 3 , SmI

rror ·2

Fig. 9. The multirod-oscillator.

Scraper

Fig. 10. Three types of unstable resonators for rod lasers.

~'T---------------------~ "''''.1_ ... Nt.& lOIN' ..... _710\/

¥ ~

400 t-.. A-

i .. 200 ... ~

........ -PIan ICIOY

2 • PUIIPII6 POWER

• 8 10

/Kovitaet [leW)

130

'1 120

E i 10

:;:

~ O'+---~--~~~--r-__ -r __ ~ o 2 •

PUIIPIN6 POMER •

/Kovitaet [leW) 10

Fig. 11. Output-power and beam parameter product of an unstable resonator, as a function of pumping power. Resonator 1: PI = -0.5 mm, P2 = -0.5 m, L = 1 m, central lens with f = -0.3 m, two rods Resonator 2: PI = +0.5 m, P2 = -0.5 m, lens f = -0.5 m.

21

4.2 THE MULTI ROD SYSTEM

The Nd-Y AG rod-laser is a well developed, reliable system which is available at a high

technological level. Y AG crystals can be grown up to a length of 200 mm and 8 - 10 mm

diameter. Such a crystal delivers a maximum output power of around 500 W. This upper limit

result from fracture due to thermally induced stress, and holds for lamp pumped systems. Diode

pumped Nd-Y AG rods can deliver up to 50 W/cm, which in principle could result in an output

power of 1 kW for such a rod. The fracture of a crystal under stress depends strongly on the

surface properties. To avoid fracture in lamp pumped systems, the output power will be limited

to about 400 W per unit.

The output power can be increased by using several rods inside one resonator (Fig. 9). If

these rods are arranged in a suitable manner, the beam parameter product is more or less the

same as that of the single unit. Powers of more than 2 kW have been produced in the lab. If this

multirod system is operated with a normal plane mirror resonator the beam parameter product is

about 40 mm mrad.

Another approach is the unstable resonator, shown in Fig. 10. This system has a rather low

beam parameter product, about 3 to 5 times diffraction limited. Its disadvantages are the slightly

reduced efficiency and the moving focus, which requires adaptive optics. The output power of a

two rod system amounts to about 600 W. Typical characteristics are shown in Figs. 11 to 13.

4.3 SLAB-LASERS

The thermal distortions of slab-lasers are one magnitude lower than those of rod lasers,

and due to its rectangular shape a slab can be pumped more intensely than a rod.

EJOO " oS " .. ..

2!50 • • * ! • 0 0 0 0 .... ~200 0

0

0- 0

;;; 150 0 0

" 0 ... CI) 100 ::> ... 0 .50 .... 0 Mr.1 (1.0) . :~~~l

0 0 600 laser output power

Fig. 12. With increasing pumping power the focus behind a lens with f = 0.2 m moves.

22

It hIaIoII, L-58cm • pion 1Iam-. L-74cm • pIon-p/an, L-74cm .. pIon-p/an, L-55cm

.. .. • .. • .. • • •

.. • •

Output pover

• •• +

(Watt).

Fig. 13. Beam parameter product vs pumping power for various systems.

However, special resonators have to be used for obtaining a high beam quality (Fig. 14).

Moreover, the technical realization of a slab-laser is rather sophisticated. Careful design of the

cooling is necessary, as the inference fringes in Fig. 15 demonstrate. These requirements result in

a very expensive system. Nevertheless, excellent results have been obtained with slabs [9].

Fig. 14. Special resonators for slab-laser [12]

23

STlAI6IIT ON TRANSIIISSION 0 til PIII'EI! SURFACE

THERMAL ISOLATION

STUI6IIT ON TRANSMISSION 10 til ZIG-ZAG TRANSIIISSION 10 til

Fig. 15. Interference structure of slabs.

4.4 ANNULAR RODS

The annular rod or tube (Fig. 16) is the system of the highest efficiency. so far. It is a very

compact system. A 1.5 kW laser requires a tube of about 55 mm outer diameter and 200 mm

length. The big disadvantage is its low beam quality. as can be seen in Fig. 17 [10].

4.5 OTHER SYSTEMS

A number of other configurations have been proposed and partly investigated in the lab:

moving slab; -disk; -rotating tube; -fiber. The principal arrangements are summarized in Fig.18.

The most interesting approach is the diode-pumped slab (Fig.19) [11]. The flashlamps are replaced by laser diodes with a pumping wavelength of about AP = 0.809 1J.IIl. This is not very far

from the laser wavelength of AL = 1.06 1J.IIl. and internal heat production is therefore strongly

reduced.

200 t _ 4 .... : •

f - 106 HI: "_,,, Ii" ., .................. . . . .

. . . .. .. : .... : ... tt·-.& ....... -.. : ...........

O'+-~~~~~~~ __ ~·"-~·-~·~·~ 0.0 1.11 z.o U

Electrical Pumping Power [kW]

Fig. 16. The annular rod laser with the output characteristics.

24

..... ~

f e .

• ROD c SLAB

o

o

•

o

• •

1:---__ ....;:.._. _________ Id!!L.-,I )"1Tr.

~~~~~~~~~~~~~~~~~ 0.01 0.1 1 10

OUTPUT POWER (kWl

Fig. 17. Beam parameter product vs output power for various systems. Fundamental mode product for

1.06 ~m is do·e/4 = A!1t.

)

Fig. 18. Other configurations: Rotating tube, moving slab, disc, fiber-laser.

25

This leads to lower thennal distortions, higher efficiency, and better beam quality (Fig. 17).

However, at present, diodes are much too expensive. Although a I kW system is running in the

slab, a commercial system at a moderate price will not be on the market within the next five

years.

5. Aspects of high average power transmission by fibers

At a wavelength of AL = 1.06 !lID absorption and scattering losses of quartz/glass fibers

are neglible. Transmission of kilowatts for distances of a hundred meters and more are no

problem. Internal damage threshold and nonlinear effects appear at intensities of GW/cm2, which

means several MW for 5()() !lID fibers. Limitations arise from the beam quality and the lower

damage threshold of the entrance surface. Laser light reflected from the material surface may also

produce damage or distortion of the laser beam. Two different types of fibers are used, the step

Fig. 19. Diode pumped lasers: end-on rod, side-on rod, side-on slab.

26

i -.

.-----'1-----. ""

• (rl

o •

i ..

Fig. 20. Step index (upper picture) and gradient index fiber (lower picture) with beam propagation.

index and the gradient index fiber - as shown in Fig. 20. The propagation of light is different in

both, which leads to the following effects:

~ index fiber: If at the input the laser spot diameter is smaller than the core diameter the

output spot is equal the core diameter, whereas the angle e remains constant if the fiber is not too

long (several meters).

Gradient index fiber: This fiber corresponds to a lens line and has imaging properties. A smaller

input diameter of the laser beam is reproduced partly at the output, whereas the divergence at the

output more or less corresponds to the NA of the fiber. The real fiber looks a bit more

complicated, if the coating is taken into account. Two examples are given in Fig.21.

leoat1D9Aazylat I . I 'UDDING I

I CORE I

[l]

mL !

1 eoat1D9 sUJDon I I CLADDING I

I CORE I m ~551,..-'

i

,~ tiL I

Fig. 21. The more complicated structure of real fibers. Step index (left) and gradient index (right).

c o

1.00,----------------. G-E) f"oser mit Acrytot-c:ooting !3-£J f"oser mit SlIic:on-Acrytat-coating

0.97

'iii 0.94 (I)

·E (I)

c ~

0.91 t-

0.88

o 1000 FIBER IIf'IlT POHER IW

Fig. 22. Fiber transmission vs input power.

27

The transmission of the fiber itself is nearly 100 % « 50 m) (Fig. 22), and losses occur

mainly at the input coupling due to mismatching and reflection. A well matched fiber will have a

transmission of about 92 %. If the previously mentioned defmition of the beam parameters is

used, only 86.5 % of the total power is coupled into the fiber. 14.5 % hits the cladding and will

be guided in special fibers only (double step fibers, in which a second coating has lower

refractive index than the cladding). Of a one kW-beam 145 W would be coupled into the

cladding and would destroy the fiber in most cases. Therefore another definition has to be used.

Experimental investigations reveal that the 98% values are a good approximation, and Eq. 5 has

to be replaced by

~d9~8~' 8-,,9~8 < _cI> ·_N_A . 4 2

This means for the Gaussian beam: d98' 898 = 2 do ·8; and requires a much lower beam

parameter product. If the above relation is met, high transmissions up to 92 % and more in the

kW range are obtained.

The main problem of fiber transmission is laser beam qUality. Of course, a larger core

diameter would allow higher power transmission with low beam quality, and fibers up to 2 mm

diameter are used. However, as already discussed, for a step index fiber the output core is more

28

or less completely illuminated. A small spot on the material surface requires reduction by optical

systems. For technical reasons (working distance, aperture, aberrations, dimensions, weight) the

reduction is limited to a factor of 2.5 to 3. Therefore, small core diameters below 0.6 mm are

desirable.

6. Summary

Nd-Y AG lasers in the multi rod configuration with 2 kW average power in cw and pulsed

mode (0.1 ms - cw) are available now for materials processing. The laser radiation can be

transmitted by fibers of about 600 IlIIl core diameter. Telescopic imaging delivers spot sizes of

250 - 300 IlIIl diameter with a focal length in the millimeter range. Below 1 kW average power,

fibers of 400 IlIIl core with smaller spot sizes can be used. Slabs lasers produce powers up to 1

kW with high and constant beam qUality. However, their reliable and technical realization is

more expensive. Compact tube lasers with high efficiency deliver 1.5 - 2 kW in the pulsed mode

(1 -10 ms), but with a rather poor beam qUality.

A review of systems under investigation is given in Table 4, and the beam parameter

products obtained are summarized in Fig. 17.

Table 4. State of the art for different high power systems

System Output Market Lab Planned Scaled Problems power (kW)

Y AG Multirod in series 1 - 2 1989 limited beam quality 3 1992

Y AG Multirod parallel 1-2 1989 very low beam aualitv

YAGSlab 0.5 - 1 1990 0 rectangular beam profile

Moving Slab Glass 1 0 mechanical 1.6 0 problems, lar~e volume

GOG-Slab 0.7 0 thermal problems 1 0

YAGtube 1 0 large volume, 2 0 homogeneity

Ftotating glass tube 12 - - 0 large volume Disk-laser 10 - - 0 gas cooling Fiber laser 0.4 0 scattering losses Fiber bundle 0.007 0 beam quality

29

ACKNOWLEOOEMENTS. These investigations were generously supported by the BMFf

(Bundesministerium flir Forscbung und Technologie) under various contracts.

References

1. Document ISO/TC 172/SC 9/WGl, International Organization for Standardization Terminoloy and Test Methods for Lasers. 2. Working draft "Test method for width, divergence and radiation chracteristicfactor of laser beam". Febr. 1991.

2. Siegmann, A. (1990): Proc. SPIE 1224, 2. 3. Yariv, A. (1975): "Quantum Electronics", Wiley and Sons, N.Y. 4. Hodgson,N. and Weber,H. (1992): "Optische Resonatoren", Springer Series: Laser in Forschung

und Technik, Berlin. 5. Kortz, H.P., Ifflander, R. and Weber, H. (1981): "Stability and beam divergence of multimode

laser with internal variable lenses" Appl. Opt. 20,4124. 6. Koechner, W. (1988): "Solid State Laser Engineering" Springer Series in Optical Sciences,

Berlin. 7. Emmett,!. L., Krupke, W.F. and Sooy, W.R. (1984): "The Potential of High-Average-Power

Solid State Lasers" Lawrence Livermore National Lab. Univ. California UCRL-53571. 8. Weber,H., Ifflander, R. and Seiler,P. (1986):"High power Nd-Lasers for industrial applications"

SPIE Proc 650, 92. 9. Chernoch, I.P.(1990): "Characteristics of a 1 kW-Nd-YAG hce-pumped Laser" I CLEO 1990,

Invited Paper. 10. Wittrock, U., Weber, H. and Eppich, B. (1991):"Inside-Pumped Nd-YAG tube laser" Optics Lett.

16,1092. 11. Comaskey, B., Beach, R., Mundinger, D., Benett, B., Freitas, B., Van Lue, D., Albrecht, G. and

Solarz, R. (1991): LLL-Labs "High Average Power Diode Pumped Slab Laser" IEEE, June 1991 12. Hodgson, N. and Eicher, J. (1989): "Efficiency of slab lasers systems", SPIE Proc. 1021, 147-52.

LASER PROCESS AUTOMATION

W.M.S1EEN Mechanical Engineering Department University o/Liverpool P.O. Box 147 Liverpool L69 3BX, U.K.

ABSTRACT. Optical energy can be delivered with minimal environmental disturbance. If any signal is obtained from a laser process then it most probably relates to the process itself mther than the energy delivery system. Thus it can be argued that one of the principal advantages the laser has as an industrial energy source is in its ability to fit into an automated system. There is currently much research into in­process sensing during laser processing as well as work on control strategies aimed at adaptive or intelligent control. Some of the sensors developed depend upon little known or expected in-process signals. Some of these are discussed together with the outline of control strategies for adaptive control.

1. Automation principles [1]

The drive towards automation is powered by the possibility of cost reductions, increased

productivity, increased accuracy, saving of labour, greater production reliability, longer

production hours, better working conditions for the human staff, increased flexibility of

production to meet the needs of changing markets and improved quality. This list is a formidable

argument for automation but it is only justified for certain production volumes. Fig. 1 gives an

idea of the stages which are most economical in setting up an automatic production facility. Into

this manufacturing pattern the laser has a place as another tool, but one with some significant

advantages:

Firstly, it is very flexible in the way it can be programmed to direct the optical energy.

Fig. 2 illustrates some of the options. One of the most flexible forms of laser beam guidance is

via a robotic beam delivery system. However the accuracy and neatness of the laser, particularly

31

S. Martellucci et al. (eds.), Laser Applications/or Mechanical Industry, 31-43. © 1993 Kluwer Academic Publishers.

32

ProcluctiOD rat.

Fig. 1. Variation of unit costs with production rate for different manufacturing strategies.

as a cutting tool, shows up the poor line following capability of current robots and this low level

of accuracy in the robot precludes its use for many applications. Fig. 3 gives an idea of the

growth of applications which would result from a successful development of an accurate robot.

There is surprisingly little effort being devoted to considering robot control algorithms for the

drive mechanisms. The main effort in robotic research seems to be spent on controlling the

sequencing of robots rather than their accuracy. Secondly, there is very little environmental

disturbance in delivering optical energy. For example, there is no electric field, no magnetic

field, (except in the electromagnetic radiation itselfI, there is no sound, no light or other optical

signal (except at the frequency of the laser beam), there is no heat, or mechanical stress. Thus

any signal in these areas will probably have come from the process itself. This gives a wide open

window for in-process diagnostics, which is unique to the laser.

Style

Cartesian gantry type

moving workpiece

laser rticuJated --, obot type 0

moving optics

moving Jaser hybrid

Fig. 2. ExampJes of automatic Jaser workstations.

sg(! .9= .... 0 tII,c ~o p,,'"' p,,'"' tII~

100r-------------~

"iU~ C)"" E-<,s -15

o ';;O:-:.5;-m-m--;::-O--;;.O~8~O=-.'="02=-m-m...J

Accuracy and smoothness

Fig. 3. Market for laser robotics vs accuracy.

33

In order to have a self regulating system for a laser or any machine it will have either an

open or closed loop controller. The open loop controller might be a clock if the sequencing is

done by time rather than events. In this case the machine actions are taken automatically but

without reference to the state of the process or the position of the machine. If it is controlled by

events then it would be a closed loop control system. A schematic of a closed loop controller is

shown in Fig. 4. It can be seen that the control sequence is: 1. A process variable or product

quality is measured; 2. The signal is compared to the desired value and an error detected; 3. This

error initiates a change in the process manipulators or drives, thus affecting the process; 4. What

is changed and by how much is decided by the controller. The start - if not the heart - of this

process is to be able to take the signal from the process while the process is running and to do so

sufficiently quickly that the error detection can be made and the machine corrected before there

is any considerable product waste. In-process signalling should be and is fast becoming one of

the strengths of the laser.

Instructions for

desired result

Analysis and decision

Actual result

Ag. 4. Diagram of the structure of a closed looped control circuit.

34

Table 1. Principle variables which characterise a laser process.

Beam Power Diameter Mode structure Location

Workstation Traverse speed Vibration, stability Focal position Shroud velocity and direction

Work iece Surface absorptivity Seam location Temperature "Quality" of product

Table 2. Some in-process sensors currently under investigation or on the market.

Signal Sensor Commercial Sensor Commercial or Reserch or Research

Beam power Laser beam analyser Commercial Leakage from cavity mirror Commercial Beam diameter LBA Commercial Perforated mirror Commercial and mode Hollow needle Commercial Location Acoustic mirror Research Modified LBA Research

LBA Commercial Scanning slot/beam splitter Research Edge thermocoul1les Research

Traverse speed and Encoders Commercial Linear Moire encoders Commercial table position Tachometers Commercial Laser interferometer Commercial

Laser Doppler Commercial anemometer{LDA}

Vibration/stability Accelerometers Commercial LDA Commercial Strain guages Commercial

Focal position Infra red Research Pressure Research Capacitance Commercial Commercial Inductance Commercial Research

Shroud gas Nozzle pressure Commercial Speckle interferometry Research Velocity Schlieren Research Surface abso!l1tion Acoustic mirror Research Back reflection Research Seam location Optical Res/Com Acoustic Research

Pressure Research Cutting quality TV camera on spark Research Acoustic mirror Research

discharge Teml!!:rature of cut face Research Viewing down beam Research

Welding quality Acoustic mirror Research Plasma charge sensor Research Acoustic workpiece Research Laser probe Research Sonic microphone Research Acoustic nozzle Research Optical emissions Research Video camera Research Electric signals Research

Surface hardening Temperature Res/Com Acoustic Research guality Infra red Research Cladding dilution Inductance Research Powder feed rate Pressure Research Vibration Research

Stress Research

35

2. In-process monitoring

For the control or monitoring of laser material processing, the in-process signals for the

variables listed in Table 1 are required. Many ideas are being and have been devised for these

tasks. Table 2 lists some of the main concepts being investigated or which have been engineered.

This is not a complete list, but nearly so. It does, I hope, illustrate the wide number of sensing

options open for laser process monitoring.

2.1. MONITORING BEAM CHARACTERISTICS

The laser power is nonnally measured when the beam is not being used, by some

technique which totally blocks the beam. For example, most lasers are fitted with a beam dump

which doubles as a calorimeter. Such devices are of no use for in-process sensing. It is amazing

that some production machines, fitted only with these methods of power sensing, have the laser

on for considerable periods of time and therefore have no way of being monitored - they are

simply running on faith that the manufacturer has made a stable product. Table 3 lists a number

of the techniques which have been patented or developed for the in-process monitoring of a laser

beam; that is monitoring while the beam is being used. Of these processes the laser beam

analyser and the acoustic mirror will now be described in more detail.

Table 3. Methods available for the monitoring of beam characteristics

Instrument Ref. Power Diam. Mode Wander Dirt Ref!. Response time

LBA 2 • • • • • • Fast Perforated Mirror 3 • • • Fast Chopper Devices 4,5 • • • Fast Heating Mirrors 6 • • Slow

7 • • • • Slow 8 • • • • Fast

Heating Wire 9 • • • Slow Photon Drag in Ge** 10 • • • Fast Piezoelectric** 11 • • • Fast Heating Gas 12 • Slow

13 • Slow 14 • Fast

Optical Scattering 15 • • Fast Acoustic Signals 16 • • • • • • Fast

*Muror or lens fouhng **In-process only if used With a beam sphtter.

36

laser LBA j -f]j--

beam dump shutter asacmbly

(b)

~ ....... beam diameter

ee)

Fig. 5. The Laser Beam Analyser (LBA). a) The principle whereby two orthogonal passes can be made simultaneously. b) An arrangement for continuous viewing of the beam. c) Example of the oscilloscope output.

2.1.1. The laser beam analyser (LBA) [2,3]. The laser beam analyser consists of a reflecting

molybdenum rod which is rotated fast through the beam. The reflections off the rod are measured

by two pyroelectric detectors placed as shown in Fig. 5a. The two detectors pick up signals

proportional to the power on two simultaneous orthogonal passes of the beam, as illustrated. It is

this ability of the instrument to collect the power distribution within around I / lOOth of a second

in two dimensions simultaneously with only 0.1 % beam interference that has made it one of the

more popular beam measuring instruments in a laser facility. It is often fitted after the output

window and before the shutter assembly so that the beam can be monitored even when it is not

being used. The arrangement is illustrated in Fig. 5b. The signals from the instrument can be

displayed on an oscilloscope or passed to a computer for further analysis. The data which can be

gained from the signals, which are illustrated in Fig. 5c, are: 1. Overall power; 2. Beam

diameter; 3. Beam wander; and 4. Mode structure. Any instantaneous variations in the above

properties can also be measured. It is quite surprising how much power vibration and beam

dilation there is in the beam from some lasers.

2.1.2. Acoustic mirror beam monitor [4]. This instrument may serve to illustrate the surprises in

store for laser engineers. It was found by Weerasinghe and Steen in 1984 that high frequency

acoustic signals were generated in mirrors which were reflecting laser radiation. The arrangement

is shown in Fig. 6. The signal responds to the power, beam diameter, position on the mirror,

state of tuning of the laser and even the gas composition in the laser cavity! What is considered

to be happening is that radiation falls on the mirror and in the action of reflecting the power,

some power is absorbed which instantaneously heats the surface atoms. This causes an expansion

and hence a stress wave which passes through the mirror (and water cooling at the back if

mirror

laser beam

piezo electric detector

recording device

Fig. 6. The acoustic mirror arrangement.

37

necessary) to the piezoelectric detector. Thus the instrument is recording only the variation in

power - not the absolute power. With this picture in mind the phenomena observed with this

instrument can be understood. Firstly an increase in laser power is usually associated with an

increase in the power variation. Some lasers are more stable than others. For example a slow

flow laser will give a smaller signal than a fast axial flow laser for the same power. A beam of

larger diameter will have a higher surface thermal stress than one of smaller diameter and hence

the signal would rise if the diameter increases with the same power fluctuations. If this device is

mounted as the last beam guidance mirror with a large view of the process then the large

fluctuations from the back reflected signal will dominate the output signal. It then becomes a

totally non intrusive in-process sensor. The only difficulty is that the signal can be caused by too

many effects and is therefore difficult to understand. However it remains a powerful beam

monitoring device and one which is ideal for alarm purposes, for example, in checking that the

beam is following the expected path. In Fig. 7 the signals from a stationary spot being melted by

I shutter opening I 24ms~

Power 1.7kW Spot size O.3mm

O~~ ______ ~ __ ~ __________ ~ __ ~

Time

Fig. 7. The acoustic signal obtained from two laser pulses on the same spot.

38

a laser are shown. The first pulse on a flat plate has a fairly unifonn signal with a possible 15 Hz

variation in it. The second pulse on the same spot has a high initial peak due to the strong

reflection from the resolidified smooth concave cavity left from the first pulse. It appears that

this took around 53 ms to remelt and then the signal again showed a faint 15 Hz variation.

Calculations by Postacioglu, Kapadia and Dowden [5] indicated that the expected natural wave

frequency on pools of this size would be around 15 Hz. Are we looking at (listening to) the

waves on the pool? Certainly the acoustic mirror would be a very significant in-process tool if

such intimate data could be obtained without any process interference.

2.2. MONITORING WORK TABLE CHARACI'ERISTICS

The work table variables are fairly straight forward as in the measurement and control of

position, travers~ speed or nOzzle gas·· velocity. Table 1 listed many of the well known

techniques. However the focal position is crucial and more subtle to measure.

2.2.1. Monitoring focal position. The focal position needs to be measured in real time because

the workpiece may warp slightly during processing or the part may be contoured and the

programming of the line to be followed may be simplified if the height above the workpiece is

controlled in-process rather than left to precise and time consuming programming. There are

several signals which are used currently for sensing and controlling the height of the nozzle

above the workpiece. They are shown in Fig. 8.

2.2.2. Seam following. In butt welding there is a need to follow the joint line. In laserwelding,

with the narrow fusion zone associated with laser welds, there is a need for a seam following

system which is accurate and fast. Several systems have been suggested. One of considerable

merit is the optical system of Lucas [6]. In this a CCD camera, using suitable filters, is able to

pick up the laser line which is fonned by passing a HeINe or diode beam through cylindrical

optics. The straightness of the line is analysed by a computer and the location of the seam found

within a few J.IS. The control system has been tested and proved to work in TIG welding [7]. An

alternative is to scan the beam, as with Oomen's method [8]. Inductive sensors have also been

developed which look at the variations in magnetic fleld around a joint [9].

2.3. MONITORING WORKPIECE CHARACTERISTICS

2.3.1. Temperature. The temperature of the workpiece is important in detennining the extent of a

a) Capacitor. inductance

~ K'SSSSSSSSS"

b) skids c) Feeler devices

d)OpUcal sensors

Fig. 8. Examples of various ways of sensing the focal position.

39

transformation hardening process or in the level of dilution expected during laser cladding. There

are several methods for examining temperature. There is a straighfforward pyrometer looking at

the interaction zone, as in Bergmann's method [10] for controlling the transformation hardening

process. There are CCD cameras which can look at the welding process and with appropriate

software compute the size of the weld pool from which the penetration might be determinable

[ 11 ]. Infrared red scanning pyrometers have been used to measure the thermal profile around the

event. Optical intensity has been viewed through the mirror by Zheng [12] using a fibre mounted

in the last beam guidance mirror. His signals indicated the extent of dross adhesion and the

formation of striations during laser cutting. Olsen [13] using a beam splitter was also able to

witness these detailed events during cutting. Miyamoto et al [14] had a system for viewing the

cut face during cutting. Their equipment was used to take some very informative film and could

be adapted to be less intrusive using fibres.

2.3.2. Keyhole monitoring. The monitoring of the keyhole by an acoustic mirror has been

discussed in Section 2.1.2. It can also be achieved using the "see through" mirror of Zheng et al

[12] . This has been developed considerably by H.B.Chen et al [15] who sensed the weld zone

with two ewavelengths; one in the infrared and one in the ultraviolet. It was found that the

ultraviolet signal responded primarily to the state of the plasma above the keyhole while the

infrared radiation was principally eminating from the weld pool. The device is capable of

identifying many weld faults in real time. What is even more intriguing is that for certain weld

faults the ultraviolet signal anticipates the infrared. Infact the collapse of the plasma predicts the

failure of the weld. For automatic control, particularly adaptive control, predictive signals are

extremely valuable. Another unexpected signal comes from an electric space charge associated

with the plasma, identified by Li et al [16] (the Plasma Charge Sensor) which can also diagnose

the general health of a laser weld while it is occurring. Since this is witnessing the state of the

plasma it too might be expected to be predictive. The plasma intensity has been observed by

many using optoelectric sensors, for example the work of Beyer [17] in Aachen. Others have

simply listened to the welding process by microphone.

40

A further alternative is the acoustic nozzle. This is different from the acoustic mirror in

that the piezoelectric detector is placed on or near the nozzle and detects mainly shockwaves

from the keyhole. The acoustic nozzle (AN) together with the plasma charge sensor (PCS) have

successfully been used to measure and log welding faults on a pilot laser can welding unit.

Together they are able to identify most welding defects in real time[18].

2.3.3. Spark discharge monitoring. The angle at which the sparks leave a cut and the cone angle

of the discharge is indicative of the health of the cutting process. ETCA [19] in Paris has

developed a TV camera system to look at the spark discharge from the underside of the cut. The

viewing angle is transverse to the cutting direction. The image is passed to an image processor

from which control data is elicited and used to control the process. The process being a multi

variable. multi option one means that some fonn of "intelligent" processing had to be developed

by ETCA. The general arrangement is shown in Fig. 9. Its weakness is that the instrumentation

has to be able to see beneath the workpiece. which is not always very convenient.

3. In-process control

It is one step to gain the diagnostic signals. the next is to use them in a closed loop control

system. These control systems come in two basic varieties. There are those which are one to one,

for example, the power is measured and the power is controlled, and there are those which have

many signals from which a choice of actions has to be calculated. In the one to one case it is

obvious what to do. For example, if the power is too low, raise the current to the discharge tubes.

If it is too high, do the opposite. The alternative control systems occur where there are many

diagnostic signals and many interrelated operating conditions to be adjusted. This latter type of

control system requires decision making software which could constitute "intelligent" processing

as discussed in Section 4. Examples of the one to one type of control is that of Li on laser power

[20], and the temperature control system of Bergmann [10] or Drefiker [21].

4. "Intelligent" in-process control

If one wishes to control a whole process such as welding. cutting or cladding, then the

Decision system

Main control system

workpIece

1Vcamera

Real time Image processing

Fig. 9. Example of the ETCA smart laser cutting system [34).

41

control system will have to handle many different in-process sensing signals and will be faced

with several control options.

For example in cladding there will be signals from the temperature. dilution, powder feed

rate, laser power, traverse speed, height of clad and possibly others. The control options are to

vary the power, powder feed, focal position and traverse speed. Thus the process will need one to

one control loops on power, powder feed, traverse speed etc in order to have control over these

variables and to know they will not wander; bat on top of this an overall diagnostic package is

required to determine what is wrong, if anything, and what to alter.

This decision is then passed to another package to determine by how much that parameter

should be altered. The " intelligent" part is for this decision making process to have its own

feedback so that it is capable of making faster and more accurate diagnosis and adjustments. The

main components for this have been seen in the arrangeament of the ETCA cutting system (Fig.

9).

The decision making can be performed by "fuzzy logic" [22] whereby a probability matrix

determining which parameter is the most likely cause of the fault condition has its values

adjusted by the programme itself in the light of the success of the previous decisions made. This

is a form of self teaching. An alternative technique is by using a neutral network to control the

process.

42

5. Conclusion

The laser is an ideal partner for automation due to the ease for gaining in-process signals

regarding the state of the process. The industrial advantage of fully automatic self correcting and

improving process are currently only partly perceived. However the technology is almost in place

for a factory working all day and night without heat or light and producing a uniform and high

quality product with only the help of a knowledgeable maintenance crew!

References

1. Steen, W.M. (1991): "Laser Material Processing" Ch. 7 publ Springer-Verlag, London. 2. Lim,G.C. and Steen,W.M. (1982): "The Mesurement of the Temporal and Spatial Power

Distribution of a High Powered CO? Laser Beam" Optics and Laser Technology, pp 149-153. 3. Lim,G.C.,and Steen,W.M. (1984): "Instrument for the Instantaneous in situ Analysis of the Mode

Stucture of a High Power Laser Beam" J. Phys.E. Sci.Instr. vol 17 pp 999-1007. 4. Steen,W.M. and Weerasinghe,V.M. (1986): "Monitoring of Laser Material Processing" Proc

SPIE conf. paper 650 22 Innsbruck publ by SPIE PO Box 10 Bellingham, Washington USA proc vol 650 pp 160-166.

5. Postacioglu,N., Kapadia,P and Dowden,J.(1988): "Capillary Waves on the Weld Pool in roduction Welding with a Laser" Journ. Phys.D. Applied Phys. Vol 22 pp 1050-1061.

6. Lucas,J. and Smith,J.S.(1988): "Seam Following for Automatic Welding" Proc SPIE Coni.Laser Technol: in Industry vol 952 ed. O.D.D.Soares Porto, Portugal, pp 559-564.

7. Sloan, K. and Lucas,J.(1982) "Microprocessor Control of TIG Welding Systems" IEE Proc. ploD l,ppl-8.

8. Oomen,G. and Verbeek,W.(1984): "A Real Time Optical Profile Sensor for Robot Arc Welding" Proc Intelligent Robots ROVISEC 3 Cambridge, Mass., USA.

9. Goldberg,F.(1985): "Inductive Seam Tracking Improves Mechanisation and Robotic Welding" Proc Automation and Robotisation of Welding, Strasbourg, France.

10. Rubruck,V., Geisler,E. and Bergmann,H.W.(1990) "Case Depth Control for Laser Treated Materials" Proc 3rd Europ. Conf. Laser Treatment of Materials ECLAr90 Erlangen, Germany, publ Sprechsaal, Coburg, Germany, pp207-216.

11. Juvin,D., de Prunelle,D. and Lerat,B. (1986): "SAO par Imagerie: Un Algorythme de Vision du Bain de Soudure pour Ie Controle et la Conduite du Procede Soudage TIG" Proc Aut des Procedes de Soudage, Grenoble, France.

12. Zheng, H.Y., Brookfield,D.J. and Steen,W.M. (1989): "The Use of Fibre Optics for In-process Monitoring of Laser Cutting" ICALEO'89 Orlando, Florida, USA, 140/ LIA vol 69 pp 140-154.

13. OIsen,F. (1989): "Investigations in Methods for Adaptive Control of LaserProcessing" Opto Electronik Magazin 4,2.

14. Miyamoto,I, Ohie,T. and Marno,H. (1988): "Fundamental Study of In-Process Monitoring in Laser Cutting" Proc CISFFEL 4 Cannes, France.

15. Chen,H.B., Li,L., Brookfield,DJ., Williams,K. and Steen,W.M. (1991): "Laser Process Monitoring with Dual Wavelength Optical Sensors" Proc ICALEO '91 San Jose, USA.

16. Li,L., Qi,N., Steen,W.M. and Brookfield,D.J. (1990): "On Line Laser Weld Monitoring for Quality Control" Proc ICALEO'90 Conf Boston, USA, published LIA Tulsa, Oklahoma USA.

43

17. Beyer,E. (1988): "Plasma Fluctuation in Laser Welding with CW CO? Laser" Proc ICALEO'87 San Diego USA May 1987 publ IPS (publ) and Springer-Verlag USA pp17-23.

18. Steen,W.M., Brookfield,D.J., Li,L., Chen,H.B. and. Keren,S. (1991): "In-Process Monitoring of Laser Welding for Can Manufacture" SERC ACME Research Conference, Leics.

19. Burg,B. (1986): "Smart Laser Cutter" Proc SPIE conf Innsbruck, Austria April 1986 vol 650 ed. Schoucker pub\. SPIE, Bellingham, Washington pp271-278.

20. Li,L., Hibberd,R.D. and Steen,W.M.(1987): "In-Process Laser Power Monitoring and Feedback C9ntrol" Proc 4th Int Conf on Lasers in Manufacturing LIM4 ed W.M.Steen Birmingham UK May 1987 publ IPS publ Ltd. Bedford UK ppI65-175.

21. Drenker ,A.,Beyer ,E.,Boggering,L.,Kramer .R.andWissenbach,K.( 1990): "Adaptive Temperature Control in Transformation Hardening" Proc 3rd Europ. Conf on Laser Treatment of materials ECLAT'90 Erlangen, Germany, publ Sprechsaal, Coburg, Germany, pp283-290.

22. Li,L., Steen,W.M., Hibberd,R.D. and Weerasinghe,V.M. (1988): "Real Time Expert System for Supervisory Control of Laser Cladding" Proc ICALEO'87 San Diego USA May 1987 publ IPS publ and Springer-Verlag in assoc LlA, Toledo, USA p 9-16.

THEORETICAL FOUNDATIONS

MODELING OF LASER MATERIALS PROCESSING

J. MAZUMDER, A. KAR Center for Laser-Aided Materials Processing Laboratory Department of Mechanical and Industrial Engineering University of Illinois at Urbana-Champaign 1206 West Green Street, Urbana, Illinois 61801, USA

ABS1RACT. This chapter summarizes the results of heat, momentum, and mass transfer for various types of processes to show the importance of transport phenomena in laser-aided materials processing. Various models for studying laser heat treatment, and fluid dynamics in laser melted weld pool are presented. Also, the mixing of externally added solute particles in laser melted pool is examined for laser surface alloying processes. A one-dimensional model is presented to study the rapid solidification phenomena during laser cladding and to obtain the nonequilibrium phase diagram. The heat and mass transfer equations are solved to model the laser chemical vapor deposition of films. An axisymmetric model for laser-induced materials damage is also discussed. Finally, this chapter describes some of the unique measurement techniques using visualization and laser spectroscopy recently developed for model verification.

1. Introduction

Several transport phenomena occur simultaneously in laser processing depending on the

incident intensity of the laser beam. For example, in laser surface hardening where power density

level is on the order of 103 to 104 W/cm2, heat transfer plays the most important role with mass

transfer detennining the limits of dwell time required for phase transfonnation. In surface

melting (or melt quenching) and welding where power density level is on the order of 105 to 107

W/cm2, momentum transfer or convection playa significant role. Convection also dominates the

process of laser surface alloying and cladding (power density is on the order of 105 to 106

W/cm2) along with mass transport which detennines the nonequilibrium microstructure and

composition of the solidified materials. Vaporization and plasma fonnation also play an

important role in detennining the surface contour, energy partitioning, and deep penetration

47

S. Martellucci et al. (eds.), Laser Applications for Mechanical Industry, 47-76. © 1993 Kluwer Academic Publishers.

48

welding. However, vaporization and gas dynamical effect become predominant in laser

processing when power density increases to 107 W/cm2 such as in drilling. Fig. 1 summarizes

the associated transport phenomena in various laser processing.

2. Transport phenomena theory for laser processing

2.1. HEAT CONDUCTION

Heat conduction plays an important role for all laser surface treatments, but for surface

hardening this is the single most important phenomenon. Several theoretical investigations have

been published with the objective of establishing the relationship between process parameters

and temperature and hardness distribution. Also, thennal analysis can be carried out to detennine

the range of various process parameters for laser chemical vapor deposition. Some of the

important models published to date are summarized here to serve as a ready reference for the

reader.

2.1.1. Analytical solutions. Different methods for solving the heat conduction equations for

various conditions are elegantly and methodically described by Carslaw and Jaeger [1]. Most of

the analytical solutions available are based on one of the many cases solved by these two

researchers and modified to suit the particular case.

2.1.2. One-dimensional transient model for flat semi-infinite body. Gregson [2] discussed a one­

dimensional model using a semi-infinite flat-plate soultion for unifonn heat source provided by

Carslaw and Jaeger [I]. He obtained a solution for an idealized heat source assumed to be

instantly applied and constant in time, which is valid if the thickness of the substrates is greater

than (4 <X t)1/2 and at the central axis of the depth, although they could be applied over the flat

hardened zone. These one-dimensional analyses may be applied to laser heat treatment processes

with unifonn heat sources which are produced by using optical systems such as a beam

integrator, dithered scanning beam, and high-power multi-mode beam with top-hat intensity

distribution. But before applying these equations, one should keep in mind that one-dimensional

solutions provide only approximate thennal distribution. For better thennal distribution, a two­

or three-dimensional analysis considering actual energy distribution and variable thennophysical

properties is required.

C = Cutting; W = Welding; A = Alloying; CL = Cladding; LM = Laser Machining; LAM = Laser Aided Machining

c:::J Convection

c::::::J Vaporization and Convection

1:::',,1 Convection, Non-equilibrium Partitioning in Mass Transfer

b,j;j'd Heat Transfer and Diffusion

Time-sec

Fig. 1. Operational regimes and associated transport phenomena for various processing techniques.

49

2.1.3. One-dimensional transient model for cylindrical bodies. Sandven [3] presented a model

which predicts the temperature distribution in the vicinity of a moving ring-shaped laser spot

around the periphery of the outer surface of a cylinder or the inner surface of a hollow cylinder.

Sandven [3] developed his model based on the flat-plate solution provided by Carslaw and

Jaeger [1] and presented graphical solutions for Z = 0 for various values of B are provided by

Sandven [3] to estimate an approximate hardened depth.

2.1.4. Three-dimensional model for semi-infinite plate. Cline and Anthony [4] solved the three­

dimensional heat conduction equation to analyze laser heating by considering the laser beam as a

Gaussian heat source.

2.1.5. Three-dimensional transient model for finite slabs. In the context oflaser chemical vapor

deposition, Kar and Mazumder [5] solved the three-dimensional transient heat conduction

equation with temperature-dependent thermophysical properties by considering both convective

and radiative losses of energy at the boundaries. The geometric configuration for this problem is

50

z

H~~------------------~

(0,0,0)

Fig. 2. Geometric configuration of the substrate and the relative position of the laser spot.

given in Fig. 2 where the laser beam moves in the x-direction at a constant velocity U. Fig. 3

shows that the chemically reactive zone, which is defined as the zone where the temperature is

more than or equal to that at which the chemical reaction that generates the film forming

material takes place, becomes narrower as the laser scanning speed increases. In this figure, the

chemically reactive zones are bounded by the curves A, B, C, and 0 for the scanning speeds 0.35,

1,0

A U =0.350 cm/s B U =0.250 cm/s

0.8 C U =0.167 cm/s 0 U = 0.125 cm/s Power = 600W

0 0.6 ~~

* A

:>. A

0.4 =:::::::~ 0

0.2

0'----'----....I.... __ --L. __ ----1 __ ---I

o 0.2 0,4 0.6 O.B 1.0 x*

Fig. 3. Variation of the width of the heated zone at the surface of the substrate along the direction of motion of the laser beam for various scanning speeds of the laser beam relative to the substrate for laser power 700 W.

51

0.25,0.167, and 0.125 em/s, respectively. There will be a critical scanning speed at which the

two boundary curves of the chemically reactive zone will collapse into one giving rise to the

narrowest possible film deposition region. With any other scanning speed higher than the critical

speed, the width of the chemically reactive zone cannot be reduced any further. So the film

deposition process has to be operated at a scanning speed lower than the critical speed. The

critical scanning speed is defined as the one at which the nondimensional peak temperature, Tp * is unity, where the peak temperature refers to the temperature at the center of the laser beam on

the top surface of the substrate. The line, Tp * = 1 is referred to as the line of the narrowest

chemical reaction zone in Fig. 4. This figure is plotted on logarithmic scales which show the

linear variation of the peak temperature (Tp *) with the laser scanning speed (U) for different

powers of the laser beam at various locations on the top surface of the substrate. The points of

intersections of the line of the narrowest chemical reaction zone with the curves of Fig. 4 gave

the critical scanning speed of the laser beam. The region which is to the right of the critical speed

is referred to as the chemically inert regime because the operating conditions of this region do

not raise the surface temperature of the substrate to the film forming chemical reaction

temperature. The region to the left of the critical speed is referred to as the chemically reactive

regime where the operating conditions are such that films offmite width can be deposited.

A x"=O,3 B x"=O,6 C x"=0'9 Power = 600W Line of the

Narrowest -Melting e c Zone ~ "Chemically Chemical

:::l "Reactive Reaction

~ ~~~~~::~;R~e~g~im:e~~~z~o~n~e ~ 100f- ~ocus of the '00.

~ Minimum LOC~ -'" o Q)

c..

Scanning Maximum Chemically Speed for Scanning Inert "-LCYD (line of Speed for Regime melting point) LCYD

uP Laser Scanning Speed (cm/s)

Fig. 4. Variation of the peak temperature with laser scanning speed relative to the substrate at various locations on the top surface of the substrate for laser power 700 W.

52

However, the surface temperature of the substrate can reach its melting temperature at a low

scanning speed for a given operating condition. Since melting the substrate is not desirable in

LCVD processes, the scanning speed has to be higher than the upper limit of the speed at which

melting occurs. The line, T p * = T mff d where T m is the melting temperature of the substrate, is

referred to as the line of melting point in Fig. 4. The points of intersections of the line of melting

point with the curves of Fig. 4 give the scanning speed above which the process has to be

operated to avoid melting the substrate. So, from the thermal considerations the operating regime

for an LCVD process is bounded by the line of melting point, the line of narrowest chemical

reaction zone, and their points of intersections with the curves of Fig. 4.

2.1.6. Numerical solutions. Numerical methods remove many of the limitations that apply to the

analytical methods. For example, the heat source does not have to be concentrated in a point, line

or plane. Temperature-dependent thermophysical properties and real boundary conditions may be

included. In spite of the inherent advantages in numerical methods, only a few numerical models

for heat flow in laser processing have been developed so far. The numerical solution to the three­

dimensional heat-transfer model for laser materials processing developed by Mazumder and

Steen [6] and later modified by Chande and Mazumder [7] allows for temperature-dependent

thermophysical properties, spatial distribution of the heat source (Gaussian, uniform, or any

other known distribution), radiative heat losses, convective heat losses, and latent heat of

transformation. A control volume approach is utilized where the heat balance on an axisymmetric

control volume as shown in Fig. 5. Courtney and Steen [8] used Mazumder's numerical model

[9] to predict the depths of the heat-treated zone and thermal cycle time as shown in Figs. 6 and

7. Knowledge of thermal cycle time is essential to calculate the carbon diffusion distance.

Mazumder's model [8] has also been successfully applied for welding [8] and melt quenching

experiments. However, the Beer-Lambert coefficient has to be experimentally determined to

account for convection which has been neglected.

2.2. CONVECTION

For processes such as melt quenching, alloying, and some welding, convection in the melt

pool dominates the process. It is also the single most important factor influencing the geometry

of the pool including pool shape, undercut, and ripples, and can result in defects such as variable

penetration, porosity, and lack of fusion. Convection is also primarily responsible for mixing,

and, therefore affects the composition of the weld pool. The heat transfer and thus the cooling

, /

E ----e-

P is the lattice point of the control volume drawn and represents the calculation centra for that alamanlal volume

Fig. 5. Nomenclature used in and around an elemental control volume.

53

rate is greatly enhanced in the presence of convection [10]. This in tum will affect the

microstructure. The homogeneity of solute redistribution during laser surface alloying (LSA) as

reported by Chande and Mazumder [7] can only be explained by the presence of convection

current

1800

~1400 .:it. 0> CI)

~1000 0-E CI)

I- 600:'

02 Q4 06 08 l.0 1.2 14 16 (a) Time (sec)

1800

1400 .:it.

0> CI)

~1000 0-E ~600

2000 02 0.4 0.6 0.8 10 12 14 16 Time (sec)

(b)

Fig. 6. Theoretically predicted thermal cycle during laser heating of EnS steel (power= 2 kW, beam radius = 3.00 mm, and reflectivity = 0.4); (a) speed = 22.5 mmls and (b) speed = 42.5 mmls (after Ref. [9]).

54

a. E <1l f-

'I--'\-~.-\--Eutectoid Temperature

800

600

400

0.30 0.75 0.93 1.10

200~L-WLllW~~-L~~LL~~~~~~ o 1 2 3 4 5

Depth. (mm)

Fig. 7. Theoretically predicted thermal profiles (power:= 2 kW, beam radius:= 3.0 mm, reflectivity := 0.4) (after Ref. [9]). Speed, mm/s: 08.0; 015.0; c" 22.5; <> 42.5.

Loser Beam

L

Scanning Direction ..

.. \

Fig. 8. Schematic diagram of the process.

55

Surface tension driven flow has been identified to be responsible for the convection within the

molten pool. While most of the work to date has been convection, a quantitative understanding

of the effect of convection on pool shape and mixing is of importance. Anthony and Cline did an

analysis on the surface tension driven flow within the molten pool. It is essentially a one­

dimensional problem and the flow field thus obtained is not coupled to the energy equation.

Therefore, no additional infonnation is obtained as far as the heat transfer process is concerned.

Dreper, et al. [12] developed a two-dimensional convection model. Buoyancy, electromagnetic,

and surface tension forces were considered. Numerical solutions were obtained based on a

specified pool proflle. It was found that the surface tension gradient is the dominant factor in

many cases. However, the solid-liquid interface is not known a priori. It is a part of the problem

to be solved. In fact, the solid liquid interface, i.e. the pool shape, is a piece of infonnation that is

of great interest. A two-dimensional transient model for laser surface melted pool was first

developed by Chan, Mazumder, and Chen [13,14]. It was shown that surface tension is

responsible for the fluid flow and convection. The cooling rate at the edge of the pool is found to

be higher than that at the bottom of the pool below the centerline which agrees with the

experimental findings that the microstructure is finer at the edge of the pool than at the bottom of

the pool. It is also predicted that the recirculating velocity is one or two orders of magnitude

higher than that the scanning velocity.

2.2.1. Two-dimensional self-consistent transient model for surface tension driver pool. A two­

dimensional transient self consistent (Le. solid-liquid interface position is a part of the solution)

model for laser melted pool was first developed by Chan, Mazumder, and Chen [14]. The

geometric configuration of the process is given in Fig. 8. The basic assumptions of this model

are as follows: 1) A laser beam having a defined power distribution strikes the surface of, an

opaque material of infinite width, thickness, and length; 2) Only part of the enrgy is absorbed; 3)

Absorbed energy induces surface tension driven flow due to the high temperature gradient; 4)

The liquid metal is considered to be Newtonian so that the Navier-Stokes equation is applicable;

5) All properties of the liquid metal and solid metal are constant, indipendent of temperature

(except the surface tension). This allows simplifications of the model; however, variable

properties can be treated with slight modifications. The dependence of surface tension on

temperature, the driving force of the flow, is assumed to be linear; 6) The latent heat of fusion is

neglected since the energy liberated is small compared to total enthalpy change associated with

temperature differences; 7) Thennal conductivity is assumed to be the same for bhot liquid and

solid phases for the sake of simplicity of the model; and, 8) The surface of the melt pool is

assumed to be flat to simplify the surface boundary condition and, hence, surface rippling is

56

neglected. A computer program SOLA [IS] is employed to study convection in laser-melted

pools. The basic method of the algorithm is presented in Ref. US]. This model provides the

details of the flow field and heat transfer on a plane perpendicular to the scanning direction. Flow

behavior predicted by this model will be valid for Regions I, 2, and 5 in Fig. 9. However, the

front-to-back motion and its effect on heat transfer also play important roles in the physical

process, especially for the fact that the recirculating velocity is so much higher than the scanning

velocity. Such a motion carmot be obtained from two-dimensional models.

2.2.2. Three-dimensional self-consistent quasi-steady state models. Recently developed three­

dimensional models [16-18] have provided better understanding of the flow behavior. It is

expected that better accuracy of the magnitude of flow will be obtained and results should be

valid for Regions I, 2, 4, and 5 of Fig. 9. Kou and Wang [18] developed a three-dimensional

model for convection in laser melted pool using both surface tension and buoyancy driven flow.

The semi-implicit method for pressure linked equation (SIMPLE) algorithm [19] was used to

calculate the velocity field. A conjugate heat transfer method was used at the solid-liquid

interface for calculation. This model predicted a velocity profile and magnitude (on the order of

m/s) similar to that of the transient two-dimensional model developed by Chan, et al. [14].

I Loser Beam

A I

3 \ 2 I li "" ................. - ..... -"'-------r- I I \ 5 I t I Fiol Surface

\ I 'I I

" I It" '--_J, ,,-4 ~_/

I

,. Stagnation Flow Region 2. Free Surlace Boundary layer Region 3. Cooled Corner Region 4. Solid·liquid Interlace Boundary Layer Region 5. Isothermallnviscid Core

Fig. 9. General features and various regions of the flow.

57

The conjugate heat transfer method uses discontinuous viscosity variation at the solidliquid

interface. Therefore, flow behavior at the SIL interface (Region 4, Fig. 8) will not be realistic

unless an extremely fine grid is used. However, this model also essentially reconfirmed that

surface tension induced by temperature gradient significantly affected the flow behavior and pool

shape. The group at the University of lllinois is taking a step-by-step approach to construct three­

dimensional models for laser melted pools. Initially a stationary spot source axisymmetric case

was considered [16]. Subsequently a perturbation solution based on low scanning velocity was

sought for the basic axisymmetric case [17]. The advantage of seeking a perturbation solution, as

it turns out, is that the three-dimensional flow is modeled by two sets of two-dimensional

equations which provide considerable computing advantages. However, the perturbation solution

is valid only for low scanning velocities. Another model which is a full three-dimensional

numerical solution of Navrer-Stokes equations, was developed using a point-by-point partially

vectorized iteration scheme. This model was then modified to accommodate a free surface [17].

These models were used to study the effect of convection on pool geometry, cooling rate, and

solute redistribution. The assumptions for materials properties for the quasi-steady state three­

dimensional models are similar to those of a two-dimensional model. These models also assume

a gray body radiative heat loss and flat surface as the starting point. It is important to note that

the solid-liquid interface is obtained as a part of the solution. This type of problem is generally

known as the Stefan problem in the literature [20]. An iterative scheme is used to solve for the

interface. The three-dimensional model is the numerical scheme that solves the above set of

governing equations. However, the perturbation model requires further mathematical

consideration. The basic solution is the axisymmetric case corresponding to a stationary laser

beam irradiating on a stationary workpiece. The perturbation is obtained by perturbing the basic

solution with a low scanning speed (low as compared to the recirculating flow due to surface

tension gradient). As it turns out, the angular dependence of the perturbation is separable. It is

more appropriate to use cylindrical coordinates since the basic solution is axisymmetric and the

perturbation solution can be reduced to an axisymmetric form. Details of the mathematical

derivation for perturbation model are available [17,21]. The Marangoni number, Ma, is the ratio

of the rate of convection associated with the recirculating velocity and the rate of conduction.

Thus, the higher the Marangoni number, the higher the convection. The Prandtl number, Pr, a

fluid property, is the ratio of kinematic viscosity and heat diffusivity. It is the ratio of momentum

diffusion and energy diffusion. The Reynolds number, Re, is the ratio of the inertia term and the

viscous term. For high Reynolds number, there would be a boundary layer phenomenon [22]. In

the case of numerical solution, this phenomenon is automatically taken care of by using a fine

enough grid. Convective heat transfer depends on the recirculating flow field. The higher the

58

recirculating flow field, the higher the convective heat transfer. However, a direct proportional

relationship cannot be assumed because of the complexity of the process due to the interaction of

temperature field and velocity field. The dimensionless melting temperature defines the solid­

liquid interface. For the same amount of heat input, the lower dimensionless melting temperature

would imply a larger molten region. The radiation factor governs the heat loss from the surface

due to radiation.

2.2.3. Axisymmetric model: numerical analysis and results. A standard alternating direction

iterative (ADI) [23] method is employed to solve the governing equation for the axisymmetric

case. The effect of convection and Prandtl number on melt pool shape is studied using this

model. The very existence of the flow field changes the mechanism of heat transfer to

convective. It tends to bring the higher temperature fluid on the surface right underneath the

beam sideways. This, in tum, melts down the solid at the edge of the molten pool, creating a

wider pool. Because of the conservation of energy, the size of the molten pool must remain

roughly unchanged, with or without convection. As a result, the molten pool becomes shallower.

The effect of an increase in the Marangoni number on the pool shape is studied. The Marangoni

number is a measure of the convection. An increase in the Marangoni number implies an increase

in the amount of heat being brought sideways. Consequently, the molten pool becomes wider

and wider. The ranges in aspect ratios (width to depth ratio) for different Marangoni numbers are

quite large, up to a 150 percent increase as compared to the conduction case. The effect of

changing the Prandtl number is also studied. The Prandtl number is the ratio of the momentum

diffusion to heat diffusion. To study its effect, the momentum diffusion must be kept constant

and only the heat diffusion is allowed to vary. This corresponds to keeping the Reynolds number

(Re ;;:; Ma/Pr) constant and changing Pr and Ma. This, in some sense, implies that the flow field

is being maintained while changing the thermal diffusion. Caution must be exercised because of

the complexity of the process due to the coupling of the energy and momentum equations. An

increase in the Prandtl number with the Reynolds number being kept constant will increase the

Marangoni number and hence the convective heat transfer. The aspect ratio will, therefore,

increase with the Prandtl number.

2.2.4. Numerical analysis o/perturbation model. Numerical solutions are sought for the two sets

of two-dimensional equations of the perturbation model. Finite difference equations, central

difference for the diffusion terms, and upwind difference for the convective terms governing the

velocities and temperature are derived. The staggered grid is used. The resultant nonlinear

algebraic equations are solved by the standard alternating direction iteration (ADI) method [23].

59

The computational procedure is as follows. For a certain temperature field where both solid and

liquid regions exist, the interface is first detennined. The velocities within the molten pool are

iterated for a prescribed number of iterations. Using this updated velocity field, the energy

equations are iterated next. Thus, the temperature field for the next global iteration is obtained.

The initial guesses are the steady state conduchon temperature and zero velocity for the basic

solution and zero temperature and velocity for the perturbation solution. The convergent

solutions are obtained when the average residual of each equation is smaller than a specified

small number, typically 10-4.

2.2.4.1. Numerical analysis of three-dimensional model. This model is a fully three-dimensional

finite difference scheme which iteratively evaluates the energy, momentum, and continuity

(modified pressure corrector) equations (Eqs. 32 through 34) and Eq. 36 for transient and steady­

state cases. For programming simplicity, a rectangular coordinate system was defined which uses

an in-line grid and primitive variable fonnulations. Because the material is assumed to be

scanning in the positive x, direction only, symmetry can be assumed about the slab midpoint (X2

= 0), thus reducing the calculation domain. Due to the large storage demands of a three­

dimensional program, the concept of extending the domain boundaries to sufficiently large

distances to simulate the infinite domain conditions is prohibitively expensive computationally.

To overcome this difficulty, a finite domain is chosen with given heat transfer coefficients at the

boundary. The heat transfer coefficients are chosen with the help of analytical as well as

numerical solutions of the conduction problem with moving boundaries for a large domain

exterior of the primary computing domain. The results represent an approximate simulation of

the infinite domain boundary condition. To penn it computations On a vectorized computer, a

mixed Jacobi/Gauss-Seidel iteration scheme was adopted to solve the finite difference equations.

The computation proceeds column-by-column according to the Gauss-Seidel scheme. Within

each column the Jacobi scheme is used to facilitate vectorization. The updating ofu,v,w is based

on the residual of the three momentum equations. Pressure correction is based on a point-wise

evaluation of the continuity equation, employing a modified artificial compressibility scheme.

The perturbation model has an advantage over the full three-dimensional model because the

fonner has two-dimensional governing equations. This allows finer mesh and shorter

computation time for the perturbation model. However, this model is limited to low scanning

speeds whereas the three-dimensional model is capable of handling high scanning speeds.

2.2.5. Perturbation model results. This model is used to study surface temperature, velocity

field, solute redistribution, temperature of the molten pool, and cooling rate:

60

a. Surface Temperature. The temperature attains its maximum at the center underneath the beam

and decreases radially outward. It is interesting to note that the temperature gradient, the driving

force, has three distinct regimes. Its magnitude increases from zero at the center in the radial

direction roughly to the edge of the laser beam, it then decreases, and finally increases again as

the flow a W roaches the edge of the molten pool. The fluid flow driven by the surface tension

gradient becomes important in the second stage where it tends to smooth out the temperature.

Finally, the flow approaches the edge of the molten pool and turns around. This phenomenon is

very similar to the heat transfer near the stagnation point and, therefore, gives rise to high

temperature gradient.

b. Velocity Field. The variation of surface temperature due to surface heating induces a surface

tension gradient thus pulling the molten materials radially outward. The general pattern of the

flow field is that the molten materials are going radially outward on the surface. As the flow

approaches the edge of the molten pool, It goes down and turns around. It then moves to the

center and comes up to complete the recirculation pattern. In addition to this recirculation, there

is also the motion of the workpiece. Therefore, material enters the molten pool from the front

portion of the pool and goes through the recirculation pattern, ultimately resolidifying on the

trailing edge.

c. ~ Redistribution. To help the reader gain a qualitative understanding of the mechanism of

solute redistribution, a particle trajectory is plotted in Fig. 10. This trajectory is presented in

three views--front, side, and top. The initial position is (xt, X2 *, X3 * = -2.97, 0.423, 0.01), a

point very close to the surface and before the molten pool. The recirculating pattern of the

particle can be clearly observed. This recirculating pattern implies that a particle will travel on a

rather long path before it freezes into the resolidifying surface. Consequently, the molten

materials can be well mixed. The composition within the melt pool is, therefore, uniform.

d. Temperature Field within ~ Molten fQQl The isotherms in the vicinity of the solid-liquid

interface are compressed near the surface of the pool and stretched apart near the bottom due to

the existence of the flow field. This distortion is due to recirculation of the molten material. The

Isotherms and the solid-liquid interface are asymmetrical due to the motion of the workpiece.

e. Cooling Rate. The cooling rate right af-er resolidification is related to the microstructure. The

higher the cooling rate, the flner the microstructure. The cooling rate (Le. the temperature

gradient) decreases from maximum at the centerline to zero at the edge of the molten pool and

also decreases from maximum on the surface to minimum at the bottom of the molten pool. Such

a variation in cooling causes a variation of the resolidified microstructure. Further investigation

of the solidification mechanism is required to fully predict the resulting microstructure.

61

2.2.5.1. Three-dimensional model results. This model is also used to study surface temperature,

velocity field, and coollng rate for higher scanning speed.

a. Surface Temperature. Isotherms are used to study the temperature distribution at the weld pool

surface. The isotherms are compressed near the solid-liquid interface due to the rapid

impingement of fluid -convecting heat) against the solid substrate. Circular isotherms near the

beam center suggests that the temperature distribution in the center of the pool is only mildly

affected by the scanning speed, thus validaeing the perturbation technique employed previously.

However, severe temperature gradients exist beneath the beam and at the solid-liquid interface.

b. Velocity field. Variation in surface tension drives the fluid radially outward along the surface

from the beam center until it reaches the solid-liquid interface. At this point the flow is turned

downward and recirculates along the pool bottom until it reaches the (beam) center where it is

forced upwards towards the surface. The large radial surface velocities are a direct consequence

of the dramatic surface temperature gradients. The thermocapillary velocities due to surface

tension effects are approximately two orders of magnitude larger than the substrate scanning

velocity.

c. Cooling ~ The results of three-dimensional model are essentially similar to those of

perturbation models although the present results show less fore-aft symmetry due to the higher

scanning speed.

~2 2 A , I S

-3 -2 -1

-3 -2 -1

~-

11 1nltl., pOIHton

f: final post tlon

Fig. 10. Trajectory of a particle as it enters, recirculates, and ultimately freezes. The trajectory is plotted in front, side, and top view in (a), (b), and (c). respectively. Initial position is (Xl' X2. X3) = (-2.97. 0.423.0.01) Ma = 1,500, Pr = 0.15, T m = 0.25 and RF = 0.0 (perturbation model).

62

2.2.6. Asymptotic analysis. Efforts at the University of illinois also include asymptotic analysis

of thetmocapillary flow under the laser beam (corresponds to Region I of Fig. 9) [24,25]. This

provides algebraic relationship for critical velocity, temperature, and shear stress. Accuracy

seems to be around 10 percent of the numerical soluti on [25].

2.2.7. Effect of flow on sUiface deformation. Most of the work discussed so far assumes flat

surface. The effect of surface rippling induced by surface-tension gradient during laser surface

melting and alloying was first studied by Anthony and Cline [11]. They essentially solved the

Navier-Stokes equation for one-dimensional steady-state condition with surface tension and

buoyancy forces, and derived an algebraic relationship for the ripple height. Chan et al. [17],

modified the three-dimensional numerical model to study the forced surface defotmation. The

method shows the following trends. Thennocapillary drives the surface fluid radially outward at

high velocities. These high velocities displace more mass from the central surface region than

can be replaced by the recirculating flow, thus causing a depression. The displaced mass builds

up at the solidliquid interface causing the surface to bulge upward where it is then forced

downward into the molten pool. For this calculation, the contact angle is unknown. The contact

edge height, however, was defined equal to the initial surface height.

2.2.8. Model for flow behavior with vaporization. Processes such as deep penetrating welding,

cutting, and drilling have vaporization affecting the flow dynamics. One of the major problems

of treating the vapor-liquid interface is the discontinuity at the interface and this makes

application of continuum mechanics rather difficult. Theoretical understanding of such problems

is at its infancy. Recently Chan and Mazumder [26] developed a one-dimensional steady-state

model where discontinuity of a low molecular mean free path at the solid-liquid interface is

incorporated using the "Knudsen Layer" jump condition and the Mott Smith type solution. In

this model, the vaporization process creates a recoil pressure which pushes the vapor away from

the target and expels the liquid. The materials are, therefore, removed in both the vapor and the

liquid phases. The materials removal rates are incotporated in the moving boundary

immobilization transfonnation. The vapor phase is assumed to be optically thin so that its

absorption of the high energy beam is negligible. However, this is a simplifying assumption at

this initial stage. Rockstroh and Mazumder [27] have observed from their spectroscopic work

that the beam does interact with the plasma. Due to the simplification, closed fonn analytical

solutions are obtained for this analysis. The effect of heat source power on removal rates,

vaporization rate, liquid expulsion rate, surface temperature, and Mach number are examined.

Results are obtained for three different materials--aluminum, superalloy, and titanium.

12.5

VI 2.8

"E

';;; > g 1.5 QJ

'" .. L

2 .. 1.9

'" ... o .. ... ~ .5

e 4------r-----r-----r----~----_, fl! 211 SI! 611

Power, kW

63

Fig. 11. Total material removal rate vs. beam power for three different materials: aluminum, superalloy, and titanium.

This analysis will be valid for drilling and onset of "keyhole" transformation for deep penetration

welding. A typical liquid layer thickness predicted by this analysis is shown in Fig. 11.

lt seems that for the power and interaction time (1 ms pulse width) the range of liquid

layer thickness is on the order of 10 to 30 f.1ID depending on the material. Fig. 12 shows the

2.5

2.9

'" E'1.5

'" 21 .. '" 1. B .. >

~ '"

• 5

10' 2D :all -II' S0 SII

Power in kW

Fig. 12. Vaporization, liquid expulsion, and total removal rates vs. beam power for aluminum.

64

materials removal rate for aluminum. The fraction of materials removed by vaporization and

liquid expulsion is compared. It can be observed from Fig. 12 that at lower power there is more

liquid expulsion than vaporization. At high power the converse is true. At very high power,

vaporization is expected to be the dominant form of material removal. Initial results of a two­

dimensional analysis show that for moderate power, liquid metal convection is important.

3. Mass transport

For process such as laser surface alloying and cladding, mass transport is a very important

phenomena. A mathematical model for mass transport during laser surface alloying would help

clarify several aspects of the problem. A set of processing conditions could be tested for

uniformity of mixing and for the resultant average compositions in liquid state. The mass flux

necessary to obtain a desired average composition in the liquid state could be computed. Powder

loss during alloying would then be estimable, knowing the actual mass flux during laser

processing. The model could be used to simulate the effect of varying the method of supplying

alloying elements. Having predicted an average liquid pool composition and measured the actual

solid state composition, the effective solute partitioning coefficient Cs * at the solid liquid

interface could be calculated. This effective Cs * could then be compared to the value determined

from the equilibrium phase diagram to check if conditions of local equilibrium existed ahead of

the solid liquid interface. This would be a useful check as local equilibrium is assumed in

predicting possible compositions during rapid solidification from equilibrium phase diagrams.

3.1. TWO-DIMENSIONAL 1RANSIENT MODEL FOR MASS TRANSPORT IN LASER

SURFACE ALLOYING

Chande and Mazumder [28] used the momentum transfer model described in the last

section and coupled that with diffusion equation to estimate the mass transfer during laser

surface alloying since convection dominates the process. For mass transfer calculations, the

following assumptions are made: 1) The effect of alloying on solute mass diffusivity was

neglected because accurate high temperature data were unavailable; mass transport by diffusion

was negligible compared to that by convection and it simplified the formulation; 2) Mass flux

was constant and uniform across the width of the pool at the surface. This rate and distribution

could be altered to allow for any other experimental condition such as wire feed or nonuniform

U

Cone from MX 40 -Pe' = 100· j = 5

--- Pe' = 10· j'= 5 --- Pel = lob· j = 10 --- Pe' = 10; j' = 10

a. j = Mass Flux - kg/m2 s _20 3

10

°o~~--~a~2--~~OA~-L~O~,6~-L~as~~--~lO III = lucid

65

Fig. 13. Summary of mass transfer calculations Showin~ growth of average composition with increasing interaction time for uniform mass flux of 5 and 10 kg/m -s and Pe' values of 100 and 1000.

powder addition; and, 3) There was no transfer of alloying elements across the solid liquid

interface This was a very good assumption that was verified experimentally using electron probe

micro aralysis (EPMA) technique. Melting of powder particles introduced into the melt was

found to be practically instantaneous [28]. Thus mass flux could be considered as being added in

the liquid state. The result of the mass transfer calculations indicate the importance of fluid flow

in determining solute distribution during laser surface alloying. The small effect of changing

solute diffusivity on the solute distribution, the uniform mixing even when mass flux was present

over only a part of the pool surface and the nature of the threedimensional plots all show that the

pattern of fluid flow controls the resultant solute distribution. The effect of laser-substrate

interaction time on solute distribution during surface alloying is shown in Fig. 13.

3.2. ONE-DIMENSIONAL TRANSIENT MODEL FOR LASER CLADDING

Laser cladding is a process used for altering the surface properties of substrate materials.

Due to high cooling rate in laser port are cladding, extended solid solution are formed whenever

a mixture of several elements is used as cladding powder. Usually, the extended solid solutions

have higher melting point, better oxidation and corrosion resistant properties than the alloys with

eqUilibrium compositions. To study the extension of solid solubility during laser cladding, a

diffusion model is developed by Kar and Mazumder [29,30,31] under the following assumptions:

a) The thermal conductivity and the thermal diffusivity for a mixture is the sum of the volume­

averaged value of the respective transport properties of each element of the mixture; b) The mass

diffusivity of each element in the liquid phase is the average value of self-diffusivity over the

room temperature and the initial temperature with modified activation energy for the mixture;

66

c) The cladding pool and the substrate are in thennally perfect contact; d) There is no mass

diffusion in the solid phase; e) The solute segregated at the solid-liquid interface moves into the

liquid phase by diffusion only (this is because a boundary layer is fonned near the interface

where diffusion of solute atoms is dominant); t) The cladding melt fonns a unifonn solution of

composition equal to that of the cladding powder mixture before its solidification begins; and,

g) Only 50% of the laser energy is absorbed by the cladding material studies [32] show that the

amount oflaser energy absorbed by different materials is 37-60%.

The procedure to solve this problem can be found in Ref. [30] where the model has been

discussed in detail and applied to the Ni-Hf system. The importance of studying the Ni-Al

system [31] compared to that of Ref. [30] can be understood by analyzing the eqUilibrium phase

diagrams of the Ni-Al and the Ni-Hf systems.

The eqUilibrium phase diagrams of various systems can be divided into several regions

based on the sign of the slope of the phase diagrams. Each of these regions can be divided further

into smaller sections based on the assumption of constant equilibrium partition coefficient (Ice)

which is true over a small range of the concentration around the point of interest on the

equilibrium phase diagram. In Ref. [30], we have presented the effects of that regime of the

equilibrium phase diagram that has negative slope. The study of Ref. [31] examines the effects of

that region of the equilibrium phase diagram that has positive slope.

The model predictions have been compared with experimental data. Laser cladding was

perfonned on nickel substrates with a mixture of -Ni-Hf powder of nominal composition by

weight, 74% Ni and 26% Hf in one case and 74% Ni and 26% AI in the other case. STEM

analysis of these samples shows that the concentration of Hf in the I phase and Al in the

martensitic solid solution regions is in excess of that predicted by the equilibrium phase diagram.

The percent model predicts the composition of the extended solid solution quite well. These

results have been presented in Table 1. A small fraction of the NiHf alloy was found

experimentally ton contain 9.3 percent (by wt.) Hf in the Ni matrix.

Table 1. Comparison of theoretical results with experimental results of extended solid solution.

Nominal Composition of the Cladding, wt. % Laser Power, (kW) Laser Beam Diameter, mm Workpiece Speed, in/min Initial Pool Mean Temperature (T2)' K Composition in the Solid Solution, wt. % Theoretical Results Experimental Results

74% Ni, 26% Hf 5 3 50 1862

7.15 Hf 6.5Hf

74% Ni, 26% AI 5 3 50 2420

30.76 AI 29 AI

67

1510

1\ ~ \ 'yNon- Equilibrium \ ::J

1500 tT I :.:j I I I

Q \ I

- 1490 \

2! I .~ I

.2 ~ \ 0 \

~ 1480 I I I

'" I l- I

\ , 1470

, r--Equi'ibrium

M60~--~--~--~~--~--~--~ o ·5 10 15 20 25 30

HF (WI%)

Fig. 14. Comparison of the nonequilibrium phase diagram of the extended solid solution of nickel­hafnium with its eqUilibrium phase diagram.

Also, the above model was used to obtain the nonequilibrium phase diagrams for Ni-Hf

and Ni-Al alloy systems as shown in Figs. 14 and 15. The characteristic parameters for these

figures have been presented in Table 2.

1866 I

i-1-Q 1865

'" ~ "0 :0 a.

~ 1864 ~

1863 26

I I

YQuid I I I

~ I I

E n' i~ I I QUII num I

N Tb' ('1 oneQUI I num I

27 28 29 AL (wt%l

I I I

SOlid~ I

30 31

Fig. 15. Comparison of the nonequilibrium phase diagram of the extended solid solution of nickel­aluminum with its equilibrium phase diagram.

68

Table 2. Solidus composition of Hf, soJidus/liquidus temperature at the substrate-cladding interface and the speed of solidification at this interface.

Solute Element Composition of Hf Solidus/Liquidus Temperature Interface Speed (wt%) (K) (cm/sec)

3.19 1505 3.26 4.83 1499 2.88

Hafnium 6.23 1489 2.74 7.15 1480 2.68 9.58 1453 2.55

30.83 1864 6.1 Aluminum 30.74 1863 5.6

30.56 1862 4.5

Deviation of nonequilibrium phase diagram from the equilibrium one shows the extension

of solid solubility that can be obtained during laser cladding. The nonequilibrium phase diagram

has been plotted in the neighborhood of the melting point of the cladding powder.

It can be seen from Fig. 14 that the width of the solid-liquid region between the

equilibrium solidus and liquidus has reduced considerably. The nonequilibrium solidus line of

this figure shows the extension of Hf concentration in Ni that can be obtained due to the rapid

cooling in laser cladding. This phenomenon can be understood from the fact that the equilibrium

phase diagram of Ni-Hf has a negative slope at the point which corresponds to the nominal

composition of the cladding powder.

Due to this, Hf is rejected fro the Ni matrix when the solution of Ni-Hf solidifies and this

results in increasing the concentration of Hf in the liquid phase. But since the solid phase retains

more Hf than its equilibrium composition, the liquid phase will have less Hf than the equilibrium

value and hence the extended solid solution (that is, the nonequilibrium) phase diagram will

shrink:.

On the contrary, it can be seen from the Fig. 15 that the width of the solid-liquid region

between the equilibrium solidus and liquidus has increased considerably. The nonequilibrium

solidus line of this figure shows the extension of Al concentration in Ni that can be obtained due

to the rapid cooling in laser cladding.

This phenomenon can also be understood from the fact that equilibrium phase diagram of

Ni-AI has a positive slope at the point which corresponds to the nominal composition of the

cladding powder. Due to this, AI is retained in the Ni matrix whereas Ni is rejected into the

liquid phase as solidification proceeds. This causes an extension of AI concentration in the solid

phase and lowers its weight fraction in the liquid phase and thus enlarges the solid liquid region

between the solidus and the liquidus lines.

3.3. THREE-DIMENSIONAL TRANSIENT MODEL FOR LASER CHEMICAL VAPOR

DEPOSITION (LCVD)

69

Laser chemical vapor deposition is a process of depositing thin films by pyrolytic or

photolytic decomposition of suitable reactants. In Ref. [35], a mathematical model for mass

transfer inside a laser chemical vapor deposition (LCVD) chamber has been presented for pure

titanium deposition from titanium tetrabromide on a stainless-steel (SS 304) substrate. A set of

five three-dimensional, transient, and nonlinear partial differential equations has been solved to

account for the diffusion of various species inside the LCVD chamber and the thermal

decomposition of titanium tetrabromide, titanium dibromide, and titanium monobromide at the

surface of the substrate in order to determine the thickness of the titanium film deposited on the

substrate. The model has been used for studying the effects of various parameters on the

deposition and shape of titanium films. The parameters which were varied are as follows: laser

power, activation energy, total pressure inside the LCVD chamber, and partial pressure of

titanium tetrabromide. Under certain conditions, the deposited titanium film has been found to

have a volcanic proftIe. This model has been applied in Ref. [36] to study the LCVD of TiN

from TiC4, and the model predictions have been found to be in good agreement with

experimental results as shown in Figs. 16 and 17.

3.4. LASER-INDUCED MELTING AND VAPORIZATION

Materials damage due to melting and vaporization during laser irradiation has been studied

in Refs. [37, 38]. It has been found that deep holes can be drilled under certain operating

conditions. In Ref. [38], a mathematical model has been developed to study the effects of

multiple reflections and the shear stress-induced liquid metal flow caused by the assist gas during

materials removal and drilling with a pulsed Nd:YAG laser, which is considered Gaussian

spatially and triangular temporally with duty ratio 0.21. To account for the effects of polarization

on absorptivity and the propagation of laser beam through the liquid metal film, an effective

absorptivity has been used in the mode1. The effects of various process parameters such as the

laser power, beam radius, and pulse-on time on the cavity depth, recast layer thickness, and

cavity tapering have also been examined. It has been found that the laser-induced materials

removal process can be described in terms of a single parameter called the "gross" laser intensity.

The cavity depth varies linearly with the gross intensity over a certain range of the process

parameters in the absence of multiple reflections, however, this variation becomes nonlinear in

the presence of multiple reflections as shown in Fig. 18.

70

30.0

2S.0

20.0

! n IS.O

J W.O

S.O

0.0

·1

p= 300 W •••• t -= 8 I

·D.S O.S Distance from beam center (mm)

...... ""

Fig. 16. Spatial variation of the thickness of TiN deposits. experimental; calculated: r = 1.0 mm, Tm = 1473 K, TM = 1640 K, E = 12.1 Kcal/mole, leo = 9.0 x 102 cm/s.

Table 3. Numerical values of the dimensionless parameters of the 9 cases of the 3-D model.

Beam radius 0.5 mm Case No.

1 2 3 4 5 6

8.0

! 6.0

j 4.0

2.0

Scanning speed (mm/s)

"" "" "

25.0 25.0 25.0 50.0 50.0 50.0

Lserpower (kW)

P=400 W t a 2s

" .. "

2 3 4 3 4 8

Distance from the beam cenler (mm)

Aspect ratio

3.375 3.51 3.62 3.66 3.79 4.0

Fig. 17. Spatial variation of the thickness of TiN deposits. experimental; calculated: r = 1.0 mm, Tm = 1473 K, TM = 1640 K, E = 11.0 Kcal/mole, leo = 8.4 x 102 cm/s.

30~--------------------------------~

E 20 .§. .J::

a. (I)

o .?:-

~ 10

--0- With Reflection, Without Liquid Metal Flow

- With Reflection, With Liquid Metal Flow

--0- Without Reflection,

Without Liquid Metal Flow

- Without Reflection, With Liquid Metal Flow

10 20 30

.12 ·2 Gross laser Intensity, Ig x 10 (W m )

Fig. 18. Variation of the cavity depth with the gross laser intensity.

4. Experimental results

71

a. ~ !lUi!.l ill ~ resolidified region: The aspect ratios based on the micrographs are

calculated and tabulated in Table 3. It can be seen that the aspect ratio increases with respect to

Fig. 19. Steel melt pool created with a 1.2 kW CW C02 laser. Image acquired during the process by illumination with both diffused and focused Argon-ion laser light.

72

(3) View from above.

Depth ~ 0.032 nun.

Width = 1.12 mm.

Length ~ 1.34 mm.

(b) Side vi ew.

Depth ~ 0.032 mm.

Length = 1.34 mm.

(e) Front view.

;)epth = 0.032 mm.

Widlh = 1.12 mm.

Fig. 20. Views of the molten weld pool surface contour when the workpiece surface is 12.99 cm away from the focusing lens.

Fig. 21. X-ray shadow graph of two successive frames showing the motion of the tungsten particle.

73

laser power which is consistent with the theoretical prediction. The experimental aspect ratio

varies from 3.375 to 4.0. The theoretical models predicts 4.0 to 5.0 for the aspect ratio.

b. Melt UQQl visualization anQ ~ surface deformation: Recently, a visualization

technique has been developed [39], which is based on the illumination of a laser melt pool

simultaneously with focused and diffused argon-ion laser light, to obtain a greatly improved

image of the molten surface as shown in Fig. 19. This method will enable one to obtain the pool

shape data on-line during the process. The method is described in detail in Ref. 40. A modified

visualization technique is further developed to provide the melt pool free surface deformation as

shown in Fig. 20.

c. X-ray shadow graph: In Fig. 21, two successive frames of the x-ray shadow graph are

shown. The locations of the center line and tungsten particle are shown by the vertical lines. The

left line directly underneath the electrode is the center line. The right line is the location of the

tungsten particle. Assuming that the particle travel with the motion of the molten materials, the

speed of the recirculating flow can be estimated. It is found to be of the order of 30 cm/sec which

is of the same order of magnitude of the theoretical prediction.

5. Concluding remarks

Comments on heat conduction, convection, and mass transfer processes as applied to laser

processing of materials can be made in the following way:

(a) Heat conduction: Although the transfer of laser energy, that is heat conduction within the

substrate materials can be analyzed by elegant mathematical techniques, more work still remains

to be done to understand the coupling of laser energy with the substrate materials.

(b) Convection: The present three-dimensional model brings additional insight to the convection

and fluid flow within the melt pool. The model gives details of three-dimensional velocity and

temperature fields. Consequently, a more realistic prediction of pool shape is obtained. The

detailed threedimensional velocity field gives a quantitative explanation of the mechanism of the

mixing process within the molten pool. A more realistic prediction of the cooling rate is now

available for laser melted pool. The perturbation model yields better spatial resolution because

two dimensionality renders it practical to employ a finer mesh. On the other hand, the three­

dimensional model is capable of handling larger scanning speeds than the perturbation model. It

is found that in the presence of thermocapillary convection, the physics of the process changes

from conduction-dominated to convection-dominated.

74

This changes the pool geometry dramatically, resulting in up to a 150 percent increase in the

aspect ratio (width/depth) as compared to the pure conduction case. The temperature gradient on

the resolidifying surface, which is directly proportional to the cooling rate, is found to be

nonuniform. This, in tum, gives rise to the non-uniform resolidified microstructure. The

mechanisms of solute redistribution are discussed qualitatively. It is found that a particle

recirculates many, many times in the molten pool before it resolidifies. Consequently, the solute

can be well mixed. Although understanding of flow dynamics of laser melted pure metal pool

has improved significantly, a vast amount of work is still needed to understand the effect of

convection on free surface and alloying element addition. Also, combined vaporization and

liquid convection needs further attention to understand keyhole formation, cutting, and drilling.

(c) Mass transfer: Although laser surface treatments are complicated processes, proper

understanding of mechanisms can be obtained by methodical modeling. The problem can be first

approached in simplified form with proper assumptions. Each step will be useful for a range in

spite of their limitation. With gradual removal of simplifying assumption, an even complicated

process such as this can be studied.

ACKNOWLEOOEMENTS. This chapter includes many years of a group (consisting of

Professors J. Mazumder, M. M. Chen, Dr. T. Chande, Dr. C. L. Chan, Dr. A. Kar, Dr R. Zehr,

and Dr. D. D. Voelkel) effort at the University of illinois sponsored by U.S. Office of Naval

Research (Grant NOOOI4-84-K0315), U.S. Air Force Office of the Scientific Research (Grant

AFOSR: 85-0333), and the U.S. National Science Foundation (Grant NSF MSM 84-12118). The

authors thank Professor O. Conde of the Physics Department of the University of Lisbon for

making the comparison between the theoretical and experimental results on TiN dots possible.

References

1. Carslaw, H. S., and Jaeger, 1. C. (1959): "Conduction of Heat in Solids", 2nd. ed.,Oxford Univ. Press, pp. 11-12.

2. Gregson, V. (1983): "Laser Heat Treatment", Laser Materials Processing, ed. M. Bass, North Holland, pp. 201-234.

3. Sandven, O. A.(1979): "Heat Flow in Cylindrical Bodies during Laser Transformation Hardening", Proc. SPIE, vol. 19B, pp. 138-143.

4. Cline, H. E., and Anthony, T. R.(1977): "Heat Treating and Melting Material with a Scanning Laser or Electron Beam", J. Appl. Phys., vol. 48, pp. 3895-3900.

5. Kar, A., and Mazumder, J.(1989): "Three-Dimensional Transient Thermal Analysis for Laser Chemical Vapor Deposition on Uniformly Moving Finite Slabs", J. Appl. Phys., vol. 65, pp. 2923-2934.

6. Mazumder, J., and Steen, W.M. (1980): "Heat Transfer Model for Continuous Wawe Laser Materials Processing", J. Appl. Phys., vol. 51, pp. 941-947.

75

7. Chande, T., and Mazumder, 1.(1981): "Lasers in Metallurgy", eds. K. Mukherjee and I.Mazumder, The Metallurgical Society of AIME, Warrendale, PA, pp. 165-178.

8. Courtney, C. G. H., and Steen, W. M.(l979): "Surface Heat Treatment of End Steel Using a 2 kW Continuous-Wave C02 Laser", Met. Technol., vol. 6, pp. 456-462.

9 Mazumder, 1.(1978): "Laser Welding of Titanium 6Al-4V-Ti", Ph.D. Thesis, London University.

10. Chande, T., and Mazumder, 1.(1982): "Mass Transport in Laser Surface Alloying: Iron Nickel System", Appl. Phys. Lett., vol. 41, pp. 42-43.

11. Anthony T. R., and Cline, H.F. (1977): "Surface Rippling Induced by Surface Tension Gradients during Laser Surface Melting and Alloying", 1. Appl. Phys., vol. 48, pp. 3888-3894.

12. Dreper, G. M. Eager, T. W., and Szekely, 1.(1983): "Convection in Arc Weld Pools", Welding 10urnal, vol. 62, pp 307-312.

13. Chan, C. L., Mazumder, 1., and Chen, M. M.(1983): "Applications of Lasers in Materials Processing", ed. E. A. Metzbower, ASM, p. 150.

14. Chan, C. L., Mazumder, 1., and Chen, M. M.(1984): "A Two-Dimensional Transient Model for Convection in Laser Melted Pools", Metall. Trans., vol. 15A, pp. 2175-2184.

15. Hirt, C. W., Nichols, B. D., and Romero, N. C.(1975): "A Numerical Algorithm for Transient Fluid Flows", UC-34 and UC-79d.

16. Chan, C. L., Mazumder, 1., and Chen, M. M.(1987): "A Three-Dimensional Axisymmetric Model for Convection in Laser Melted Pool", Mat. Sci. Engineering, vol. 3, pp. 306-31l.

17. Chan, C. L., Zehr, R., Mazumder. 1.. and Chen. M. M.(1986): "Three-Dimensional Model for Laser Weld Pool". Proc .. 3rd Engineering Foundation Conference on Modeling and Control of Casting and Welding Processes. eds. B. Kuo and R. Mehrabian, TMS-AIME, Warrendale. PAt pp. 229-246.

18. Kou. S .• and Wang. 1. H.(1986): "Three-Dimensional Convection in Laser Melted Pools". Metall. Trans .• vol. 17 A. pp. 2265-2270.

19. Patankar. S. V .• and Spalding. D. B.(1972): "A Calculation Procedure for Heat. Mass, and Momentum Transfer in Three-Dimensional Parabolic Flows", Int. 1. Heat and Mass Transfer,voI.15,pp.1787-1806.

20. Viskanta. R.(1984): "Solar Heat Storage: Latent Heat Materials", ed. G. A. Lane, CRC Press, New York, pp. 153-222.

21. Chan, C. L.(1986): "Thermocapillary Convection during Laser Surface Heating", Ph.D. Thesis, University of Illinois at Urbana-Champaign.

22. Schlichting, H.(1960), "Boundary Layer Theory", McGraw Hill, New York. 23. Carnahan, B., Luther, H. A., and Wilkes, 1. 0.(1969): "Applied Numerical Methods", 10hn

Wiley & Sons, New York. 24. Chan, C. L., Chen, M. M., and Mazumder, 1.(1988): "Asymptotic Solution for Thermocapillary

Flow at High and Low Prandtl Numbers due to Concentrated Surface Heating", ASME 1. of Heat Transfer, vol. 110, pp. 140-146.

25. Chan, C. L., Chen, M. M., and Mazumder, 1.(1985): ASME/AIChE National Heat TransferConference, ASME, New York, 85-HT-21, p. 1985.

26. Chan, C. L., and Mazumder, 1.(1987): "One-Dimensional Steady-State Model for Damage by Vaporization and Liquid Expulsion during Laser Materials Interaction", 1. Appl. Phys., vol. 62, pp.4579-4586,1987.

27. Rockstroh, T. 1., and Mazumder, 1.(1987): "Spectroscopic Studies of Plasma during Laser Materials Interaction", 1. Appl. Phys., vol. 61, pp. 917-923.

28. Chande, T., and Mazumder, 1.(1985): "Two-Dimensional, Transient Model for Mass Transport in Laser Surface Alloying", 1. Appl. Phys., vol. 57, pp. 2226-2232.

29. Kar, A., and Mazumder, 1.(1987): "One-Dimensional Diffusion Model for Extended Solid Solution in Laser Cladding", 1. Appl. Phys., vol. 61, pp. 2645-2655.

30. Kar, A., and Mazumder, 1.(1988): "One-Dimensional Finite-Medium Diffusion Model for Extended Solid Solution in Laser Cladding of Hf on Nickel", Acta Metall., vol. 36, pp. 701-712.

76

31. Kar, A., and Mazumder, J.(1988): "Extended Solid Solution and Nonequilibrium Phase Diagram for Ni-A1 Alloy Formed during Laser Cladding", Metall. Trans. A, vol. 20A, pp. 363-371,

32. Li, L. J., and Mazumder, 1.(1984): "A Study of the Mechanism of Laser Cladding processes", in Laser Processing of Materials, eds. K. Mukherjee and 1. Mazumder, The Metallurgical Society of AlME, Los Angeles, CA, pp. 35-50.

33. Boettinger, W. J., and Perepezko, J. H.(1985): "Fundamentals of Rapid Solidification", Rapidlv Solidified Crystalline Alloys, eds. S. K. Das, B. H. Kear, and C. M. Adam, The Metallurgical Society of AlME, Warrendale, PA, pp. 21-58.

34. Aziz, M. 1.(1982): "Model for Solute Redis-ribution during Rapid Solidification", 1. Appl. Phys., vol. 53, pp. 1158-1168.

35. Kar, A., Azer, M. N., and Mazumder, 1.(1991): "Three-Dimensional Transient Mass TransferModel for Laser Chemical Vapor Deposition of Titanium on Stationary Finite Slabs," 1. Appl. Phys., vol. 69, pp. 757-766.

36. Conde, 0., Kar, A., and Mazumder, J.(1992): "Laser Chemical Vapor Deposition of TiN Dots: A Comparison of Theoretical and Experimental Results," 1. Appl. Phys. (accepted for publication).

37. Kar, A., and Mazumder, 1.(1990): "Two-Dimensional Model for Material Damage due to Melting and Vaporization during Laser Irradiation," 1. Appl. Phys., vol. 68, pp. 3884-3891.

38. Kar, A., Rockstroh, T., and Mazumder, 1.(1992): "Two-Dimensional Model for LaserInduced Materials Damage: Effects of Assist Gas and Multiple Reflections inside the Cavity," 1. Appl. Phys., vol. 71, 15 March.

39. Voelkel, D. D., and Mazumder, J.(1990): "Visualization of a Laser Melt Pool," Applied Optics. vol. 29, pp. 1718-1720, 1990.

40. Voelkel, D. D.(1992): "Laser Weld Pool Optical Diagnostics: Topography and Visualization used for Surface Contour Analysis and Process Monitoring," Ph.D. Thesis, University of Illinois at Urbana-Champaign.

ANAL YTICAL MODEL FOR ABERRATED DIFFRACTION IN HIGH POWER CW LASER BEAM TRAINS: LASER CAVITY TO WORK PIECE

J.R.PALMER University of Alabama Center for Applied Optics Huntsville, Alabama 35899, USA

W.M.STEEN University of Liverpool Mechanical Engineering Department P.O. Box 147 Liverpool L69 3BX, U.K.

S. MARTELLUCCI Dipartimento di Ingegneria Meccanica 1l Universita' di Roma "Tor Vergata" Via O. Raimondo, 00173 Roma, Italy

ABSTRACT. This chapter describes an analytical model for evaluating the beam quality necessary for welding, cutting, hard surfacing, and similar commercial industrial processes. The high power lasers used in commercial industry are attaining even higher power levels and considerations of beam distributions for multiple stations is becoming necessary for large manufacturing plants. Most of these high power laser systems will require a beam train containing two to ten optical elements to direct the beam to the work piece. As the power increases, so does the optical distortion at each optical element. The model developed in this chapter will allow the designer to evaluate the amount of distortion budget that can be allowed for each optical element in the train, and determine the amount of diffraction that will ultimately be present at the focusing optics. This will allow an analysis of the beam waist and focusing capability of the focusing optic on the work piece. Depending on the mode characteristics of the laser used, the designer will be able to perform an end-to-end analysis of the beam quality at the work piece.

1. Introduction

There has been substantial discussion of beam quality and how it is to be quantified in the

high power commercial laser arena. With very high power military lasers, the concern has been

77

S. Martellucci et al. (eds.J, Laser Applications/or Mechanical Industry, 77-97. © 1993 Kluwer Academic Publishers.

78

mainly to provide a wave front that would provide a near diffraction limited Airy disc in the far

field. That is, how does one take the laser beam from the cavity and manipulate the beam quality

and wave front so that the power of our laser is maximized in the far field to destroy a target at

some distance away.

Our problem in high power commercial laser processing is not unlike that of high power

laser weaponry, in terms of wave front quality that has to be maintained. In the case of high

power commercial lasers, the user wants to maximize the laser power onto the workpiece. In

order to perform this function effectively, the wave front quality of the beam must be maintained

from the laser cavity through the optical train and focused on the material to be processed.

Fundamentally, we are talking about the diffraction image. "If the pupil function is a constant, i.

e., if the transmission of the system is uniform over the aperture and the system is aberration -

free, the illuminance distribution in the image becomes, in the far field,"[I]

I = 10 exp[- p\ J 2RO

(I)

where: Ro = 1.22 N.A. A. , cm, is the un-aberrated Airy disc radius at focus of the optic; P =

~2 (1.22.N.A. A.) 2 0.1743533 , cm, is the central lobe of the un-abberated Airy disc, cm

(Note: The central lobe carries 84% of the total flux density. The remainder is contained in the

outer lobes.); A. = Wave length of the laser, cm; and, N.A. = Numerical aperture = Focal

length/Diameter of optic. For the case where the optical component has a focus at infinity, the

Numerical aperture will equal [2],[3J N.A. = Diameter of optic / 2.

We must also accommodate the diffraction image of the multiple transverse modes of the

laser beam emitted from the laser cavity. The central lobe begins to diminish in amplitude as the

lobe begins to increase in radius because of diffraction effects on the wave front quality. The

diffraction of the laser mode is found from [1-3]

P = ~ 2{1.22 N.A. A. [2P+Q+I]}2 0.1743533 (2)

where p is in cm, P = see Fig. I - ordinate, and Q = see Fig. I - abscissa.

We must also take into consideration diffraction effects caused by figure error, ripl"le caused

by cooling channels, and bowing caused by axial temperature gradients across the optical

components. These effects. manifest themselves as a defocusing term <1 • The original equation,

then, is expanded to include the defocusing term, such that (1),[4-6]

79

P=~2{1.22 N.A. A[2P +Q+l]+ Loi 0.1743533 (3)

where p is in cm, lli = sum of defocus due to mirror train aberrations, cm.

The last component to be included in our equation is the decollimation angle of the laser

beam coming out of the laser cavity. Ideally, this decollimation angle will be very small.

However, as the beam travel distance increases, with a distributed system, this value can become

a significant element in the calculation .. At large distances, an un-aberrated Gaussian beam will

provide [6,7] an angular spread of 4A.,htDO, in radiants.

The diameter of the laser beam at the focusing optic, after having passed through the beam

path from the laser cavity, will follow from

2 P 2Dp

D • _ 2(0.1743533) P -1.22[2P+Q+l] ADO

(4)

where: Dp = Diameter of the laser beam at the laser cavity, cm; dw = Diameter of the beam waist

Q

0 1 2

0 e - -~ - .~ • TEMoo TEMol TEM02 TEMol-

0 ~ -1 ,-:- t ~ ~

TEMIO TEMll TEMl2

~ ... ~ - 'I~I , -2 --

TEM20 ~M21 ~M2:l

Fig. 1. Modified matrix for typical transverse mode patterns in laser output beams

80

at the focus of the wode piece, cm; D~' = Diameter of the laser beam on the focusing optic, cm;

and, Do = Diameter of the beam train optical components, cm. The beam waist diameter at the

wode piece is a very important parameter. The focusing optic will be expected to have a focal

length of some known value. All of the wave front aberrations that are presented to the focusing

optic carry right on through the focusing system. If the energy in the central lobe of the wave

front has been aberrated, i. e., diminished, the beam waist and concentration, or focusability, of

the beam will be equally increased as to the waist diameter and subsequently diminished

commensurately as to the concentration. The beam waist at the quasi focus of the refracting

focusing lens (of focal length f, in cm) will follow from

f dw = 2.44 Op' A[2P+Q+l]+rcr (5)

It is not uncommon to use refracting optics to focus the beam. As the power of the laser

increases, the usual materials used for refractive lenses become vulnerable to laser damage, and

must be replaced by reflecting optical focusing systems. In both cases, there are additional

aberrating third and fifth order terms that would come into play for the wave front aberration

scheme. The focusing optics may be subject to spherical aberration, astigmatism and field

curvature and coma. The most acute problem, however, is the first order term of defocus.

There are some conventions that are presently in vogue for characterizing the aberrations

in the system from the laser cavity to the wode piece. One of these is the M2 factor. This value

relates to the ratio of the actual beam divergence to that of the un-aberrated Gaussian beam

divergence[8]. When we provide a sample problem we will compare the M2 technique to the one

we have just described. Another technique is to apply a k value. This value is very much the

same as the M2 technique in that it is reciprocal of M2. In both cases, however, the aberrations

caused by the optical components in the beam train are not taken into consideration.

2. Equation components

2.1. RATIO OF OPTICAL COMPONENT TO LASER BEAM DIAMETER

For laser beam train optics, the numerical aperture is known to be D/2. There is yet

another set of criteria that one should observe when designing optical components for laser beam

transmission. The diameter of the optical components should be on the order of 1.63 to 2.0 times

81

the diameter of the laser beam outcoupled from the laser cavity. In the one case of a value of

1.63, we suggest that the aberrations will impair the "blur-spot image" at the best focus of the

system [1]. Those authors favoring the value of 2 times the diameter of the laser beam state: "At

large distances, the angular spread of a Gaussian beam is 4AhtDo between the e-2 points where

Do is the e-2 beam diameter at the optical system. If a Gaussian beam is truncated, or stopped

down, the Gaussian distribution gradually disappears, approaching the distribution as a

uniformly illuminated aperture. If the clear aperture of the optical system is equal to at least

twice the e-2 beam diameter, the flux density distribution of the beam is within a few percent of

a true Gaussian shape" [6,7].

Recently, H. Hugel and his team at Stuttgart University, completed some experiments that

suggested that in fact these values are correct. From Fig. 2 one can see that the relative aperture

diameter should be something between 1.75 to 2 times the beam diameter, depending upon the

amount of wave front error that can be tolerated or anticipated [9]. Fundamentally, we are

discussing the Rayleigh quarter wave limit for image quality. In this limit, if the wave front

aberration varies no more than one quarter-wavelength over the aperture of the optical system,

the image will be almost perfect. The Strehl criterion is somewhat broader in scope: the Strehl

ratio is defined as the ratio between the illuminance at the peak of the diffraction pattern of an

aberrated point-image and the illuminance at the center of an aberration-free image. A Strehl

1.5-.--------.--------,r--------,

1.4

1.3

1.2

1.1

1.0

0.9

0.8

S beam diameter F focus diameter K K-number

0.7+--------,--------,--------1 1.0 1.5 2.0 2.5

relative aperture diameter

Fig. 2. Beam quality as a function of the ratio of laser optic diameter to laser beam diameter.

82

ratio of 0.8 is equal to the Rayleigh quarter-wave limit, i.e., a reduction in the central lobe from

84% of the total energy to 68% of the total. At this point we need to look at the character of ~

for an optical beam train. A substantial amount of work. has been undertaken by Bennett [10] and

others [7,11-13] relative to wave front quality for very high power laser systems.

2.2. BASIC STREm.. CRITERIA

Bennett states, "More stringent tolerances are required to focus diffraction limited

intensities than are required to give diffraction limited resolution." For the most part, the

Marechal criterion is the basis for design in a multi-element optical system. Marechal argues

that RMS wave front error not exceed 1/14, giving a Strehl ratio of 0.8 where the Strehl ratio is

described by [14]

(6)

where: 10= Intensity without wave front error; 1= Intensity in central lobe with wave front error;

and, a= RMS wave front error, (a and A. in ~).

Bennett states that "The total RMS wave front distortion a is the sum of contributions

from the N components in the optical train . .. Thermally induced wave front errors in different

components are often correlated . .. We will assume thermally induced wave front errors

between components are correlated . . . there will be no correlation between figuring errors in

different components." We would add to this concept that the pressure ripple, thermal ripple, and

fin ripple compose the RMS wave front error and are, consequently, correlated as well. Bennett

is able to offer a relationship for the allowable RMS wave front error and states, "Assuming that

the initial figure and optical inhomogeneity errors are uncorrelated with each other or with the

thermally induced figure errors so that the squares will be additive, assigning the same figure

error budget to each component and using the Marechal criterion we then have"

(7)

where: af= RMS wave front error introduced by initial figure errors in a single component; N=

Number of optical components; OT= Thermally induced wave front error; and, A.= Wavelength,

~.

The allowable peak-to-peak initial figure error of for a mirror is described as

where: 'tfm = Initial figure error of individual mirrors;

Nm crT= L RMS(a)+.1S

N=1

83

(8)

(9)

where: N I-m= Total number of mirrors in the system; RMS(A) = Mirror ripple; and, AS =

Mirror bowing.

However, to show how much energy will reach the system end point we need to include

another correlated component for each optical element. Each optical component will have

different flux density, or fluence, to contend with and, as such, will have different temperature

gradients. Also, absorption and scattering from the coatings will be different for the various

components which may have different reflectivities. For absorption [7,14,15]

Nm { 4'1' crso = L 1-R - 2N

N = 1 <p( 11L / 11H)

o 2N}2 41t; [1.15 - 2.565 .10-6(5 .103 -A.)] (10)

where: for high index material next to the substrate, 'V = 1.0, <p = 11S ; and, for low index

material next to the substrate, 'V = l1S' <P = 11H2 , l1S = Substrate index, 11H= High index of

dielectric film, 11L = Low index of dielectric film, N = Number of high - low coating pairs, and

for scatter, i. e., diffuse reflection, in a coated optic [7,15]

If the optical components are not coated with a quarter-wave stack of optical thin films,

84

the average specular reflection from the optical surface at nonnal incidence in air will follow

from, [7, 14, 16,17]

R (n-1l+k2 [41t0'0cosS]2 1,2 ... N = (n+1)2 +k2 - A. (11)

For reflective surfaces with a laser beam impinging at an angle S , reflectance Rs and Rp (the

intensity reflectances corresponding to light polarized perpendicular and parallel to the plane of

incidence, respectively), and no the refractive index of the external medium

RS no cos2 S+(n2 + k2)(a2 + b2)_ 2nO cosS(n a - k b)

n02 cos2 S +(n2 + k2)( a2 + b2)+2 nO cosS(n a - k b) (12)

where: n = Optical index of metallic surface; k :::: Imaginary coefficient of metallic surface; no =

Index ofextemal medium, (air= 1); a = [( Jp2 +q2 +p} 2 t2; b = [( Jp2 +q2 -p)/2 t2; p = 1.0+(k2 _n2) [no sinS l(n2 +k2) t; and, q:::: - 2n k [no sinS/(n2 +k2)J2 . For copper

at 10.25 J.UIl [17] , n = 11.0 and k = 60.6.

The scatter, i. e., diffuse reflection, from the surface will follow from

. 2 Nm [41t0'0{COSS}]

O'S:::: I-N=l A.

(13)

where: 0'0 = RMS surface roughness, A; S = Angle of incidence, radians; and, A. = Wavelength

of interest, A. For optical components that are, or are not, provided with optical thin films, we can look

at the end point intensity by rearranging Eq. (7), so that

85

(14)

2.3. MIRROR BOWING - MIRROR RIPPLE

An important part of Eqs. (9) and (12) are the factors of bowing and ripple. Bowing is

precisely what one would think it to be. It is the distortion resulting from an axial temperature

gradient through the optical substrate. If one has a round optic which is simply supported, the

radius of curvature will follow from [7,18,19]

R=_t_ ~AT

(15)

where: ~= Linear coefficient of expansion, cm/cmoC; R = Sag radius, cm; A-r= Temperature

difference between front and back, °C; and, t = Substrate thickness, cm. For this discussion 0 is

considered small compared to R, which will provide [18-20] a 00 according to 00=R02/2R

where Ro = Radius of optic, cm. 00 also equals CJoRo 4/16D(I+~) , which may be written,

and the total deflection will follow from (17)

001 = QoRo 4 [2. + I ] = 00 (I +~) [2. +_1_] 8 8 D(I+~) 4 1+~

I sothatool=-oo(1+~)+oo .

4

(16)

(17)

For the case where the edges of the optic are clamped, a moment arm is developed at the

edges. Since the edges are precluded from rotating we find the moment from

M = .:.....~ _A-,,-T_D--=(c-I +--=~....::..) t

where: D= E t3 / 12( 1 + ~ 2); E= Modulus of elasticity, Nw/cm2; and, ~= Poisson's ratio.

(18)

86

The bending stress that is developed follows from

6M <rmax =2

t

which results in

<r - .:...~..,...A--,T,--E.,... max - 2(1-11)

(19)

(20)

From Eq. (16) we find there is a maximum thermal stress which is fundamentally

proportional to the coefficient of thermal expansion. the temperature difference. and the modulus

of elasticity E. We also know that the maximum stress at the edge is given by

The apparent load. then. would follow from

2 Q _ 4<rmax t 0- 3R02

The maximum deflection at the center of the clamped faceplate follows from

o - Qo R04 01- 64 D

(21)

(22)

(23)

The apparent load and maximum deflection may also be found from. M = <4JR02/8 and

<4J= 8M/Ro2 .

For a circular plate that is supported on the edge at three points 120° apart. with the load

uniformly distributed. the deflection at the center would be

0= 0.0362n 9oRo 4

D (24)

87

and RMS(A) = 00 1 /2J2. For Eq. (15), the value of O'f for the thennally induced wave front

error, would only involve thennal bowing as a result of clamping geometry for each component

in the system using uncooled optics.

When cooled faceplates are used in optical components for high energy laser systems,

there can be an additional correlated thennally induced wave front error known as ripple. This

distortion arises when cooling channels are used. There are two components for this distortion,

temperature gradient and pressure bowing. The absorbed flux density, or fluence, is partitioned

between the open channel and the web, or channel flange, so that

(25)

(26)

where: tl = Webb thickness, cm; and, t2 = Flow channel width, cm. The axial steady state

temperature gradient for that portion of the faceplate over the open channel will be, [7 ,21]

a) for continuous wave beams

(27)

b) for repetitively pulsed beams,

AT = to(21t Q Foq,)channel + 21t Q Foq,

K hc (28)

where: K= Faceplate material thennal conductivity, W/cm °C; 41= Pulse width, sec; 10= Faceplate

thickness, cm; FO= Absorbed fluence, W/cm2; hc= Film coefficient on the back surface, W/cm2

°C; and, Q is in Hz. For the most part, these channels will be rectangular and will have a ratio

of 10 / t2 > 3.0.

The bending moment will follow from Eq. (22) and the bending stress will follow from

Eq. (24). The deflection at the center of the plate which is rectangular and fully clamped will

follow from [7,18]

88

crl = Qot24 2.6.10-3 D

(29)

The web portion of the configuration will have a different gradient than the open channel. For

the web, the steady-state gradient will be [7,21,22]:

a) for the continuous wave case,

(30)

b) for repetitive pulses

(31)

where': Ag= tl(L), cm2; L= Total beam path length, cm; t3= Channel height, cm; P= (2t3 + tl),

em; and, m=.JhcP /K Ao ' cm- l . The web will attempt to grow in accordance with, cr2 = ~

AT2 t3 (1 - I! 2) where ~ = linear coefficient of expansion, cm/em °e. The last component of ripple is the pressure bending. As one might expect, this

component results from the pressure drop required to drive the cooling medium. Then,

Qo=PI- APx, (32)

where PI = Entrance pressure [23], and APx =(1.5+f Lx/DH) P V2 /2g, where: f= Friction

factor; Lx = Channel length to point of evaluation, cm; DH= Hydraulic diameter, cm; p= Density

of cooling medium grams/cm3; V= Velocity, cm/sec; and, g= Gravitational acceleration, 981.4

em/sec. Thus cr3 = 2.6.10-3 CQo t24 /D). For each optic,

(33)

The thermal pressure induced distortion will follow from crT = RMS(Aj + As. Given that these

various components of thermal induced distortion may be approximated to a reasonable degree,

we can substitute into Eq. (13) and obtain a relationship for the allowable peak-to-peak initial

figure error (-tftn.) for a mirror:

89

't - A: -~cr 2 [ 2 ]0.5

fm- 101t2N N T (34)

3. Examples

3.1. BACK SURFACE FORCED CONVECTION COOLED REFLECTIVE MIRROR

This type of mirror flow (coplanar flow) is the most commonly used technique for cooling

optical faceplates. Basically, we have a faceplate with flow channels beneath the absorbing

surface. For the most part, these channels are very small and require a great deal of pressure drop

through the mirror. In some instances, the pressure drop is so great as to cause frictional heating

of the heat transfer medium. In any event, we would use the equations shown above for

evaluating the heat transfer film coefficient for a flowing medium in the channels.

Fundamentally, then, we determine the required heat transfer film coefficient from the

amount of absorbed flux density that would be expected from the surface. Based on allowable

temperatures that we would want the mirror to operate within, we would determine the amount

of flow and the subsequent heat transfer film coefficient that such a flow would provide. Very

often the major determining factor will be the allowable optical distortion that may be provided

by the optic within the optical train. The optical distortion will result from the required pressure

drop, coupled with the axial temperature gradient through the faceplate. It is not uncommon for

the designer to use two rows of channels, one below the other, in order to reduce any distortion

resulting from overall bending of the optical component due to the axial gradient through the

whole optic. A certain periodic structure also results from the difference in the faceplate

immediately over the channel and the faceplate portion that is over the channel wall. These

various deflections have been described by Palmer [20].

The pressure drop through the system, knowing the amount of flow required, with very

smooth tubes, would be determined from [7,20,23,24,25]

(35)

where: Lo= Length of the flow path, cm; Dtt= Hydraulic diameter of the flow channel, cm = 4

90

(flow cros-sectional area/wetted perimeter); V= Flow velocity, em/sec; p= Density of flowing

medium, g/cm3;g = 981.41 cm/sec2; ~k = Entrance and exit losses, generally 1.5; and, fc=

Frictional coefficient. For a Reynold's number greater than 10000, fc= 1/(1. 84 In Re)2.0, see Ref.

[25]; fc= 1/(1.84 In Re - 1.64)2.0, see Ref. [26]. For a Reynold's number less than 2000, fc=

64IRe, see Ref. [25].

We already know the temperature gradient through the faceplate, simply from

AT= Po t K

(36)

where: FO= Absorbed flux density, W/cm2; t= Thickness of the faceplate, cm; and, K= Thermal

conductivity of the faceplate, W/cm °e. The radius of curvature developed by the axial gradient will be found from [7,18,27]

R=_t_ f3 AT

(37)

where: R= Radius of curvature, cm; f3 = Linear coefficient of expansion of the faceplate material,

oel ; M = f3AT D (l+ll)/t; Po=8M/(1 +1l)(y/2)2; and, A=2.6 . 10-3 Po y4/D. The pressure at the point of evaluation is found from [7,20]

(38)

where: PI= Inlet pressure to the mirror, Nw/cm2; and, Apo= Pressure drop at the point of

evaluation, Nw/cm2.

The deflection at the center of a rectangular element in a fully clamped geometry is found

from an equation similar to Eq. (30) [7,18]:

4 P y4 00 (-l)(m-l)/2 Al =-- L cos (m1tx/y)

1t5 D 5 m=1,3,5.. m (39)

. 1- cos(m1tx/y)+ sinh(m1tx/y) [ (2+um tanhum) m1tx/y ]

2 cosh urn 2 cosh urn

For a ratio of channel width to channel length ;::: 3.0, A 1 in Eq. (39) can be expressed as we have

shown before: Al = 2.6· 10-3 p(y)4 / D, where: D = E t3/l2 (1-1l2); y = Channel width, cm;

91

E = Young's modulus, Nw/cm2; J.L = Poisson's ratio; and, Om = mtcz/2y, (where, y= The short

leg of the rectangular side, cm, z= The long leg of the rectangular side, cm) .

The amount of flux density absorbed over the vertical channel wall will provide a different

temperature gradient than that provided by the absorbed flux density over the span of the

faceplate. We can evaluate the temperature gradient from the faceplate to the bottom of the

vertical channel wall using an analog of the wall acting as a fin with no heat transfer from the

end [22]:

[ e mx e -mx 1

T - Too = (Ts - Too) 2mL + 2mL l+e 1-e

(40)

where: m = .Jhc P / KA ; L = Height of the channel wall, em; P= Perimeter of the wall, cm; A=

Area of the wall, cm 2; he = Film coefficient in the channel, W /cm 2oC; and, Ts = Temperature at

the back of the faceplate, at the water interface, °C. The amount of distortion caused by the

growth of the channel wall will follow from

(41)

The total amount of distortion to give the Root Mean Square distorsion of the optical faceplate,

will be

(42)

We can use an example to describe how all of these equations come to together to provide

a beam quality budget through an optical beam train. We will use a diamond turned copper

optical faceplate set of optics that are not coated with an optical thin film quarter-wave stack.

The laser will be a 5.5 kW stable resonator design with an outcoupled beam diameter of 2.765

cm and a decollimation half angle of 10-6 rad. The beam mode will be TEMoo. The optical

components in the beam train will be 6.0 cm in diameter. We will include four optical

components, excluding the focusing optic. The RMS surface roughness of the diamond turned

optical components will be 20 A. We will be using cooling channels in the optical faceplate and

assume that the channels are 0.100 cm wide by 0.100 cm deep and the channel wall is 0.02 cm

thick. Structurally, these dimensions are probably not very sound; however, we are only using

them for this example. We will assume an arbitrary water velocity of 1000 cm/sec. Our hydraulic

diameter will be 0.10 cm. Our optical components will be fully clamped at the edges. The

92

faceplate will be 0.10 em thick. We will assume that the flow is fully developed. The absorbed

flux density on the optical components that are at nonnal incidence will be found from our

equations such that,

Ab -I {I (n-l)2+k2 [41t0"0cose]2}_812W/ 2 sOIptance- 0 - 2 2 - - . cm (n+l) +k A

However, most of our optical components will have an angle of incidence of 45°. Consequently,

our absorptance will come from Eq. (12), so that

n02 (a2 +b2) + (n2 +k2) cos2 e-2 nO cose (n a + k b) Rp=--~~~~~~~--~--------------­

n02 (a2 + b2) + (n2 + k2) cos2 e+2 nO cose (n a + k b)

where: n = Optical index of metallic surface; k = Imaginary coefficient of metallic surface; no =

Index of external medium, (air= 1); a = [( Jp2 +q2 +p)/2 t2; b = [( Jp2 +q2 p)/2 t2; p = 1.0+(k2 +n2) [no sine l(n2 + k2)y; and, q = -2nk [no sine l(n2 + k2) y. For copper

at 10.25 J.lIIl [17] n = 11.0 and k = 60.6.

Undertaking the calculation reveals that, Rs = 0.992 and Rp = 0.984, giving an Ravg of

0.988, which is close to the value shown by Drummeter and Hass in their CUIVes for the

reflectivity ofvapor deposited copper [17]. Lynch and Hunter [28] provide a somewhat different

value of nand k from their measurements, i. e., n = 10.8 and k = 47.5. Using these values, Rs =

0.987 and Rp = 0.975, providing an Ravg. of 0.981 . For this analysis, however, we will stay

with the values supplied by Drummeter and Hass. For our 45 ° mirrors the absorptance will be

equal to 9.85 W/cm2, i.e.,"" 10 W/cm2. We see that the absorptance on our mirrors increases.

Even though the specular reflectance of the mirror is pretty much the same as that of the nonnal

incidence mirror, the diffuse reflection, or scatter, has gone down. As a consequence, the amount

of flux density absorbed by the mirror increases. Bennett and Bennett state, "As the angle of

incidence increases, the effective size of the surface irregularities, 0"0 cose II. , decreases and

therefore the coherent reflectance increases. As a result, a surface which exhibits no specular

93

reflectance at nonnal incidence may be a good reflector at grazing incidence." [16]. For this

analysis we will assume that we are only absorbing 10 W/cm2. The Reynold's number in the

flow channels through the mirror will be [7,20,21]

VDp NRe=-­

v (43)

In our example (V =1000 cm/sec, D = 0.100 cm, p =1.0 g/cm3, v = 9.87 .10-3 g/cm sec) NRe =

10.13-103. From our equations, then, we will find the film coefficient for our flowing medium.

Thus, for water entering at 21.3 °c, using the Dittus-Boelter equation [29]

(44)

for liquids cooling values of Reynold's number and Prandtl's number, hc = 4.24 W /cm 2 °c .

Since we are absorbing 10 W/cm2, the temperature gradient at steady state across the

water laminar boundary layer will be 2.357 0c. The temperature gradient across the channel span

of the faceplate, according to our equations, will be 0.255 °C. The temperature gradient through

the channel wall from our equation will follow: T - Too = (T - Too) [emx/(1+e2mL)+e-mx/

(1_e2mL)] = (Ts -Too) (0.450+0.450) = 23.42 °C; and, m = .Jhc P / KA = 4.674, with the

following values of: L = Height of the channel wall, 0.10 cm; P = 20.20 cm, perimeter of the fin;

A = 1.0, cm2 area of the fin; he = 4.24 W/cm2 °C, Film coefficient in the channel; and, Ts =

23.66 at the water interface, °C.

The temperature gradient from the channel span to the bottom of the channel wall (fin)

will be 23.66 - 23.42 = 0.236 °C. The coefficient of expansion for copper is [7], 16.6 .10-6 °C-l.

The Poisson's ratio for copper is 0.3. The Young's Modulus is 11.7 .106 Nw/cm2. The radius of

curvature developed by the axial gradient is found from our equations, so that R = t / ~AT; M =

~ AT D (l+u) 1 t = 13.032 Newtons; Po = 8M 1 (1+ /l)(y/2)2 = 32.079 .103 Nw/cm2; D = 1071

Nw cm; and, A= 2.6 .10-3 Po y4/D = 7.788.10-6 cm.

For the purpose of this example, we will assume that the pressure in mirror at the point of

evaluation is 250 Nw/cm2. From our equations, Al= 2.6 .10-3 P y4/4 = 6.069 .10-8 cm. The

expansion of the channel wall will follow from our equation, so that A2 = (T - Ts) ~ L = 2.745 .

10-7 cm.

The Root Mean Square will be RMS = (A+A1+A2)/2"2 = 0.0298 Jlffi. The allowable

peak-to-peak initial figure error for each mirror will follow from Eq. (14), so that,

94

'tfm = J'J..2 /201CN - 2aT2 / N = 0.6686 ~ for each of the four mirrors in our system example. From Eq. (8) we find the allowable figure

error for our budget, so that af = 'tfm/...J2 = 0.47277 ~. Unfortunately, because these are copper

mirrors, it is unlikely that such a figure can be obtained. It is much more likely that the mirrors

will have a figure 5 to 10 times the required amount From Eq. (7) we can see the impact on the

loss to energy distribution in the central lobe of the wave front.

If we are able to maintain the figure of 0.472776 ~ on each mirror and the pressure

ripple and pressure bowing are not greater than calculated, the flux density in the central lobe

would be I = 0.6846 10 • Looking at our original equations for distribution of energy in the central

lobe, p= ~2{1.22 N.A. 'J..[2P+Q+l]+ Iaf(0.1743533) we would find for our example that

the original value of p would be P = 2.291 .10-3 em .

The total value of Ia will be 2.0104 ·10-4 cm for the mirror figure and thermal ripple and

bowing for the four mirrors. The total value for p will be 2.4892 .10-3. The value of Ro is =

3.8796.10-3 cm ; then, from Eq. (1) exp (_p2/2Ro2)= 10 0.814.

We can tum the problem around and see how much figure error we could stand in order to

achieve 64% of the energy in the central lobe. Then, 8 = 11.16 ~ or a little over one

wavelength at 10.6 ~ .

We see from our example that the Strehl ratio is very much more conservative than using

the central lobe energy distribution technique. However, if the figure error is greater than in our

example, or there is loss of reflection due to oxidation on the mirrors so that there is more

absorption than calculated, we would see a substantial diminution in the energy distribution in

the central lobe. For example, if the figure error were on the order of 2.233 .10-6 for each

mirror, the total budget for figure error and thermal ripple and bowing would be 9.0512 ·10-4

em. The value of p would be 3.19612.10-3 cm and, 10 eXp(-p2/2Ro2) = 10 0.7122. It would

seem, then, that we can take the aberrations that are created in the figure error of the optics and

the thermal ripple and bowing to distort the diffraction image from the Strehl Criteria and extract

this information for use in the central lobe technique. In this fashion, it would seem, we can

remove some of the conservatism of the Strehl criteria.

3.2. FOCUSING OPTIC

As we have suggested earlier, the aberrated wave front presented to the focusing optic

carries through the focusing optic to the work piece. The diameter of the laser beam at the work

95

piece will follow from Eq. (4) : for our example, then, the laser beam diameter on the focusing

optical component will be D~' = 1.0832 D~ = 3.0 cm.

The beam waist at the focus of the focusing optical element will follow from our equation, such that dw = 2.44 A[2P + Q + 1] ID~+ lli . If we have a focusing element with a focal length

of 12.5 cm, the beam waist for our example will be dw = 0.01097 cm at the best focus, and

81.5% of the energy is contained in a central lobe that has a p' = 3.5084 ·10-3 cm.

We can see from our example that we would have a very good beam quality if we had a

1EMoo beam with optical components with a 0.4728~ or better figure error and our absorption

in the cooling optics were as little as the example shows. We can now look at the M2 concept

and see wehether its prediction are consistent with the energy in the central lobe technique.

Fundamentally, the M2 concept is a ratio of the actual beam waist diameter at the focus of the focusing optic compared to that of a diffraction limited un-aberrated beam waist. Steen suggests

that the ratio should look like [8]: DW /2f = Actual; 2Mt(4fA11tD~) = Gaussian. For our problem,

then, DW!2f = 0.120 and 4fA11tD~= 6.10142 .10-3, so that M2 = 1.085 .

Hugel [9] uses a beam quality factor k which is equal to 1.0/M2. The wave front quality

shown in our example will diminish by 2.765/3.0 == 0.92167 on finding the reciprocal = 1.085.

That is, the beam diameter increased by a factor of 1.085 and diminished the energy in the

central lobe from 84% to 81 % for our example. It is not enough, however, to simply look at the

far field beam spread. One must look at the impact of the optical components on the diffraction

image as it transports through the beam train. Since our real concern is the amount of flux

density that is contained in the central lobe, it would seem that the best measure of beam quality

would be the ratio of the actual lobe diameter to the Gaussian lobe, i. e.,

(45)

Based on this definition, the M2 value for our example would be, M2 = 1.15 and k would be,

k=1.0/1.15 =0.8695.

Using Eq. (45) for a figure of merit for an optical beam train for a particular high power

laser system provides for a direct comparison between the flux density in the central lobe for an

un-aberrated system that can be focused on a work piece and an aberrated system of an actual

high power laser beam train. The reader will note that we used the very best beam mode from the

laser. All the other modes would provide a beam quality less than what we have shown in our

example.

96

4. Conclusion

We have shown in this chapter how one can detennine, from the theoretical model, the

image quality of a laser beam as it goes from the cavity to the work piece through the optical

beam train.

Because the far field transfonn of the image requires that the flux density be concentrated

in the central lobe of the Airy disc, any and all aberrations that detract from the central lobe,

forcing flux density into the outer rings of the Airy disc, cause a substantial problem relative to

cutting, welding, etc. In order to properly evaluate the impact of the laser cavity mode shapes

and defonnations in the optical components, it is necessary to use the percentage of flux density

in the central lobe as the criterion for detennining the beam quality of the system.

The present conventional wisdom is to use the M2 relationship or the Strehl ratio for

detennining the appropriate beam quality of a system. What our equations show is that the Strehl

criterion is probably too conservative and does not really detennine the amount of flux density in

the central lobe, except inferentially. The present M2 technique does not address the quantity or

quality of the optical components and, as a consequence, does not give a true picture of the flux

density in the central lobe at the work piece.

By combining the Strehl ratio aberration technique of the optical components with the

characteristics of the M2 technique and employing the Airy disc equations, we are able to take

the best of both techniques and provide a more accurate description of how the flux density in

the central lobe changes with the change in modes of the laser and the defonnations resulting in

the optics to cause a change in the percentage of the flux density that is in the central lobe, or

being distorted and subsequently removed to the outer rings. As the flux density is removed from

the central lobe and increases in the outer rings, the wave front of the laser beam looks more and

more like a Lambertian radiator and there is no capability of focusing the energy onto the work

piece.

In short, maintaining the highest quality in the Airy disc, which is the same as saying that

one wants to maintain as flat a surface in the optical components as possible, becomes the most

critical element in detennining the ability to use the laser beam as a working tool.

References

1. Driscoll, W. G., and Vaughan, W., (1978):, "Handbook a/Optics", McGraw-Hill Book Co., New York,NY.

2. Hardy, A. C., and Perrin, F. H., (1932):, "The Principles a/Optics", McGraw-Hill Book Co., New York,NY.

97

3. Svelto, O. and Hanna, D. C., (1982):, "Principles of Lasers", Plenum Press, New York, NY. 4. Duley, W. W., (1976):, "C02 Lasers; Effects and Applications", Academic Press, New York, NY. 5. Ready I.F., (1971):, "Effects of High Power Laser Radiation", Academic Press, New York, NY. 6. Wolfe, W. L. and Zissis, G. J., ed.,(1978):, "The Infrared Handbook", Office Naval Research,

Department of The Navy, Washington, D.C. 7. Palmer, J. R., (1990): "High Power Laser Optics; A Study in Transient Heat Transfer", Pro Se

Publishing Co., San Diego, CA. 8. Steen, W. H., (1992): "Laser Materials Processing", Springer-Verlag, London, England. 9. Hiigel, H., (1992): "Recent Development to increase the efficiency and tlrxibility of Laser

Materials Processing, published in this volume. 10. Bennett, H. E., (1976): "Thermal Distortion Thresholds For Optical Trains Handling High Pulse

Powers," in Laser Induced damage in Optical Materials: 1976, H. E. Bennett, A. H. Guenther, D. Milam, and B. E. Newnam, ed.,Nationai Bureau of Standards Special Publication No. 462.

11. Palmer, J. R., (1986): "Analytical Model For Transient Strehl Ratio Distribution In Soft X-Ray Optical Systems," Proc. International Conference On Soft X-Ray Optics And Technology, SPIE No 733, Berlin, Germany.

12. Palmer, J. R., (1987): "Hostile High Energy Visible Laser Environment Providing Destruction Of Optical Signal In Imaging System," Int'I Symposium On The Technologies For Optoelectronics, SPIE No. 867, Cannes, France.

13. Lloyd, J. M., (1979): "Thermal Imaging Systems", Plenum Press, New York, NY. 14. Wolf, M. and Wolf, E (1959): "Principles of Optics", 1 st ed., Pergamon Press, New York, NY. 15. Palmer, J. R., (1986): "Theoretical Model For Determining Temperature Transients In Multilayer

Optical Thin Films Subjected To High Power Continuous Wave And Repetitive Pulsed Lasers; Part 1- Continuous Wave," Proc. 3rd International Symposium On Optical And Optoelectronic Applied Science And Engineering, SPIE Vol. No. 655 and 656, Innsbruck, Austria.

16. Bennett, H. E. and Bennett, J. M., (1967): "Precision Measurements In Thin Film Optics," Physics of Thin Films, Vol. 4, G. Hass and R. E. Thun, ed., Academic Press, New York, NY.

17. Drummeter, L. F. and Hass, G., (1964): "Solar Absorptance And Thermal Emittance," Physics of Thin Films, Vol. 2, G. Hass and R. E. Thun, ed., Academic Press, New York, NY.

18. Timoshenko, S. and Woinowsky-Kreiger, S.(1959): Theory of Plates and Shells, 2 nd ed., McGraw-Hill Book Co., New York, NY.

19. Harvey, J. F. (1967): Pressure Vessel Design: Nuclear and Chemical Applications, D. van Nostrand, Inc., Princeton, New Jersey.

20. Palmer, J. R.,(1983): "Continuous Wave Laser Damage On Optical Materials," Optical Engineering, Vol. 22, No.4

21. Kreith, F., (1966): "Principles of Heat Transfer", 2 nd ed., International Textbook Co., Scranton, PA.

22. Kern, D. Q. and Kraus, A. D., (1972): "Extended Surface Heat Transfer", McGraw-Hill Book Co., New York, NY.

23. Schlichting, H. (1968): "Boundary Layer Theory", McGraw Hill Book Co., New York, NY 24. Denn, M. (1980): "Process Fluid Mechanics", Prentice - Hall Publishing Co., New Jersey. 25. Olson, R. M. (1967): "Engineering Fluid Mechanics", International Textbook Co., Scranton, PA 26. Idel'Chik, I. E. (1960): "Handbook Of Hydraulic Resistance", Israel Prog. For Scientific

Translations, Tel Heviv, Israel. 27. Boley, B. A. and Weiner, J. H. (1960): "Theory Of Thermal Stresses", John Wiley and Sons, New

York, NY. 28. Lynch, D. W. and Hunter, W. R., (1985): "Optical Constants Of Metals," Handbook Of Optical

Constants Of Solids, E. D. Palik, ed., Academic, Press, Inc., New York, NY. 29. Dittus, F. W. and Boelter, L. M. K., (1930): Univ. Of California Pubs. Eng., Vol. 2, Berkley, CA.

ANALYTICAL MODEL FOR EV ALUA TING DUAL CO-AXIAL REPETITIVE PULSED AND CW LASERS IN DRILLING AND MACHINING OF MATERIALS

1. R. PALMER University of Alabama Center for Applied Optics Huntsville, Alabama 35899, USA

ABSTRACT. The direction of this paper is to describe an analytical model that provides for the precision drilling and machining of materials using dual wavelength co-axial lasers. The model describes how the radial vector of the diffusion is removed when using a C02 CW field laser to heat up the local area close to melt and the injection of the repetitive pulse doubled Nd:YAG laser to force the material to evaporation. By using the two lasers in concert, the Gaussian temperature distribution is taken away from the working area of the Nd:YAG and allows for the evaporation of the material to provide a clean straight hole. The model will describe the time-temperature history of the material using the Reverse Thermal Wave Transform [1-3]. The time-temperature history, of course, is dependent upon the various materials thermal transport properties. Illustrations are provided by contrast to show the required levels of flux density and fluence for three different materials. This technique was first discussed by V. N. Anisimov et al. and gave rise to this present concept [4].

1. Introduction

When flux density, or fluence, is applied to only one surface of a work piece either a large

amount of flux density must be applied over a short period of time if the laser beam is small, i.e.,

R < 6 .JCrl. 't) . Or, the flux density must be maintained over a period of time necessary to raise

the temperature through the melting point of the material and provide for evaporation of the

volume. A basic element of the scheme proposed in this paper is to raise the temperature of the

material in a very localized area without causing the surrounding material to rise in temperature

to a point where there could be a change in phase or grain structure. In the drilling of a hole,

then, we would want to raise the temperature of the material surrounding the hole to a point

somewhat below the melting point. By maintaining the temperature of the field, the working

99

S. Martel/ucci et al. (eds.), Laser Applicationsfor Mechanical Industry, 99-111. © 1993 Kluwer Academic Publishers.

100

laser, that will drill the hole, need only provide sufficient Joules to cause the evaporation of the

volume that constitutes the hole. The temperature rise of a finite piece of material that is imposed

with a laser beam on one surface will follow from the Reverse Thermal Wave Transform.

2. Analytical model

This set of equations relates to those conditions and with those cases wherein the thickness

does not satisfy the semi-infinite plate boundary condition, i. e., where the thickness is

something less than to < 6 J(a t). For our first iteration, the temperature rise on the first

surface of the round flat plate, using a round continuous wave laser beam, for Ra < (5 -V(a t) +

Ro); to < 6 -V(a t) and Ro < 6 -V(a t) can be found using the following expression [1-3]:

2 Po J(at){ierfc(x) - ierfc(xI) + ierfc(x2)} ATI=--~~~--~----~~--~~

K

2 Po ~ { ierfc(y) - ierfc(Yl ) + ierfc( Y2)}

K

. {erfc(y) - exp (toHo+H02 at) erfc(z)}

2 PoJ(a t){ierfc(y)-ierfc(YI)+ierfc(Y2)} +--~~~~~~--~~----~~

K

2 PoJ(a t){ierfc(x)- ierfc(xI) + ierfc( X2)} +--~~~~~~--~~----~~

K

. {erfc(y) - exp (toHo+Ho2 a 'C) erfc(z)}

(I)

where: Fo = absorbed flux density, W/cm2; a = theffilal diffusivity, klpCp, cm2/sec; p= density,

glcm3; Cp = specific heat, Jig °C; K = theffilal conductivity, W/cm °C; to = thickness of the

substrate, or workpiece, cm; t = laser run time, sec; z = [to / 2Jret + HO Jret]; HO = 0.5

..[itiCii ~ to, cm-I; x = 1I..[i; Xl = RO / 2Jret; x2 = Ra I 2.JOi; y = to I 2.JOi; YI =

~t02 + R02 /2.JOi; Y2 = ~t02 + R02 /2Jret .

The temperature rise at some depth into the component at t > 0, for Ra < (5 ~ + Ro);

to < 6 ~, and Ro < 6 ~ will follow from the following expression:

2 Po '(ex.'t){ierfc(yo)-ierfc(YOl)+ierfc(Y02)} aTI = _..;...V..!...-_~-:...-:... __ :....-..:..-_---.:....---.:~ K

_ 21b~{ierfc(x)-ierfc(xt}+ierfc(x2)} K

· {erfc(y) - exp (toHo+Ho2 a 1:) erfc(z)}

101

(2)

where: YO = t / 2&; YOI = ~to2 + R02 /2& ; Y02 = Jt2 + Ra2 / 2&; 0 < t S; to; Ro = radius of the laser beam, cm; and, Ra = radius of the component or workpiece, cm .

The second iteration, then, will come from the following expressions: for the first surface

the required temperature rise for the particular thermal mass, given the appropriate finite

geometry, for Ra < (5 ,,(a 1:) + Ro); to < 6 ,,(a 1:) and Ro < 6 ,,(a 1:) will follow from the

following expression:

21bJ(a 1:)'PO {ierfc(x) - ierfc(xI) + ierfc(x2)} AT} = -.::.....:...--..:.....:..----...:........::-'-----'-::..:...:..

K (3)

21b~ {ierfc(y) - ierfc(YI) + ierfc(Y2)}

K

· {erfc(y) - exp (toHo+Ho2 at) erfc(z)}

2FoJ( a 1:)'P{ierfc(y) - ierfc(Yl) + ierfc(Y2)} +-~----~--------'-~--~~

K

2 FoJ(a 1:)'PO{ierfc(x)-ierfc(xt}+ ierfe(x2)} +-.::.....:...--~----~:....--~~

K

· {erfc(y) - exp (toHo+Ho2 at) erfc(z)}

where: z = [to /2& +HO&l; and, HO= O.sJ1t/ a1t ~ to, em-I.

The temperature rise at some depth into the component at t > 0, for Ra < (5 ~ + Ro);

to < 6 ~,and Ro < 6 ~ will follow from the following:

2 FoJ(a 1:)'Po{ierfc(yo)-ierfc(yOl) + ierfc(Y02)} A~=-~--~-~~-~~--~~

K (4)

2 FoJ( a 1:)'PO{ ierfc(x) - ierfc( Xl) + ierfc( X2)} +-~--~-----'--~-~~

K

· {erfc(y) - exp (toHo+Ho2 at) erfc(z)}

102

where: Yo = t / 2Jfii; YOl = ~t2 + RO / 2Jfii; Y02 = ~t2 + R0.2 / 2Jfii, 0 < t::;; 10 ; and, '1'0 will follow from the following Eq. (9) .

The temperature rise at some radius greater than the laser beam and less than, or equal to,

the outer boundary radius would follow from the following Eq. (5), so that: for Ro < Re ::;; ROo

and

a) for 30 > o.rclRo2 > 0.60515

ATRe =AT~RO erfc(~)+ AT(Re ~:o)~ ierfc(~) Re 4RO . Re·

b) for o.'t IRO < 0.60515

~RO AT(Re -RO)!cXt ATRe = AT - erfc(~) + 0 5 15ierfc(~)

Re 4RO· Re·

c) for 0.1: 1Ro;;::: 30

ATRe =AT~RO erfc(~) Re

where: ~= (Re -RO) / 2Jfii .

(5)

(6)

(7)

Having completed the first iteration, using Eqs. (I), (2), (5), or (6) or (7), which provides

us with the temperature distribution throughout the volume encompassed by the laser beam and

the radial peripheral fin into which the flux density is escaping through thermal diffusion, we

would then determine the enthalpy balance using [1-3l the following equation:

Ib = P Cp1t R0.2 to ATavg

'tR02rc (8)

which can be simplified to FO = AT avg (R0.1Ro)2 P Cp t / rc't, or, re-arranging, AT avg =

(ATlaser+ATfin) / 2 = FO 'tR02 / pCp t Ri. We would, then, calculate the average temperature in the volume under the laser beam and

the average temperature in the radial peripheral fin and find the average temperature in the the

complete volume, i.e.,

AJ: _.!.{LATlaserbeam LATradial} avg - 2 No. of Nodes + No. of Nodes

'fI _ AT Eq.(8) 0- AT Eq.(9)

103

(9)

Having found the value of 'Po we would proceed through the second iteration using Eqs.

(3), (4), and (5) or (6), or (7). It should be pointed out, at this point, that these equations assume

that there is no heat loss from the component or work piece. These equations can best be shown

by using an example. We will use a piece of copper that is 10 cm thick and 30 cm in diameter.

The laser beam will be 14 cm in diameter and we will run the laser for 100 seconds. The

absorbed flux density will be 100 W/cm2. The temperature profile after the second iteration will

look like the following:

t Ro 7,8 8.6 9.4 lQ,2 11,0 11.8 12,6 l3.4 14.2 15

Q 135 112 105 23 83 14 66 52 53 41 43

2 93 82 12 64 51 51 46 41 37 33 29

4 79 69 61 54 48 43 38 34 31 28 25

6 68 59 52 47 41 37 33 30 27 24 21

8 52 52 46 41 36 32 29 26 23 21 19

lQ 52 45 4Q 36 32 28 25 23 20 18 16

We have a total of 66 nodes where we have calculated the temperature rise. We have a

total sum of all the temperatures which is 3147.54 °C AT and we divide the total temperature by

the number of nodes which will provide an average temperature in the component of 47.69 °C

AT . Adding together, [(135 + 93 + 79 + 68 + 59 + 52)/6 + 47.69]/2 = ATavg = 64.25. The required average temperature is calculated from Eq. (8) so that FO=pCp Ra2 t 1 't Ro2 . For

copper p =8.92 glcm3; Cp = 0.38 W sec/g °C; t = 10 cm, Area of the laser beam = 153.938

cm2, and volume of the wOrkpiece is 7068.583 cm3. The required AT calculated from Eq.(8) is

AT = 100 ·100 ·153.94 1 8.92 ·0.38 . 7069 = 64.245 °CAT vs 64.25 °CAT ; then,

fb = ATavg{Ra I RO)2p Cp t /t, Fo = 100 W/cm2 which provides an accuracy of 99.99 %.

The value of '1'0 after the first iteration was 0.428429 from Eq. (9).

The next set of equations will deal with the repetitive by pulsed laser. Using the equations

for the Reverse Thermal Wave Transform, we can find the temperature profile for the workpiece,

or component, from the following expressions.

104

The temperature rise of the first surface, for the first iteration, will follow from the

following expression [1-3]: for Ra <{5 ...j(a n-1 11) + Ro } or to :::; 0.6 -Va n-1 11 , Ro < 6

-V(a n-1 11) ,

~Tl = 211 FOJU {ierfc(x) _ ierfc(xJ) + ierfc(x2)}' {n-O.5 _ (n-1 - 4»0.5} K

- 4 1tnIj>Fa-Va n-1 Tl{ierfc(y) - ierfc(yJ) + ierfc(Y2)} K

{erfc(y)- exp (toHo + H02 a n-1 Tl) erfc(z)}

+ 4 1tnIj>F0-Va n-111 {(ierfc(y))-(ierfc(yI))+(ierfc(Y2))} K

+ 4 1tnIj>F0-Va n-111{(ierfc(x)) - (ierfc(xIl) + (ierfc(x2))} K

{erfc(y) - exp (toHo+Ho2 a n-111) erfc(z)} .

(10)

The temperature rise at any depth into the workpiece to;;::: t > 0, for Ra <5 -Va n-1 11+Ro

to:::; 0.6 -Va n-1 Tl , Ro < 6 -Va n-1 Tl, will follow from the following expression:

~Tt > 0 = 4 1t n Ij> Fa -Va n-1 11 {ierfc(y) - ierfc(yJ)+ ierfc(Y2)} K

+ 4 1t n Ij> Fa -Va n-1 Tl{ierfc(x) _ ierfc(xI) + ierfc(x2)} K

{erfc(y) - exp (toHo+H02 a n-111) erfc(z)}

(11)

where: Fo = peak pulse absorbed flux density, W/cm2; a = thermal diffusivity, KlpCp ,

cm2/sec; p = density, glcm3; Cp = specific heat, J/gOC; K = thermal conductivity, W/cmoC; to =

thickness of the substrate, or workpiece, cm; n-1 11 = laser run time, sec; Ij> = pulse width, "Full

Width Half Band", sec; z = [to / 2~an-l11 + Ho -Va n-1 11 ]; Ho= 0.5 ~1t / an-I" ;;::: to'

cm-1; " = number of cycles; n = Hertz rate, Hz; Yo = t/2 ~a n-1" ;Yol=

Jt2 + R02 /2Ja n-1Tl ; Y02 = Jt2 + Ra2 / 2~a n-1" ; 0 < t :::; to ; Ro= radius of the laser

beam, cm; Ra = radius of the component or workpiece, em; x = 1. 0/ .fit; xl = RO / 2J an-I,,;

x2 = Ra/2Jan-1,,; y = to! 2~an-l,,; Y1 = J t02 +R02j 2Jan-1,,; and, Y2 =

Jt02+Ra2 j .2Jan-1".

105

Following the first iteration, again, we would balance the enthalpy as we did with Eq. (8)

and (9), so that [1-3] :

(12)

We would, then, calculate the average temperature in the volume under the laser beam and

the average temperature in the radial peripheral fin and find the average temperature in the the

complete volume, i.e.,

AT 1 {rAT laser beam rAT radial } avg = "2 No. of Nodes + No. of Nodes

'Po = AT Eq. (12) AT Eq. (13)

(13)

The second iteration would follow, then, using the value of ATavg. for the calculation.

The final temperature distribution for Ra < 5 ,,(ex 0-111) + Ro. for 10 :c;:; 0.6 "ex 0-111 ; Ro < 6 ,,(ex 0-111) would follow from the following equations: [1],[3]

ATI = 211 FoJa 'PoWerfc(x) _ ierfc(xt) + ierfc(x2»}{0-0.5 _ (0-1 - cp)O.5} (14) K

_ 4 1tOcpFo 'PO "ex a-111{ierfc(y) _ ierfc(yt) + ierfc(Y2)} K

·{erfc(y)- exp (loRo + Ho2 ex 0-111) erfc(z)}

+ 4 1tOcpFO 'PO "ex 0-111 {ierfc(y) _ ierfc(yI) + ierfc(Y2)} K

+ 4 1tO «llFo 'PO "ex a-1 11{ierfc(x) - ierfc(xO + ierfc(x2)} K

{erfc(y) - exp (loRo+Ho2 ex 0-111) erfc(z)}

106

The temperature rise at any depth into the workpiece to ~ t > 0 , for Ra < 5 ..J(a a-1 11) + Ro '

to::;; 0.6 ..Ja a-1 11 , Ro < 6 ..Ja a-1 11, will follow from the following expression:

ATt> 0 = 4 na«pFo '1'0 ..Ja a-IT] {ierfc(y) - ierfc(Yl)+ ierfc(Y2)} K

+ 4 nO«PFo '1'0 ..Ja a-1 T]{ierfc(x) - ierfc(xI) + ierfc(x2)} K

{erfc(y) - exp (toHo+Ho2 a a-111) erfc(z)}

(15)

The temperature rise at some radius greater than the laser beam and less than or equal to

the outer boundary radius would follow from Eq. (16), so that for Ro< Re ::;; Ra and:

a) for 30 > a 0-1T] IR02 > 0.60515

(16)

b) for an-1T]1R02 < 0.60515

(17)

A TRe = AT fEfr erfc(~) VRe (18)

where: ~ = (Re - RO)/ 2 ~C1. 0-111

We can begin by evaluating the problem by finding the latent heat of evaporation, either

from the Table I in Section 5 "Numerical Appendixes" or with Trouten's rule [9), such that,

AH A H* = - = 88 J/g-mole K

Tb (19)

the gram molecular weight of aluminum is 26.98 gig-mole and H*= 88(2520+273.15)/26.98 =

9110 JIg vs the Table which provides a value of 2.9343 .105/26.98 = 10,876 J/g. The amount of

107

energy necessary for the adiabatic entropy increase which brings the thennal mass to its boiling,

or vapor state, can be calculated [1,5,6,7] from the latent heat of vaporization as

Evap=mH* (20)

where: m = 1t Ro2 P to, and Ro = radius of the hole to be drilled, cm. The field laser must bring

the material up to near the boiling point before we can make the working laser useful. The CW

field laser must provide Qo = m"'Cp'" AT (in Joule) and the working laser must provide Qo =

m*H* in (Joule) . The reciprocal velocity of melt through time for the repetitive pulse laser, after

the material has reached the vaporization time, will be found [10]

'" sec p (Cp ATO + m H ) -= cm 2 dl Fopeak cP

(22)

where: ATo = boiling temperature - melting temperature. The total time to bum-through the

material will be 't total = time to reach vaporization temperature from the equations of the

Reverse Thermal Wave Transform and Eq. (22) to .

3. Experimental model and examples

The initial experiment that was undertaken related to the use of a COz field laser that

outcoupled 100 W in a 1.0 cm beam. The worldng laser was a doubled Nd:YAG laser that

outcoupled 10 J/pulse at a rate of 100 Hz and a pulse width of 100 n see sec-FWHB. The beam

diameter of the working laser was 0.025 cm at 1/e2. The work piece was 0.2 cm thick and had a

diameter of 5.0 em. The sample was painted with a black absorbing high temperature paint to

give approximately 95% absorption on the surface. The average flux density of the field laser

was Fo avg= 1o'0.s2x or 127.32 W/cm2 ; the peak fluence of the working laser is Fo peak=

Io/O.0252xcp or 5.093 ·1010 W/cm2 .

We find from our equations that the aluminum under the field laser beam achieves melting

temperature at 250 sec, or 4.16 min, of run time assuming a starting ambient environment of 20

DC. The volume of material to be drilled will be 0.0252 . 1t . 0.2 = 0.0003927 cm3 the number of

grams to vaporize, then, will be 2.67 glem3 . 0.0003927 cm3 = 0.0010485 g. The required

number of Joule to vaporize all of the material will be 10.876 .103 JIg· 0.0010485 g = 11.387 J.

From our equations, we find that the C02 laser was able to bring the field up to 660°CAT in 250

108

sec, i.e., 4.167 min, the working laser will take about 250 cycles, i.e., 2.5 sec to achieve the

temperature of 2520°C, and the bum through will come at 2.7391 ·10-4 seconds following that

time. The temperature profile in the workpiece at the end of the 250 second field laser run is:

t Ro 0.95 1,40 1.85 2.3 2.75 3.2 3.65 4.1 4.55 5.0

0 641 458 311 318 281 253 230 212 197 184 112

0,04 fl38 45fl 310 31fl 212 251 222 211 126 183 111

0,08 fl31 455 362 31fl 212 251 222 211 126 183 111

0.12 fl3fl 455 3fl2 31fl 212 251 222 211 126 182 171

O.Hi 636 454 368 315 278 251 222 210 125 182 111

0.20 635 454 368 315 278 250 228 210 195 182 111 The actual time that was measured was 253 sec to turning off of the field laser and turning on the

working laser, and the bum-through time, after turning the field laser on, was on the order of

2.83 sec. We then attempted to bum-through the same material with just the field laser only.

We found that it took 985 sec for the laser beam to come through the back surface.

4. Conclusion

The analytical model reasonably predicts the time temperature history of the dual

wavelength relationship of two co-axial lasers working in concert. This technique will allow for

maintaining the diameter of the hole from front to back, instead of the Gaussian curve that one

achieves with just one laser: the superior technique, however, would be to have the field laser

working on both the front and back surface. That will have to be the subject of a different paper.

If the part becomes sufficiently thick, however, one will have to have a laser beam on the back

surface concentric with the laser beam on the first surface.

5. Numerical appendixes

Table I Values of density, gram molecular weigh, boiling temperature, melting temperature and latent heat of vaporization from Refs. [1,2,7] and [8]

Material p(g/cm3) g-mol wt T °C Boil T °C Melt kJ/Mole-LH Evap Aluminum AI 2.699 26.98 2520 660 293.43 Gold Au 19.30 197.2 1564 1063 335.03 Beryllium Be 1.83 9.013 472 1278 292.41

109

Table I (cont'd)

Bismuth Bi 9.81 209.0 1564 271 178.74 Cadmium Cd 8.642 112.41 1040 321 99.54 Cobalt Co 8.800 58.94 2928 1495 376.60 Chromium Cr 7.200 52.01 2672 1890 344.30 CopperCu 8.92 63.54 2566 1083 300.29 Iron Fe 7.86 55.85 2900 1535 349.56 Manganese Mo 7.30 24.32 2062 1260 127.40 Magnesium Mg 1.738 54.93 1107 651 225.90 Molybdenum Mo 10.2 95.95 4607 620 592.45 Nickel Ni 8.90 58.69 2914 1455 370.30 Lead Pb 11.30 207.21 1750 327 177.80 Tin So 7.310 118.70 2623 232 296.20 Tantalum Ta 16.6 180.88 5365 3027 761.91 Titanium Ti 4.50 47.90 3289 1800 420.91 Vanadium V 6.10 50.95 3409 1710 451.87 Tungsten W 19.35 183.92 5900 3370 824.25 Zinc Zo 7.133 65.38 911 419 115.56 Silver Ag 10.50 107.88 1950 961 250.63 Rhodium Rb 12.40 102.90 3727 1985 493.70 Platinum Pt 21.45 195.23 4300 1774 509.60 Zirconium Zr 6.49 91.22 2900 1900 581.60

Table II. Value of specific heat, thermal conductivity and thermal diffusivity from Refs. [1,2,7] and [8].

Material Cp(Joules/g-°C) K (W/cm-°C) a(cm2/sec) Aluminum Al 0.870 2.37' 0.9800 Gold Au 0.140 3.18 1.2000 Beryllium Be 1.884 1.61 0.4670 Bismuth Bi 0.121 0.086 0.0720 Cadmium Cd 0.2302 0.921 0.4470 Cobalt Co 0.385 0.698 0.9500 Chromium Cr 0.50 0.950 0.2500 CopperCu 0.380 3.92 1.1400 Iron Fe 0.50 0.750 0.2000 Manganese Mo 0.486 0.078 0.0220 Magnesium Mg 1.026 1.536 0.8615 Molybdenum Mo 0.250 1.34 0.5200 Nickel Ni 0.45 0.850 0.1800 Lead Pb 0.130 0.350 0.2300 Tin So 0.226 0.64 0.3800 Tantalum Ta 0.140 0.570 0.2500 Titanium Ti 0.560 0.210 0.0830 Vanadium V 0.498 0.310 0.1020 Tungsten W 0.130 1.650 0.6600 Zinc Zo 0.226 0.640 0.3800 Silver Ag 0.230 4.200 1.6800 Rhodium Rb 0.248 0.900 0.2920 Platinum Pt 0.130 0.720 0.2500 Zirconium Zr 0.290 0.220 0.1100

110

Table III.

(X) Interval

o to 0.299 0.30 to 0.5499 0.55 to 0.999 1.0 to 1.599 1.6 to 2.0 2.001 to 2.4

(X) Interval

o to 0.299 0.30 to 0.5499 0.55 to 0.99 1.0 to 1.599 1.610 2.09 2.1 to 2.5

(X) Interval

o to 0.299 0.30 to 0.5499 0.55 to 0.999 1.0 to 1.599 1.6 to 2.0 2.001 to 2.4

(X) Interval

o to 0.299 0.30 to 0.5499 0.55 to 0.999 1.0 to 1.599 1.6 to 2.1 2.001 to 2.4

References

2 ierfc, 4i 2 erfc, 6i 3 erfc and erfc functions.

2 ierfc (X)

1.1284 - 1.998(X) + 1.11(X)2 1.115074 - 1.90863(X) + 0.9543iX)2 1.012 - l.5525(X) + 0.64346(X) 0.6271- 0.76914(X) + 0.241964(:s2 0.21573 - 0.21165(X) + 0.0525(X) 0.05124 - 0.041425(X) + 0.008393(X)2

erfc(X)

1.0 - 1.1408(X) + 0.1377(X)2 1.0167 - 1.2658(X) + 0.38269(X)2 1.0394 - 1.35775(X) + 0.47606(X)2 0.84562 - 0.98479(X) + 0.29598(X)2 0.359744 - 0.34066(X) + 0.081586(X)2 0.045779 - 0.032871(X) + 0.005892(X)2

4 i2erfc (X)

1.0171964 - 2.497286(X) + 2.4798872(Xl 0.74768 - 1.21279697(X) + 0.522727(X) 0.9441262 - 1.854143(X) + 1.049524(X)2 0.3642714 - 0.4490714(X) + 0.1417143(X)2 0.09558 - 0.093086(X) + 0.022857(X)2 0.021334 - 0.0174143(X) + 0.0035714(X)2

6 i3erfc (X)

0.5635655 - 1.457276(X) + 1.289495(Xl 0.517278 -1.133698(X) +0.715885(X) 0.3797907 - 0.6568938(X) + 0.2997376(Xl 0.1496087 - 0.1882095(X) + 0.059970(Xi 0.02861 - 0.02754(X) + 0.006655503(X) 0.021334 - 0.0174143(X) + 0.OO35714(X)2

1. Palmer, J. R. (1990): "High Power Laser Optics: A Study In Transient Heat Transfer", Pro Se Publishing, San Diego, CA .

2. Palmer, J. R. (1992): "Transient Heat Transfer in Flat Plates; Vol. I - Continuous Flux Density", Pro Se Publishing, San Diego, CA. (In Process).

3. Palmer, J. R. (1993): "Transient Heat Transfer in Flat Plates; Vol. I1/- Repetitive Flux Density", Pro Se Publishing, San Diego, CA .(In Process).

4. Anisimov,V. N., Arutyunyan, R.V., Baranov,V. Yu., Bolshov, L.A., Velikhov, E. P., Dolgov, E. P. Ilyin, A.I., Kovalevich, A. M., Kraposhin, V. S., Malyuta, D. D., Matveeva, L. A., Mezhevov, V. S., Pismennyi, V. D., Sebrant, A. Yu., Stepanov, Yu. Yu. and Stepanova, M. A. (1984): "Materials Processing By High-Repetition-Rate Pulsed Excimer and Carbon Dioxide Lasers," Applied Optics, Vol. 23, No.1.

111

5. MacDougall, F. H. (1936): "Physical Chemistry", The MacMillan Company, New York, NY. 6. Castellan, G. W. (1971): "Physical Chemistry", 2nd ed., Addison-Wesley Publishing Company,

Reading, Massachusetts. 7. Lyman, T. (1961): ed., Metals Handbook, Vol. I, 8th ed., "Properties and Selection of Metals,"

American Society For Metals, Metals Park, Ohio. 8. Hodgman, C. H. (1952): ed., Handbook of Chemistry and Physics, 34 th ed., Chemical Rubber

Publishing Co., Cleveland, Ohio. 9. Gaskell, D. R., (1981): "Introduction To Metallurgical Thermodynamics", 2nd ed., McGraw­

Hill Book Co., New York, NY. 10. Schneider, P. J.,(1985): "Conduction," Handbook of Heat Transfer Fundamentals, Section 4-1,

2nd. ed., W. M. Rohsneow, J.P. Hartnet, and E.N. Ganic', ed. McGraw-Hill Book Co., New York, NY.

LASER CUTTING AND WELDING APPLICATIONS

LASER MACIDNING OF COMPOSITE MATERIALS

F. JOVANE Istituto Sperimentale per Ie Macchine Utensi/i del CNR Via A.M. Ampere 56,1-20131 Milano, Italy

A.DIILIO Dipartimento di Energetica, Universita di L'Aquila Monteluco di Roio, 1-67040 L'Aquila,ltaly

V. TAGLIAFERRI Istituto di Ingegneria Meccanica, Universita di Salerno Fisciano, 1-84084 Salerno, Italy

F. VENIALI Dipartimento di 1ngegneria Meccanica,Il Universita di Roma "Tor Vergata" Via O. Rainwndo,l-00173 Roma,ltaly

ABSlRACT. In this chapter the fundamental technological aspects of the laser machining of composite materials are reviewed, including: laser material interaction; laser cutting performances; maximum cutting speed evaluations; cut quality parameters; and. evaluation of damage.

1. Introduction

The experimental results available in the literature [1-6] support the usefulness of laser

technology in cutting Glass Fibre Reinforced Plastics (GFRP) and Aramid Fibre Reinforced

Plastics (AFRP) so far as quality. possibility of automation and high cutting speed are

concerned; on the contrary. at present time, there are some difficulties. essentially concerning cut

quality, which remain to be overcame in the case of Carbon Fibre Reinforced Plastics (CFRP),

It is a typical feature of the production of Fibre Reinforced Plastics (FRP) that it is

possible to obtain a product almost in its final shape in the transition from the soft to hard

115

S. Martellucci et al. (eds.), Laser Applications for Mechanical Industry, 115-129. © 1993 Kluwer Academic Publishers.

116

Table 1.

Material

Resin Aramid fibres Carbon fibres Glass fibres

Thermal properties of FRP constituents: Density, p; Conductivity, K; Heat capacity, C; Diffusivity, k; Vaporization temperature, Tv; Vaporization heat, Hv

P K C k Tv Hv (g/cm3) (W/mK) (J/kg K) (cm2/s) x 10 -3 (0C) (J/g)

1.25 0.20 1200 1.30 450 1000 1.44 0.05 1420 0.24 950 4000 1.85 50.00 710 380 3300 43000 2.55 1.00 750 4.6 2300 31000

material stage. This, in general, takes place by the curing of a thermoset matrix or by cooling of

thermoplastics. In classic fields of FRP applications such as the aerospace and shipbuilding

industries, in which the production of single parts only in small quantities is common, the

necessity for fully automated and economic production methods does not exist. However,

because of the tendency to introduce FRPs in mass production, i.e. in the automotive industries,

automated machining, such as cutting and drilling. of cured FRPs is needed because the required

accuracy and surface quality of rim sections may not be achieved through reasonable effort in the

curing cycle [1, 2] . The structural complexity of FRP, essentially an inhomogeneous and

anisotropic material, explains its peculiar behaviour under conventional machining conditions:

for instance, it is well known [1-4] that FRP laminates exhibit poor quality cut surfaces, due to

spalled fibres, fuzzing, and delaminations, when drilled or routed by conventional tools. These

failure modes are absent in metal working owing to the completely different deformation

mechanisms and tool-material interactions. Some special tools [2, 5] and controlling systems [I,

3] have been designed for mechanical machining of FRP materials by conventional and non­

conventional methods.

The problem of the optimization of cutting parameters in order to obtain good quality

surfaces has not been completely solved; this is particularly true for AFRPs, which are difficult

to cut by mechanical action due to the toughness of reinforcing material and to the poor

interfacial bond between fibre and matrix. Laser cutting has been proposed as a viable method

for cutting fibre-reinforced materials [1,5,6].

The expected advantages of lasers depend on the thermal nature of the cutting process,

which does not involve any mechanical force applied to the material. Nevertheless, some

difficulties arise because of the difference in the thermal properties of fibre and matrix,

sometimes resulting in thermally altered zones at the kerf edge. In this chapter the major

parameters determining the cutting results, especially with respect to quality, are discussed and

explained.

117

2. Fundamental aspects of laser-composite material interaction

The effects of the laser beam on a material are generally connected with the following

characteristics of the beam and material properties: power density, wavelength of emission,

interaction time, polarization of the beam, absorption coefficient at the given wavelength,

melting and vaporization temperature, thennal conductivity and heat capacity. FRPs generally

exhibit a high absorption of infrared rays typical of those produced by C02 laser. Moreover,

their thennal properties are such that the vaporization process occurs at much lower specific

powers (103-105 W/cm2) than it does in metals [9]. The thennal degradation mechanism leading

to material removal is strongly influenced by the nature of the constituent materials (fibres and

matrix). The thennal properties of various polymeric resins, which constitute 40-60% by volume

of an FRP, are similar to each other. They are characterized by low values of thennal

conductivity, thennal diffusivity, and, decomposition heat. As can be seen in Table 1, differences

of thennal properties between fibres and matrix are extremely high for both graphite and glass

fibres while such differences are low for aramid fibres. The energy needed for the vaporization of

the fibres is higher than that required for the matrix; hence the laser power required for cutting

FRP will be strongly dependent upon the kind of fibres and their volume fraction. At high

specific powers the time to vaporize the FRP constituents is very short but, due to their different

thermal properties, fibres and matrix can exhibit very different values of vaporization times.

Theoretically, the time t that elapses before vaporization conditions are attained on the material surface is given [6] by t= KTv/4Pok where: K = thennal conductivity; Tv = Vaporization

temperature, Po = Laser Specific power; and, k = Thennal diffusivity. In Fig. 1 the vaporization

time values are reported for a typical matrix material and for three kinds of fibres. It is possible

to observe two limit conditions under constant specific power: both fibres and matrix exhibit

slightly different vaporization times as in the case of polyester resin and aramid fibres and

therefore the composite behaviour can be considered homogeneous; when fibres and matrix show

very different vaporization times (graphite-resin and glass-resin), the resin reaches its

vaporization temperature while the fibres are still unaffected. The results obtained by using a

thennal model and the thennal properties of unidirectional composites, reported in Table 2, show

that heat penetration into the laminae and extension of the heat-affected zone are larger by almost

an order of magnitude for CFRPs due to the high conductivity of fibres [5]. For 0/90 laminates

the thermal degradation is depicted in Fig. 2 as a function of the interaction time: this is the time

that the laser beam illuminates a zone of material, and is calculated as the laser beam diameter

to the feed rate ratio. In Fig. 2 we observe that the depth of the layer reaching the resin

118

~10·w-"------------------------------~

E o "­~10· "-'"

10'" 10 -l (S1) Interaction time 10 t

Fig. 1. Specific power of the laser beam versus the time to vaporize for the FRP constituents.

vaporization temperature, inside a lamina oriented along the kerf, i.e. perpendicularly to the

thermal flow, is much smaller than that of a lamina oriented perpendicularly to the kerf. At lower

speed an increase of heat penetration occurs for both 0° and 90° laminae and a larger difference

of penetration between the two directions arises. AFRPs have low thermal conductivity and

therefore the depth of heat penetration is very small. Besides, due to their low degree of thermal

anisotropy, the thermal gradients between 0° and 90° laminae are slightly different (Fig. 3).

Consequently, for these composites, the size of the heat-affected wne will be very small, and a

more homogeneous degradation will occur inside laminae at different orientations.

I l i 1=6 ms

~W"i I~:: : : : : : :-:::-'1 1'1\1 ••••••••••• I \lUI •••••••••••

l ~\\:g~rg~ ',1/1'1

'"i 1"',111 __ -

; TL" T~" \ \ \ \ V _"\~ .... ' ~~ d ~ ~ d

Fig. 2. Scheme of thermal degradation for 0/900 laminates as a function of the interaction time.

119

Table 2. Thermal Properties of unidirectional FRPs: Density. p: Conductivity. K; Specific Heat C; Diffusivity. k (subscripts refer: I parallel. t perpendicular to fiber direction).

FRP

Aramid/resin Carbon/resin Glass/resin

N 'Q .. ....

a

b

P (g/cm3)

1.35 1.55 1.90

K C (yV/m K) (J/kg K)

Kl Kt

0.13 0.10 1300 25 0.59 950 0.60 0040 1000

0.3

d (mm)

Fig. 3. Thermal degradation of 0/90° AFRP (a) and GFRP (b) laminates.

3. Laser cutting results

3.1. EXPERIMENTAL DETAILS

k (cm2/s) x 10-3

kl kt

0.74 0.57 170 4.00 3.20 2.10

GFRP, AFRP and CFRP panels with different thicknesses (2 to 8 mm) and 50% fibre

content in the warp and in the weft directions were tested. Two C02 laser systems were used for

120

Fig. 4. SEM micrograph of AFRP cut surface obtained at 1 m/min.

Fig. 5. SEM micrograph of GFRP cut surface obtained at 0.5 m/min.

Fig. 6. SEM micrograph of CFRP cut surface obtained at 0.5 m/min.

121

low (0.15-0.5 kW) and for high cutting powers (0.5-3.0 kW) focused on lhe material surface:

wilh gaussian beam energy distribution. Three focal spot diameters (0.25 mm, 0.35 mm and 0.5

mm) were used to attain energy densities of 104-106 W/cm2. An inert gas jet coaxial wilh lhe

laser beam impinged orthogonally on lhe sample lhrough a nozzle 2 mm in diameter.

3.2. ARAMID FIBER REINFORCED PLASTICS

Scanning electron micrograph of a cut surface is shown in Fig. 4. Macroscopic grooving

is visible. The striations lie almost parallel to lhe cutting direction and seem to be more

pronounced at lhe entry zone of lhe beam. The surface appears uneven due to loss of material,

particularly in lhe longitudinal direction, Le. parallel to lhe laminae and this seems to be greater

at lhe entry zone of lhe beam. The longitudinal shape of lhe removed material as well as lhe

scanty presence of longitudinal fibres suggest lhat many fibres in lhe weft direction were

removed togelher wilh lhe resin. The fibres in lhe warp direction do not show a clean cut but

ralher a carbonized end. On lhe olher hand, lhe bottom of lhe longitudinal grooves appears to be

smoolh and free from any carbonized residues. At higher cutting speed lhe surfaces are more

uniform. Striations are not visible in eilher lhe direction of lhe laser beam or lhe longitudinal

direction; moreover, lhere are no differences between lhe entry and exit zones of the beam. The

weft fibres are visible while slight resin loss appears in the richest matrix zones.

3.3. GLASS FIBER REINFORCED PLASTICS

Fig. 5 shows the cut edge obtained at 05 m/min: the burnt layer resulting from the cut is

visible on lhe right while lhe lower zone is shown on lhe left of lhe figure. The latter appears to

be very irregular, a remarkable loss of matrix being evident while the fibres do not show a clean

cut. At cutting speeds> 0.5 m/min no evident improvement was observed.

3.4. CARBON FIBRE REINFORCED PLASTICS

A typical cut edge obtained at 0.5 m/min is shown in Fig. 6. The matrix loss between

laminae is much higher lhan that observed for the other materials; however, warp fibres appear to

be cut at lhe same lenglh. The bottoms of the cavities are smoolh and free from any burnt

material. Micrographic observations show lhat for a given laser power, lhe morphology of lhe cut

edge is strongly influenced by the cutting speed. The surfaces are characterized by long fibres

122

Cut quali ty

X-Z WI

Slope of the I--!-t r!ffJX-X Matrix cut surface [Ej tfJ Recession

H Wj-Wo Wo tga~~

1-/

I i y-z Wo V

Craters <:::> X-Z Wd

X-z Heat

L\\ llJ I ~ af'fected

Delaminati on zone

Fig. 7. Principal cutting quality parameters.

1.0 ,........ E 08 E·

'--'

.r:. 0.6 4-' -0 .3: 0.4

't: Q) 0.2 ~

0.0 0

Power 500 JOO 2<lO 150

50 100 150 200

Cutting speed (mm/s)

Fig.8. Kerf width at inlet, Wi and outlet, Woof the laser beam versus cutting speed.

""""'E 0.6..,--------.-----------, aeeee s - 2.0 mm

E EIEIEE£I S = 3.3 mm '-" ttHf S = 4.5 mm

::5 0.4 :2 ~

Q)

010.2 o E o o

0.0 -+----.----,-----.--.----r--...,.----.---; 50

Cutting

Fig. 9. Heat-affected zones W d at the inlet of the beam versus cutting speed.

123

leaning out of the matrix and surrounded by large zones with loss of material. As the cutting

speed increases, this morphology tends to disappear.

4. Quality assessment

4.1. EXPERIMENTAL DETAILS

As we have previously observed, the morphology of FRP cut edges is characterized by

fibres pultruded by the matrix surrounded by large zones with loss of material. The kerf width is

not constant and a slope of the cut surface is evident. The presence of the charred materials (resin

and fibres) has been observed and, in some cases, thermal cracks were present. On the basis of

experimental and theoretical results obtained, we have identified the principal quality criteria, as

reported in Fig. 7. The quality parameters proposed are the kerf width at inlet Wi and outlet Wo

of the laser beam; the heat-affected zone size, W d, characterized by the presence of fibres

debonded from the matrix, or matrix recession, and thermal degradation of the fibres and matrix;

the slope of the cut surfaces; and the presence of craters, delaminations and charred materials.

The heat-affected zone is assumed to be a zone where regular transfer of load from matrix to

fibres does not occur.

4.2. INFLUENCE OF CUTTING PARAMETERS

Fig. 8 shows Wi and Wo versus cutting speed: the values of Wi and Wo both decrease as

the cutting speed increases and there is evidence of a cutting speed value up to which both Wi

and W 0 are almost constant. The Wi limit values are about the same as the laser beam spot

diameter. At higher speeds the kerf width at outlet of the laser beam tends to zero. The Wi and

W 0 values seem to be less sensitive to power changes. These results suggest that the

modification of the kerf width is more sensitive to interaction time, which can exhibit different

values by an order of magnitude.

In Fig. 9 the widths of the heat-affected zone (W d ' Fig.7) at the inlet of the beam are

reported versus cutting speed. As can be seen, the damage in the material tends to decrease as

speed increases. This behaviour can be correlated with interaction time and thermal properties of

the material by assuming a thermal mode of degradation [5]: the damage diminishes when the

energy input is lower, due to a shorter interaction time or a lower power level. At the lowest

124

user be.m

I

T v

a .. I ,

..... 1 I .... --_J .....

Fig. 10. Qualitative model showing the mechanism of formation of cracks induced by thermal stresses.

200r-----~W~d------------------~

150

q; > Q) -100 ~

~ '-'

50

~ .. ---- ... _/

170 190 210 230 250 Pixel number

Fig. 11. Grey profile (continuous line) and its derivative (dot line) for a laser cut obtained by Digital Image Processing; 0 = black, 255 = white.

125

speed the cut surfaces appear covered by a layer of charred material. At the highest speed the

charred material is almost absent, while some matrix cracks inside the lamina perpendicular to

the kerf are present. Delamination and fibre cracks are not evident.

4.3. DAMAGE INDUCED BY THERMAL STRESSES WHEN CUTTING AFRPs

As mentioned above further damage may occur which is induced by thermal stresses

arising during laser cutting. This damage consists of transverse cracks in plies laying at large

angles with respect to the cutting direction. Taking into account that a single lamina of

kevlar/epoxy exhibits a high transverse coefficient of thermal expansion and a negative value in

the longitudinal direction, the mechanism of damage can be understood with the aid of Fig. 10,

where the temperature T and the normal stress inside 90° plies at two different points are plotted

vs time t. Assuming that the matrix undergoes deformation when the temperature exceeds its

glass transition temperature (T g), it can be seen that for point P a plastic deformation occurs for

T>Tg, i.e. the time interval between tl and t2' In the figure it is assumed that the maximum

temperature at the cutting surface is equal to the vaporisation temperature of the fibres Tv' As a

consequence of the aforementioned plastic deformation, the 90° plies will reach a free stress

condition during cooling at the time t3' when the temperature is still much higher than ambient.

After this time, the normal stress becomes positive producing transverse cracking, with an

average distance which decreases as the previous length reduction increases. The propagation of

cracks below the cutting surface stops at a distance where the maximum temperature reached

during laser interaction is lower than the T g of the matrix (below point P').

4.4. EVALUATION OF DAMAGE IN LASER CUTTING OF AFRPs

The extent of the damaged zone, as measured with an optical microscope, is reported in

Fig. 9 as a function of the cutting speed. The damage can be also evaluated by means of digital

image processing (DIP) [12,l3]. Two different methods of DIP can be used for evaluating the

damage: the first one is based on the evaluation of different threshold levels of grey, the second

one is based on the evaluation of the grey profile that permits one to measure the damage also

from an analytical point of view. This latter method is also less sensitive to background noise

and to local fluctuations of brightness, without limiting the efficiency of the analysis when it is

performed taking into account the derivative of the grey profile (Fig. 11). Therefore.

measurement of the damaged zone can be readily performed by measuring the Integrated Optical

Density (IOD) or by measuring the Mean Absolute Deviation (MAD) which gives similar results

126

[13). An increase of damage has been obselVed at the highest cutting speed which is related to

the following aspects: the increase of the amount of hot reflected gases, and the reduced

effectiveness of the assist gas flow in taking away the products of cut, which both lead to an

increased heating of the top surface of the laminate.

By using DIP, the measurements of the kerf width are reliable when compared with those

obtained from optical microscopy and exhibit the advantage of the absence of operator skill. Also

the Fast Fourier Transform (FFf) can be used to measure the kerf width [13] but with lower

reliability.

5. Technological models of cutting

5.1. MAXIMUM CUTIING SPEED

For a given power density and material thickness a limit feed rate exists, above which a

through cut cannot be obtained: this limit feed rate will hereafter be referred to as the maximum

cutting speed Vmax. It is evident from the above considerations that theoretically predicting the

maximum cutting speed would have twofold importance: V max determines the best conditions

for high productivity and good cut qUality. Furthermore, an explicit correlation between V max

and cutting parameters would permit the choice of the best laser system and the most economical

operating conditions to achieve the best results.

A model based on energy balance is generally adopted to evaluate V max [9);

unfortunately, it results in an analytical relation where the maximum cutting speed is calculated

as a function of the mass of evaporated material, which is unknown a priori. In Ref. [II] a one­

dimensional thermal model was proposed. The maximum cutting speed is given by V max=P/'¥

SD, where P is the laser power, D is the focal spot diameter S is the thickness of machined

material. and 'JI is the material parameter. For AFRPs 'JI is 3730 J/cm2: for GFRPs 'JI is 11,100

J/cm2; and for CFRPs 'JI is 40,000 J/cm2. The quantity VmaxD plotted as a function of PIS

results in a straight line passing through the origin, with slope l/'JI (Fig. 12). No distinction is

made between cuts carried out in the warp and weft direction because maximum cutting speed

was found to be coincident for both. The model is expected to work well for high power density

and feed rates. In fact under low interaction time heat conduction losses can be neglected, and the

cut process can be considered quasi-adiabatic.

127

150 acoco

125 ~ ~ ,--..

(f)

,,100 N

E 75 E

'--"

0 50

)( 25 0 E

> 0 0 250 500 7

P / S (W/mm)

Fig. 12. Maximum cutting speed at different machining parameters.

5.2. INFLUENCE OF FIBRE DIS1RIBUTION

We have developed a micromechanical formula (14) which is capable of predicting the

vaporization energy of composites as a function of vaporization energy of constituent materials

(fibre and matrix) and volume content. The formula, based on energy considerations, is formally

identical to the well known rule-of-mixture and it increases the usefulness of the previous model,

allowing for the calculation of the optimum cutting conditions, irrespective of the fibre volume

ratio of the composite under examination: S = agPN[(Hf - Hm) FrtHm)), where: S is

~2000'-(~I~I~I~O~~~,-=--~o.--~------------------'

E ~ F, = 1 .... E ~ F, = 2711 ~ F,= ......

,,1500 ~ "--"

1000 (f)

" 500 D...

25 50

V / 9 75 2 100

(mm Is) 125

Fig. l3. Influence of fibre content on the maximum cutting speed for GFRP.

128

the depth, a is the coefficient of laser beam absorption, g is a characteristic of the laser system, P

is the beam power, V is the cutting speed, Hf and Hm are the vaporization energy or fibre and

matrix respectively, Ffis the volume fibre fraction.

The experimental data, obtained on GFRP panels with different fibre contents (Fig.l3)

show an excellent agreement with theory. We have also used this theory to predict the response

to laser cutting of composite material with a fibre content varying along the thickness; it results

that, when the fibres are not uniformly distributed in the matrix, the interpretation of the

experimental results may be erroneous, thus causing errors in the evaluation of the material

thermal properties, or in the prediction of the kerf depth.

6. Conclusions

The quality of laser cuts depends on the interaction time between the beam and the

material and therefore on the translation speed of the beam. The quality of the cut, in terms of

the uniformity of the surface morphology and extension of the heat-affected zone, is better for

those composites which exhibit low differences between the thermal properties of the constituent

materials. The best results have so far been obtained for AFRP in which, because of the organic

nature of the fibres, the constituents have similar properties. GFRP and CFRP have shown

poorer results. Therefore, laser cutting seems to be particularly useful for materials such as

AFRP that normally show some difficulties in being machined by conventional tools. The

geometrical arrangement of the fibres, their weave, and their bonding to the matrix are less

important for the laser cutting process than for the mechanical process. The intensity profile of

the laser beam influences the cutting results. It should be observed that high-power laser systems

plus high-speed feed rates would give best productivity. However, in this case, the cost of the

laser system could be prohibitive compared with other cutting systems. Consequently the best

solution must be carefully evaluated for the specific ease.

References

1. Koenig,W., GraB, P., Wulf, Ch. and Willerscheid, H. (1985) "Machining of Reinforced Materials", Annals of the CIRP vol. 34 , p. 537.

2. Schwartz, M.M. (1983) "Composite Materials Handbook." McGraw-Hill, New York.

129

3. Koenig, W. and Justen, S. (1984) "Improvements in machining quality and tool life al drilling and countersinking of Kevlar composites" Int. Rep. Fraunhofer Institute fur Produktionstechnologie, Aachen. West Germany.

4. Caprino, G., Diterlizzi, A. and Tagliaferri, V. (1988) "Damage in Drilling Glass Fiber Reinforced Plastics" Advancing with Composites, Proceedings of Int. Conf. on Compo Mat. May 10-12, Milan, Italy, pp. 493-503.

5. Tagliaferri, V., Di Ilio, A. and Crivelli Visconti, I. (1985) "Laser Cutting of Fibre Reinforced Polyester" Composites vol. 16, p. 317-325.

6. Tagliaferri, V., Di Ilio, A. and Crivelli Visconti, I. (1985) " Cutting of Aramidic Fibres Composite with A CO? CW Laser" Proceedings of the International Conference "ECCM", Bordeaux.

7. Tagliaferri, V., Di Ilio, A. and Crivelli Visconti, I. (1987) "Machining of Fibre Reinforced Materials with Laser Beam: Cut Quality Evaluation" Proceedings of the International Conference "ECCM", London. ,

8. Tagliaferri, V., Di Ilio, A. and Crivelli Visconti, I. (1987) "Laser cutting of Aramid-FRP" Proceedings of the International Conference "Laser Advanced Material Processing", Osaka.

9. Duley, W. W. (1977) "C02 Laser Effects and Application", Academic Press, New York. 10. Di Ilio. A. and Tagliaferri, V. (1989) "Thermal Damage in Laser Cutting of (0/90h~ Aramid

Composites" Composites, vol. 20, pp. 115-119. 11. Caprino, G. and Tagliaferri, V. (1988) "Maximum Cutting Speed in Laser Cutting of Fibre

Reinforced Plastics" Int. Journal of Machine Tool and Manufacture, vol. 28, 4, p. 389-398. 12. Di Ilio, A., Tagliaferri, V. and Veniali, F. (1990) "Machining Parameters and Cut Quality in

Laser Cutting of Aramid Fibre Reinforced Plastics" Materials and Manufacturing' Processes, vol. 5,pp.591.

13. Di Ilio, A., Tagliaferri, V. and Veniali, F. (1990) "Digital Image Processing of Laser Induced Damage in Cutting of Aramid Composites" Proceedings of the International Conference "Laser Systems Application in Industry", Torino, Nov. 7-9.

14. Caprino, G. Covelli, L. and Tagliaferri, V. (1993) "The Importance of Material Structure in the Laser Cutting of Glass Fibre Reinforced Plastics Composites" Composite Manufacturing, (in press).

CHARACTERISTICS OF LASER HEATING OF MATERIALS

I.Yu. SMUROV Ecole Nationale d1ngenieurs de Saint-Etienne 58, rue Jean Parot 42023 Saint-Etienne Cedex 2, France

ABSlRACT. The present paper deals with how heat dynamics under laser action are influenced by: (a) temperature dependence of absorptivity and heat properties of metals; (b) surface oxidation and (c) reduction. It is shown that the joint influence of temperature dependences (or surface thermochemistry) can lead to quantitative and qualitative differences from the results obtained on the basis of linear models.

1. Introduction

From a general point of view, the heating of materials by concentrated energy flows, such

as those from lasers, plasmas, electron beams, concentrated solar energy etc, can be considered

as the action of some effective heat sources with different spatial-temporal parameters

q=q(x,y,z,t) [1-2]. The characteristics of these effective heat sources are determined from the

consideration of interaction of energy flows with materials on the microscopic level [3-4]. The

penetration ability of a heat source, i.e., the manner in which its characteristics vary with z

coordinate, depends on the properties of the target material (metals, dielectrics, semiconductors),

the nature of the energy flow, and its parameters (for example, laser wavelength). Laser action

and solar irradiation correspond to a so-called surface heat sourc~, because the absorption of

electromagnetic waves in metals occurs in a thin surface layer (for example, a few fractions of a

micrometer thick at a laser wavelength of about 1 Jim). This thickness is much smaller than the

typical spatial dimensions of the temperature field distribution. On the contrary, electron beam

heating, as a rule (to be more precise, if accelerating voltage is larger than 20 kV), is

131

S. Martellucci et al. (eds.), Laser Applicationsfor Mechanical Industry, 131-149. © 1993 Kluwer Academic Publishers.

132

characterised as a so-called volume heat source, because of the sufficiently larger propagation

depth of electrons in metals.

Calculations of temperature distributions in material bodies are often made on the

assumption that the effective heat source has ideal properties, which can lead to appreciable

deviations of calculated data from experimental results. For most practical cases, the form of the

effective heat source can be represented as a product of a time-dependent function and a function

of surface coordinates q=Aql (t)q2(x,y,z), where A is the absorptivity. Examples of typical

spatial distributions that are widely used in simulations are: Gaussian (normal) Q2=exp( _kr2) and

uniform distribution Q2=const over the focused hot spot.

In the case of laser heating, the function Q2(x,y) is consistent with the above mentioned

cases and, in general, to their superposition (for example, Q2(x,y) can have a minimum in the

centre of a heating spot). The function ql (t) for C02 laser action is usually considered to be a

constant, but ql (t) is considerably more complicated in the case of pulsed laser action. The most

typical pulse structure consists of a train of flashes or spikes of various powers and duration of

about 1 J.Is, which is characteristic of free-running, or normal pulse, operation. As a rule, a

characteristic pulse envelope can be observed.

One of the typical properties of a laser, as a heat source, is the strong dependence of the

absorptivity both on the state of the target surface (namely, the degree of mechanical, chemical

etc treatment), and on the surface temperature [4-6]. Variation in the chemical composition of the

surface layers during laser action also has an important influence on the energy input. Under

such conditions the parameters of an effective heat source start to vary as a function of the

surface temperature and some other initially unknown values (which depend upon the specific

initial conditions, such as the variation of the chemical composition of the surface), and this

makes the heat transfer problem essentially nonlinear on heat dynamics under:

The present chapter deals with the influence of the above mentioned factors on heat

dynamics under: (a) temperature dependence of absorptivity and heat properties of metals;

(b) surface oxidation; and, (c) surface reduction.

2. Temperature dependences of absorptivity and thermal properties of metals

One and two-dimensional problems in the laser heating of metals with partial allowance

for nonlinearities have been studied in a number of papers (see, for example, refs. [7-9]).

Simultaneous allowance has been made in the present approach for all nonlinearities (the

133

absorptivity, the thennal conductivity, and volume heat capacity) which are inherent in the

process of laser heating of metals. It has been shown that the combined effect of the

nonlinearities leads, in a number of cases, to substantially new results.

Let us examine the heating of a semi infinite body by a moving laser beam. In the

coordinate system (x, y, z) positioned at the centre of the heating spot of a surface heat source

and moving with a constant velocity V in the positive direction of the xO axis (the system xO, YO,

zo is attached to the body), the mathematical fonnulation of the problem is as follow:

Cv(T) dT _ V aT = div(A.(T) grad T) at ax

(1)

aTI -1..1-az z=O

T(x = 00, y,z, t) = T(x, y = 00, t) = T(x, Y ,z, t = 0) = TO

Here T(x, y, z, t) is the temperature; t is the time; x, y, z are the spatial coordinates; V is the

velocity displacement of the beam; qO is the energy density flux; k is the degree of concentration;

A(T) is the absorptivity; A.(T) is the thermal conductivity; Cv(T)=p (T) cp (T) is the volume heat

capacity; TO is the initial temperature.

For a sizable class of metals and alloys the temperature dependences of the thennal

conductivity and volume heat capacity are approximated well by third-degree polynomials

[12,13] (see Fig.l). An approximately linear dependence of the absorptivity on temperature

follows from theoretical concepts (see, for example, ref.[1O]) and from an analysis of the

experimental data (see, for example, ref. [11,13]).

For molybdenum:

A.(T) =173.8-9.20 .1O-2 T +4.29.10-5 T2 -7.59.10-9 T3, Wlm K

cp(T)=216.7 + 0.103 T - 6.80 .10-5 T2 + 2.01 .10-8 T3, J/kg K

p(T)=1.02 ·104 - 3.8 .10-2 T, kg/m3

A(T) =0.99 .10-4 T

For tungsten:

A.(T) =196.4 - 0.135 T +4.63.10-5 T2 -3.62 .10-10 T3, Wlm K

cp(T)=130 + 1.36 .10-2 T + 4.04 .10-6 T2, J/kg K

peT) =1.93 ·104 - 3.3 .10-2 T, kg/m3

A(T) =2.4 .10-2 + 1.03 .10-4 T

134

Fig.I. Temperature dependences of the absorptivity (curves 1), the thennal conductivity (2) and the volume heat capacity (3) for tungsten (solid curves) and molybdenum (dashed curves).

It should be noted that sizable changes in the reduced thennal conductivity

Ared(T)= AoA(T) I AOA(T) over the temperature interval examined can serve as a criterion that

one must calculate the heating of a given metal using the nonlinear model. For example, the

value of Ared(T) for molybdenum decreases by a factor of 20 with increasing temperature

(Fig.2). One should note that the relations obtained consist of combinations of elementary

functions. They take an especially simple fonn in the calculation of the temperature at the centre

of the heating spot of a fixed heat source (V=O) for 't=kat« 1 [14, 16-18].

The results of a calculation of the transient temperatures of molybdenum and tungsten

during heating by a fixed laser beam are presented in Fig. 3. The calculations showed that, for

the initial stage of heating, the temperature values calculated in the nonlinear approximation

Ared (T)

I,D

°:100 1000 ZOOO :1000 T, II.

Fig.2. Temperature dependences of the reduced coefficients of thennal conductivity Ared (T).

135

agree with the temperature values calculated with the initial values of the absorptivity and

thermophysical coefficients, as follows from general considerations. In the same initial stage of

heating, the temperature values calculated in the linear approximation and for mean-integral

values of the parameters are much larger than the temperature values found from the solution of

the nonlinear model. This is connected with the fact that in this linear approximation the

absorptivity is greater (and the thermal conductivity is smaller) than its true value at the initial

temperature. The situation is reversed as the temperature increases, and, consequently, at a

certain moment of time the temperatures being compared coincide (Fig. 3b), but this may occur

beyond the limits of the temperature range examined (Fig. 3a).

The calculations that have been carried out showed that during the laser heating of

molybdenum and tungsten, one cannot select constant values of the absorptivity and of the

thermophysical coefficients such that the linear model describes the heating process in an

acceptable way over the entire temperature range from the initial (room) temperature up to the

melting point temperature, since the discrepancies will be large either at low or at high

temperatures.

It should be noted that whereas in any linear problem the temperature at the centre of the

heating spot is described by the time dependence arctg(2(kat) 1/2), the graph of which does not

have inflection points, the solution of the nonlinear problem has a significantly different time

dependence and can have inflection points. From a physical point of view this circumstance is

connected with the sharp changes in the rate of heating that are caused by significant changes in

. -3 1jo,o,o,t}'10 ,K

Mo

2

1

T(o,o,o,t}·'0~14. w

q

7,0 7,'1 f; '10~ S

Fig.3. Calculation of T(r=O, z=O, t) during the heating of molybdenum (Fig.a) and tungsten (fig.b); %= 2.105 W/cm2; k=1O (a), 102 cm-2 (b); curves 1 - results of calculation taking into account temperature dependences of absorptivity and heat transfer properties (I-approximate analytical solution, 2-numerical solution); curves 3 and 4 - results of calculation using the linear model for the mean integral and initial values. respectively. of the absorptivity and the thermophysical properties.

136

the absorptivity and thennophysical properties of the metal. The local maximum on the heating

rate curve for tungsten (FigA) is related to the local minimum on the curve of its thennal

conductivity. It may be noted that the nonlinear model has a narrower range of parameters where

the heat transfer problem can be considered one-dimensional [17,18]. This is related to the

question discussed below.

The increase of the degree of concentration in the temperature distribution on the surface

of the body being heated is another interesting feature that is inherent to nonlinear heating

problems. The distribution of temperature, nonnalized to its value at the point r=O, during the

heating of molybdenum is presented in Fig. 5. The curves 1 and 2 are, respectively, the solution

for the nonlinear model and the corresponding linear model at time t= 15 ms. Evidently, when the

nonlinearity is taken into account a sharpening of the temperature distribution on the surface of

the body occurs, whereas the solution of the problem in the linear approximation always shows a

decrease of the degree of concentration of the temperature distribution as a result of thennal

conductivity. The indicated effect is associated with an increase of the absorptivity with a

simultaneous decrease of the thennal conductivity.

The simulation of laser heating of Cu and chromium-nickel steels has also been carried

out. The heat transfer coefficient of Cu decreases with increasing temperature, which is why the

temperature dependences of A(T) and A.(T) in the boundary condition on the irradiated surface

are supplementary, and intensify one another. The above mentioned nonlinear effects take place

in laser heating of Cu, but their influence is low compared with W and Mo. This is the result of

0,001 o,oos

Fig.~. Heating rates in the middle of the heating sRot (r=O); tu¥sten - QO=2 '106 W/cm2, k=102 ~m-2 (soM curves), and molybdenum - qo=2 '105 W/cm2, k=lO cm- (dashed curves); curves 1 - nonhnear model; 2,3 - linear model.

137

a slower rise of the absorptivity with temperature, but mainly of the narrower temperature

interval under consideration (for Cu, Tm=1256 K).

Simulation of chromium-nickel steel heating shows a comparatively small difference

between the results that are obtained from the nonlinear model and from the linear one (for lhe

mean integral values of lhe absorptivity and lhe lhennophysical properties). As a rule, this

difference is of about 20%, wilh lhe exception of lhe initial stage of heating, where lhe ratio of

lhe temperatures is detennined mainly by lhe ratio of lhe initial and lhe mean integral values of

absorptivity. The relatively weak influence of nonlinearities is explained by lhe fact lhat bolh

absorptivity and lhennal conductivity are increasing wilh temperature. That is why lhe

nonlinearities in lhe boundary condition on lhe irradiated surface are tending to compensate one

another.

As a general conclusion, it is possible to note lhat lhe sharp temperature dependences of

absorptivity and heat transfer properties cannot by lhemselves guarantee the appearance of lhe

above mentioned nonlinear regularities, lhat distinguish the linear and nonlinear models of laser

heating. The interaction of these temperature dependencies is also quite important, because lhey

can either intensify or compensate one another. As is shown by lhe above mentioned examples,

more important are the dependences of the absorptivity and the thennal conductivity (this can

easily be detennined from a general point of view); but it is important to underline that lheir

individual temperature dependencies A(T), A.(T) are not as significant as their assemblage

temperature dependence A.red (T)=AQA.(T)/AoA(T) (this follows directly from the boundary

condition on the irradiated surface). The sharp variation of A.red (T) with temperature is lhe

indication of a significant influence of the nonlinear effects (similar to those mentioned above)

on laser heating dynamics.

Fig.5. Temperature distribution on a molybdenum surface at t=15 ms (qo==2 .105 W/cm2, k=lO cm-2 ); curve 1 - nonlinear model; 2 - linear model; 3 - Gaussian distribution of the incident laser radiation.

138

3. Laser heating of a metal in an oxidizing atmosphere

The interaction of laser radiation with metals in an oxidizing atmosphere depends

significantly on the highly absorbing oxide film which is formed on the surface of the metal [19-

22]. As shown by experiments, oxide layers can appear on the target surface even when they

interact with rather short pulses- 10-6 s. Pulses of - 1 ms duration promote the growth of oxide

films 0.01 to 0.1 ~ thick, and CW laser beam can encourage oxidation to a depth of tens of

microns. The isothermal surface oxidation of metals proceeds in several stages: surface

adsorption of oxygen; binding of the free electrons of the metal or of the oxide with oxygen;

diffusion and migration of metal and oxygen ions through vacancies and interstitial sites of the

reaction products to phase boundaries; and the chemical reaction proper that yields a new oxide

layer. The metal oxidation rate depends on the temperature, oxide thickness, oxygen pressure and

other factors. The kinetics of the isothermal oxidation of metals can be described by the equation

(2)

where h is the oxide thickness, Bm a constant, T a the temperature of activation at the slowest

stage of the reaction, which commonly ranges from 1 to 5·104 K; m and n - the parameters.

Most of the known relations that define the oxidation process can be expressed in the form

of Eq.2. When m=n=O, the equation describes linear oxidation, for which the reaction rate is

independent of the oxide thickness. When m=O and n=-2, equation is linear too, but here it is

thermionic emission from the metal rather than the emission of ions from the metal to the oxide

that limits the reaction rate. The case m=l corresponds to the Mott formula if n=l and to the

Wagner formula if n=O. These formulas conform to the parabolic relations of oxidation

controlled by diffusion or ion migration in the oxide. If m=2, a cubic relation for isothermal

oxidation follows from Eq.2. Many characteristic parameters of the heating process as w~ll as

approximate relations and the results of computer simulations of separate stages of the process

are presented in refs. [23-27].

Let us consider the nonlinear transient spatial problem of heating of a semi-infinite body

by C02 laser radiation, taking into account the growth of the oxide film as well as the

concomitant change in the absorptivity. The flux density of the incident radiation is assumed to

be high enough (qO> 105 W/cm2) that the heat losses from the surface can be neglected. The

139

chemical reaction heat, as well as the complex structure of the oxide layer being grown, and

temperature gradients in the oxide film are neglected.

The mathematical formulation of the problem has the following form:

(3)

-A ~I = qO A(h) exp( _kr2) aZ z=o

T(r = co,z,t) = T(r,z=oo, t) = T (r,z,t=O) = TO

Here T(x, y, z, t) is the temperature; t is the time; r and Z are the spatial coordinates; qO is the

energy density; k is the degree of concentration; A is the thermal conductivity; a is the

coefficient of thermal diffusivity; h(r,t) is the thickness of the oxide film; Ao is the absorptivity

of the metal in the absence of the oxide film on the surface; TO is the initial temperature.

According to Eq.3, the thickness of the oxide film varies parabolically, where d is the constant in

the parabolic oxidation law and Td is the activation energy of the chemical reaction (in degrees).

This law is valid for the growth of "thick" (h > 0.1 J.I.Ill) films. The estimates show that the initial

period of growth of the oxide film does not significantly affect the dynamics of laser heating,

since at this stage the absorptivity of the system oxide fIlm+metal remains virtually constant and

the heating can be described with the help of a linear model, i.e., ignoring the surface

thermochemistry. The absorptivity of a two layer (metal-oxide) system is

2X A(h)= 1- (l-Ao)exp(-cx.h) - -- sin ~h exp(-2cx.h) (4)

11

where a=(47t1AtiJ) 1m (£)1/2, ~(47t1A.m) Re(£)1/2, (£)1/2= 11 + iX, 11 is the real part, X is the

imaginary part, E the oxide permittivity, Am the laser wavelength, 1m the sign of the imaginary

part of the number, and Re the sign of the real part.

From the general relation that describes the dependence of the absorptivity on the

thickness of the oxide film, it follows that in the limit of thin films (~h2 «1, this corresponds to

140

A(h),%

o 1 1 ;r 'I 5 fj h, 11m

Fig.6. Dependence of the absorptivity of the oxide+metal system on the thickness of the oxide film; curve 1 - the exact value of the function A(h), 2 - quadratic approximation.

~~K ________________________________ ~

Fig.7. The temperature at the centre of the heating spot as a function of time for different values of the

incident radiation flux density: '10= loS (curve 1); 2 .105 (2); 3 .105 (3); 4.5 ·loS (4)W/cm2.

h,l1m

J

2

oL-~~~~--~----~ 5 t, TO'" S

Fig.8. The thickness of the oxide film at the centre of the heating spot as a function of time for different

values of the incident radiation flux density: '10= 5.105 (curve 1); 4.105 (2); 3 .105 (3) W/cm2.

141

h« 0.5 ~), the dependence has the fonn A(h)=AQ (l+~h2), where J3= 41t2 (112 -1)/A,2m. This

case is examined in [25), where an approximate analytic solution of the spatial laser heating

problem is presented. These assumptions limit the region of applicability of the relations

obtained to small thicknesses of the oxide film and, therefore, small changes in the absorptivity

and small increases in the transient temperature due to the chemical reaction.

In the present chapter, instead of the exact relation, an approximation of the dependence

A=A(h) for oxide film thickness up to 3 j.Ull is used, which makes it possible to perfonn

calculations for a wide range of laser-heating conditions (Fig.6). This approximation is in fact

present in Eq.2, where B is the coefficient of the parabolic approximation. Estimates show [27]

that in the range of temperatures (TO<T<Tm, Tm is the melting point temperature) and radiation

flux densities studied, there is not enough time for the thickness of the oxide film to grow

appreciably (it does not exceed 3 j.Ull).

From the relations obtained follow a number of interesting characteristics, which are

inherent to the spatial laser-heating problem in a nonlinear fonnulation. Heating of the metal by

continuous C02 laser irradiation in an oxidizing atmosphere stimulates the development of a

layer of oxide film on the surface of the metal, whose rate of growth depends markedly on the

value of qO ; in addition, the absorptivity of the oxide+metal system that appears increases

markedly and as a result the rate of heating increases, which, in its turn, causes the thickness of

the oxide film to grow, etc. It is evident from Fig. 7 that the presence of an oxidation reaction on

the surface leads to the existence of characteristic heating stages, depending on the magnitude of

the incident radiation flux [18, 28-29). In the case of relatively "small" values of qO (for Cu,

A.m=1O.6 ~, qo< 105 W/cm2), absorptivity hardly increases at all and the transient temperature

at the centre of the heating spot is described by the well-known dependence arctg(2(kat)l/2)

(Fig.7, curve 1). As qO increases, the temperature increases nearly linearly from the stationary

value corresponding to the solution of the problem ignoring thennochemical reactions on the

surface of the metal. The rate of growth of the temperature increases with qO (curve 2). With

further increases in the flux density, beginning at a certain time, there is a sharp growth in the

temperature, which corresponds to a considerable increase in the absorptivity due to the growth

of the oxide film (curves 3,4).

Another interesting feature, following from the solution of this problem, is the

dependence of the thickness of the oxide film on radius and time. Fig. 8 shows the thickness of

the oxide film as a function of time at the centre of the heating spot for different values of qO.

Evidently for t<tst the thickness of the oxide film grows rapidly; in addition, the rate of growth

of the oxide film increases substantially with qO, which is caused by the more intense flow of the

oxidation reaction on the surface of the metal (here tst is the characteristic time for establishing a

142

stationary temperature in the absence of a thennochemical reaction). For 1>tst, which

corresponds to lower flux densities, the time dependence of the thickness of the oxide film is

weaker. This is explained by the limitation of the rate of growth of the oxide by the film layer

that has already appeared. Fig.9 shows graphs of the thickness of the oxide film versus the

distance from the centre of the heating spot, nonnalized to the value at the point r=O, for different

times. For comparison, the Gaussian distribution of the flux density of the incident radiation is

shown on the same graph (curve 3). It is easy to see that for short heating times the spatial

dependence of the thickness of the oxide film and, therefore, of the absorptivity also becomes

much sharper (curve 1), which is the result of the more rapid growth of the oxide film at the

initial stage of heating, but later (with increasing time) this sharpening will be less distinct, Le.,

the spatial distribution of the thickness of the oxide film is "smeared" and "smoothed" (curve 2),

which happens due to heat conduction. The analogous distribution of the transient temperature

on the surface of the body has a complicated character: remaining at all times a function that is

more concentrated than would be a spatial temperature distribution that ignores the thenno­

chemical reaction on the surface, at the same time it varies compared with the fixed flux-density

distribution. The specific spatial distribution calculated (the irradiated material is copper) will at

first coincide with the fixed flux-density distribution, then become sharper and, ultimately,

become smeared due to the inherently high thennal diffusivity of copper.

The present analysis of the problem pennits one to describe qualitatively the characteristic

regimes of heating of metals in an oxidizing atmosphere, previously observed experimentally."

h(r,t)/h(O,t)

(k) 1/2 r

Fig.9. Normalized distribution of the thickness of the oxide film on the surface of the metal at different times: kat=O.6 (curve 1), 10 (2); curve 3 - distribution of the flux density of the incident radiation.

143

4. Surface reduction of metals

The processes of laser induced reduction of metals, primarily refractory metals, from

oxides is another typical example of laser thermochemistry. The reducing mediums that can be

used here include gases such as hydrogen and carbon monoxide, liquids such as alcohol and

sometimes solids. The main feature of these heat processes is a degenerative feedback resulting

from changes in the oxide surface absorptivity with the growth of the reduced metal layer. A

number of works have treated the process of CO2 laser reduction of metal ftlm from an oxide,

for example, see the review of Ref. [251.

Let us consider the following model of the process. The surface of a semi infinite uniform

oxide body, immersed in a reducing atmosphere (H2), is heated by laser irradiation with a

Gaussian distribution of energy density flux. As a result of thermochemical effects a metal film

appears on the surface; its thickness increases in time obeying the general law

(5)

where B is a constant, T* the activation temperature of reduction, in K; h the thickness of m

the reduced layer. For a thin metal film n=O (linear law); as the film thickness increases, the

process is limited by diffusion, n=1 (parabolic law). In general, the assumptions of the model are

practically the same as for the oxidation process.

1 aT 1 a aT aZT --=--r-+­a at r ar ar az2

dh - = v dt

T. exp (--)

T

T (r=oo,z,t) = T (r,z =oo,t) = T (r,z,t =0) = To

(6)

144

The general method for determining the absorptivity of the two-layer system is the same as in the

previous section. The absorptivity of an oxide-metal film system for small film thickness is

given by the simplified expression

where A* is the absorptivity of a pure oxide, n is the refractive index, and AO is the metal

absorptivity. The analysis shows (30) a rapid decrease in absorptivity of the two-layer system as

the metal film thickness increases (Fig. 10). For a metal film thickness of about 0.01 Jll11 the

absorptivity of the system is close to that of the pure metal. That is why it is possible to assume

A(h)=AO = const, and consider the heat model of a massive body with the thermal properties of

oxide and with the absorptivity of metal. The same approach is not valid in the case of laser

oxidation, because of the sufficiently smoother dependence of absorptivity of the system "oxide

film on the surface of massive metal body" as a function of oxide thickness. The growth of oxide

films up to 0.1 Jll11leads to an absorptivity increase of about 5%. In the centre of the heating spot

the temperature growth is determined by

T(O,O,t)=T 0 + T *arctg(2(kat) 1/2),

where A., a - correspond to oxide, AO to the metal.

0,6

0,2

o 8 h, to-2).tm

Fig.lO. Absorptivity of a two-layer "metal film on oxide surface" system as a function of metal film thickness.

145

Analysis of transient temperature growth reveals the existence of typical regimes of laser

reduction that specify the structure and morphology of surface layers. For relatively small <In' the

maximum surface temperature is smaller than the melting point temperature of pure metal (for

Cu, T =1356 K), and the reduction process occurs in the solid phase (Fig. 11, curves 3,4). As m

<In increases, a melted metal film appears on the surface of a solid oxide body (for CU20,

Tm =1506 K), and the reduction process continues in the liquid phase. With further <In increases,

the melting of the oxide surface begins, and the system "liquid metal+liquid oxide" appears on

the irradiated surface. Further increase in CJo leads to an increase in oxide temperature (curve 1 in

Fig. 11) up to the decomposition point (for Cu20, Tdec:::;2073 K) and, as a result void fonnation

under the metal film, which curves the surface in the zone of laser irradiation (Fig. 12).

The processes of water evaporation on the reduction front under the metal film also

influence void fonnation. A further increase in oxide temperature leads to mass removal from the

zone of action, and to crater fonnation. The above mentioned stages of the process are illustrated

in Fig. 12 in the range of relatively high energy density values. The continuous thin metal film

(top layer in Fig.I2b) can cover practically the entire zone of laser action; evaporation can

destroy the continuous layer of metal and fonn a "volcanic" relief (Fig.12a); a droplet of liquid

metal fonns at the bottom of the crater on oxide surface (the crater diameter is about that of the

heating spot).

We should mention that there are optimum parameters of laser action for obtaining

comparatively thick layers of metal film. For relatively low values of <In the temperature of the

T(f:::;O, z:::;O,t), 103 K

21

f .

o t, 10-3 s

Fig.lI. Temperature at the centre of a heating spot as a function of time under different regimes: curve 1 - QO=4.8 .105; 2 - 3.2 ·105; 3 - 1.6.105; 4 - 0.8 .105 W/cm2.

146

surface and the rate of metal film growth are low (Fig.13, cUlve 2). On the contrary, for the high

values of 'la' destruction processes dominate. The optimum conditions are situated between the

above mentioned asymptotes.

2mm 0.2 mm

a

1 mm

b

1 mm

Fig. 12. Microstructure of a copper film recovered from a CU20 surface in hydrogen. (a) at a Jlressure of 2 ·loS fa; incident energy flow density 5 ·104 W/cm2; (b) 5.105 Pa; 105 W/cm2; (c) 5.105 Pa; 2 .106 W/cm.

147

h(D,t), J.lm

20

/0

o t, to-3 s

Fig. 13. Metal film thickness as a function of time at different laser fluences. curve 1 - Q0=4.8·105; 2 -3.2'105 W/cm2.

5. Conclusions

One of the pricipal characteristic properties of a laser as a heat source is the strong

dependence of the absorptivity both on the state of the target surface and on the surface

temperature. The variation of the energy input versus the surface temperature and the chemical

composition of the surface layers makes the heat transfer problem significantly nonlinear. In such

cases even simple laser heating exhibits some characteristics that cannot be obtained on the basis

of well known (and widely used) linear models. Among these characteristics are: the inflection

points on the CUlVe of the temperature in the middle of the heating spot; the extremes of the

heating rate CUlVes; and the increase in the degree of concentration of the surface temperature

distribution. It can be shown that, in general, one cannot select constant values of the

absorptivity and of the thermophysical coefficients such that the linear model describes the

heating process in an acceptable way over the entire temperature range from the initial (room)

temperature up to the melting point temperature, since the discrepancies will be large either at

low or at high temperatures. It should be noted that sizable changes of the reduced thermal

conductivity Ared (T)= AoA(T)/A.oA(T) over the temperature intelVal examined can selVe as a

criterion for the fact that one must calculate the heating of a given metal using the nonlinear

model. In addition, surface thermochemistry can lead to positive (for example, oxidation) or

negative (for example, reduction) feedback between the temperature and absorptivity.

148

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LASER ASSISTED MACHINING

I.Yu. SMUROV Ecole Nationale d'[ngenieurs de Saint-Etienne 58, rue Jean Parot 42023 Saint-Etienne Cedex 2, France

L. V. OKOROKOV Institute of Metallurgy, Russian Academy of Sciences Leninsky prospect, 49 Moscow, Russia

ABS1RACT. Experimental results are presented concerning the characteristic features of mechanical treatment of a wide range of materials (Mo, W, Ti, ceramic materials, steels) preheated by laser radiation. Various physical phenomena which accompany laser assisted machining (LAM) are discussed: plasticization, structural and phase changes of materials, variation of friction conditions with temperature, effects of the laser plasma, etc. It is shown that under optimum conditions LAM significantly increases the quality of machining: roughness of the treated surface, depth and rate of cold working decrease; and shears and microcracks disappear. LAM permits: interrupted cutting of materials with increased strength characteristics; turning of soft and/or thin-walled materials; finishing of materials for which structural and phase changes beneath the cut line are not desirable. Generally, the productivity of machining can be increased up to 4-6 times.

1. Introduction

One of the methods for improving the machinability of metals is to preheat the workpiece

by a suitable source of energy. Resistance heating and plasma flow action are the most widely

used techniques of heat pretreatment; these effectively reduce the cutting force and increase the

cutting speed [1,2]. In the process of resistance heating, the heat source raises the temperature of

the material ahead of the cutting tool and does not essentially affect the surface adjacent to the

transient surface. The plasma arc can heat up the layer being cut throughout its thickness before

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s. Martellucci et al. (eds.), Laser Applications/or Mechanical Industry, 151-163. © 1993 Kluwer Academic Publishers.

152

the tool arrives at the area of chip formation. On the other hand, the large heat-affected wne

leads to thermal deformation of the workpiece and to changes in the structure of the work

surface. The optimum application for plasma assisted machining is, for example, the preliminary

rough turning of hardly machinable alloys. The typical features of plasma flow, as a heat source

(taking into account its possible nonstability), do not permit using it for finish machining,

because of poor quality of the treated surface.

Evident advantages of the laser for material preheating are the opportunity to focus the

beam to a small spot, and the availability of a high energy density flux. It is possible to note

other promising features of laser, such as opportunity to vary and to control spatial and energy

parameters on the irradiated surface, high stability of the process, general use for a wide class of

materials. radiation supply to places where access is difficult. possibilities for proces automation.

etc. On the other hand. it is necessary to emphasize from the beginning that to use laser assisted

machining (LAM) for ordinary materials. as for example cheap steels. is not optimum because of

the high prices. even for today. of laser equipment.

One of the requirements for mechanical treatment with preliminary heating is a maximum

decrease in the cutting forces. while the temperature of the tool-workpiece interface is kept at

some optimum level.

The contact temperature 9c arises from an additional heating temperature 9R. and the

temperature 9d conditioned by the processes of deformation and friction in the zone of chip

formation. It is necessary to emphasize that 9d. in general. is a function of 9R.

The focused laser beam heats the material being treated in the wne of chip formation up to

a temperature higher than 1000 K. considerably decreasing cutting forces and the value 9d; at the

same time. the heating temperature near the contact surface is about 400-800 K. Thus we have

conditions for increasing tool lifetime at a given efficiency. or for increasing the cutting speed. or

for improving the quality of machining.

2. Methods and materials

The use of additional laser heating is effective while performing various kinds of

mechanical treatment: external longitudinal and face-plane turning. hole boring. drilling. thread

cutting. rolling. milling. and planing.

A schematic arrangement for LAM (turning) used in the experiments is shown in Fig. I.

C02 (5 kW) and Nd:YAG (400 W) lasers were used. The device is mounted on the base of the

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turning lathe. The laser beam is focused at distance I (in typical conditions approximately 0.5

mm) from the cutting edge.

The following elementary estimation can roughly determine the optimum distance t. The

time for the workpiece to pass a distance I between the heating spot and the tool is IN, where

y= cutting speed. On the other hand, during the same time heat must propagate from the surface

to the cutting depth z, so z=(at) 1/2, a= thermal diffusivity (corresponding to the optimum

conditions). This gives I ",yz2/a; the experiments, in general, confirm this functional

dependence. When I has a lesser value, the material has insufficient time for being heated up,

when t is a greater value, the material gets cold before coming to the tool (Fig.2).

The design of the optical system makes it possible to stabilize the focal spot position

relative to the tool, while its supports move at random, and, using the piezodrive of the mirror

(4), to scan the focal spot on the cutting surface. Laser beam alignment on the workpiece and

adjustment of the optical system is carried out by a visible A = 0.63 ~ He-Ne laser, the optical

axis of which is made coincident with that of infrarcd laser by mirrors (2), (3). Focusing is

carried out by means of KCl lenses, as well as metal optics: a spherical mirror, arrangements of

flat and spherical mirrors, and an off-axis telescope with compensated aberration.

In most cases the focal spot has an elliptical shape on the workpiece. In such cases the

shape of the temperature field is determined by the ellipse orientation relative to the direction of

movement. When the major ellipse axis is parallel to the movement direction, a narrow and deep

y- ..... 1 He - Ne I I 7 ,

Fig.1. Schematic diagram of a LAM set-up: (1), C02 laser; (2)-(6) optical system: mirrors, lens; (7), He­Ne laser; (8), cutting tool; (9), plasma-generator; (10), lathe.

154

111l1l

JIJIJ L-._L-.........L._--'-_-'----':--'

!l -415 -451J -,1,75 -1.171J J,MM

Fig.2. The dependence of the temperature of additional heating of Mo by 5 kW C02 laser irradiation on the distance from the centre of heating spot (in the direction of movement) on the various depths z (V=60 m/min). Calculated values: curve 1 - z=O.2 mm; 2- 0.1; 3 - 0.05; 4 - 0;- experimental results.

zone of heating is fonned. If the major ellipse axis is nonnal to the movement direction, a wide

heating zone of a smaller depth is fonned. Experiments and calculations show that the maximum

temperature is about four times higher in the first case than in the second (for typical

conditions).

The choice of either orientation is determined by the heating scheme and the dimensions

of the layer being cut. Thus, when heating from the side of the surface being treated and having a

high cutting speed, it is preferable to choose the first variant, and when heating from the cutting

side of the surface with deep cutting and lower cutting speed, the second variant is preferable. If

the cutting ann is much larger than the focal spot, the character of cutter wear depends on the

focal spot position on the cutting ann. To obtain maximum cutter lifetime, it is necessary to

provide heating along the entire cutting ann. In our experiments this was achieved by scanning

of the focal spot. The following cutting edge geometry was used: tool clearance a= 12° (8°), tool

rake y=15° (10°), tool cutting edge ~=45°, cutting edge radius r=0.2 mm (0.5 mm). For

zirconium hydride we used cutting tools with W-CO tips: a=100, y=20°, ~=90°, ~1=30°, r=0.2

mm. The workpieces were cylinders with length 50-150 mm and diameter 12-20 mm of various

materials: Mo alloy with 3% Nb (monocrystals), W (obtained by the precipitation from gas

phase), Ti, zirconium hydride, ceramic materials, and various steels.

155

One of the main problems in LAM is the low efficiency of laser heating (keeping in mind

its high price), caused by the low absorption of metals at the C02 laser wavelength A. = 10.6 l.l

m. There are several ways to address this problem: the use of absorbing coatings, a shorter

wavelength laser (for example, A. = 1.06 1J.Ill), preheating or oxidation (to increase the

absorptivity), and the use of radiation reflected from the workpiece.

From a general point of view, to effectively use absorbing coatings in LAM, it is necessary

to fulfil the following requirements: 1) Applying a coating, and its removal during the LAM

process; 2) Fine dispersion of coating particles, to provide a steady flow of the particle stream; 3)

Localization of coating deposition (on the cutting arm); 4) Rapid adhesion to the material being

treated, as radiation is applied after't = (lO-C 10-2) s; 5) Good thermal and beam stability, as

well as minimum evaporation (evaporation leads to screening of the laser beam by the vapor); 6)

High absorption at the wavelength A. = 10.6 1J.Ill; 7) Inertness with respect to the materials being

treated; and, 8) Minimum abrasive effect on the cutter.

In our experiments the best results were obtained with a K2Si03 coating, deposited on the

cutting arm by the stream method during the treatment process. Fulfilling the many requirements

mentioned above is rather difficult, which is why other methods of increasing the efficiency of

laser heating were used in our further work. Concerning the use of radiation at wavelength A.

=1.061J.Ill: at this wavelength, the absorption is about 40% for such metals as W, Mo, Ti.

Another approch mentioned above for increasing the efficency of LAM is to use radiation

reflected from the workpiece. This radiation is concentrated in the zone of heating by a spherical

mirror mounted on the cutter. As preliminary experiments showed, in LAM of tungsten,

molybdenum, titanium, 60-65% of the power is absorbed and it is possible to replace the

powerful C02 laser with a smaller laser.

For preheating and oxidation, a small plasma generator «9), in Fig.1.) of 1 kW power,

with the plasma arc diameter about 2 mm, was used. The spot of laser heating served as the

anode spot of the plasma arc; this increased the laser radiation absorption due to oxidation and

heating of the workpiece; in addition, this provided greater localization of plasma heating and

stability of the arc position relative to the cutter. In our experiments we were not able to bring the

arc closer to the cutter nearer than 4 mm, where the peripheral part of the heated gas column

reached the cutter. Cutters with welded plates from W-Co, W-Co + boron nitride, hexanite were

used in the experiments. The main parameters of LAM were controlled during the process.

Temperature in the heating spot centre was measured by a pyrometer with high spatial resolution,

temperature of the cutter-workpiece contact by the method of the natural thermocouple, forces of

cutting by a dynamometer, wear rate of the cutter by the method of acoustic emission (AE) using

an acoustic signal meter, and the oscillation amplitude by the device Balatron 2003.

156

After LAM, the roughness of the treated surface, depth and rate of cold hardening,

presence of shears, microcracks, structural and phase transformations on the treated surface were

observed. Comparison of the results of metallographic analysis with characteristics of the AE

signal showed that control and retention of an optimum level of treatment quality by change of

heating parameters can be successfully performed by AE methods.

Our experiments were made using an ordinary and comparatively low quality mechanical

lathe; therefore it is possible to expect considerable improvement in the machining quality if

another, more precise lathe is used. The similar is the situation with the laser equipment. So,

more important are not the exact values of the parameters obtained that were improved. Thus, are

less important than the general tendencies of the main physical phenomena accompanying LAM.

3. Physical phenomena accompanying LAM

LAM is accompanied by the following physical phenomena: heating and cooling of the

material, structural and phase transformation, plastic deformation, destruction, friction,

generation of stress waves, electron emission, plasma formation over the material surface, etc.

All these phenomena are interrelated and determine the efficiency and treatment quality of LAM.

On the basis of our experiments and literature data one can single out a number of phenomena

which most influence LAM of a given material [3-11].

3.1. TEMPERATURE WEAKENING OF THE MATERIALS

Decrease of the strength characteristics of the materials being treated is one of the main

phenomena defining improvement of the material workability, especially in the case of very

strong, structurally stable metals, e.g., for example, tungsten. molybdenum. Change of metal and

alloy strength versus temperature is characterized by a relative ultimate resistance

s=ap(6)lap(6), where ap(6) and ap(60) are the ultimate resistance at temperature 6·and

room temperature 60' respectively.

It is possible to use the following expression: s = exp (-k(6 I 6m)n), where: k, n are

numerical coefficients; em - is the temperature of melting. Our estimations for LAM of

tungsten and molybdenum show that with a laser of 2.5 kW (absorptivity of about 10%), s is

decreased by a factor of two at a depth z= 0.1 mm.

157

Therefore, we may to expect that the maximum effect can be obtained in the processes of

finishing or semi-finishing. For example, in the process of final turning of W and Mo, the

component of the cutting forces Pz has been decreased by 45-56%, and treatment efficiency has

been increased 6.5 and 3.5 times, respectively. The results of a typical experiment (with

conditions not corresponding to the optimum values) are presented in Fig.3.

Temperature dependence of the shear stress 'to for Inconel 718 alloy is different from the

above mentioned formula; 'to remains practically constant up to the temperature of 650°C.

Above 650 °C, a sharp decrease of the shear stress takes place. LAM of the alloy Inconel 718

showed that Pz decreases by 40%, and turning efficiency increases by 35%. At some

temperatures heat-resistant deformable nickel and titanium alloys have a decrease of plasticity

characteristics.

It was found in our experiments that in the LAM process, the cutting temperature

corresponding to the maximum efficiency of turning coincides with the plasticity decrease

temperature to a precision of a few percent. Decrease in the strength characteristics of the alloys

under consideration, gave the possibility of increasing 4-6 times the efficiency of draft and final

turning due to an increase in cutting speed.

3. 2. PLASTICIZATION OF THE MATERIALS

In the process of mechanical finishing of such materials as W, Mo, Cr, and zirconium

hydride, cast iron chip formation is determined by brittle separation caused by the periodic

process of outstripping crack development. Tear-outs and microcracks appear on the treated

surface. Preheating of the layer to be cut, up to and beyond the temperature of brittle plastic

transition, can change the process of chip formation. Destruction of the material follows plastic

P, N .--------==9

2

J

/J£l£l /51/1/

Fig. 3. The dependence of the components of the cutting force on the additional heating temperature in tungsten and molybdenum (+3% Nb) turning: V=60 m/min, S=0.05 mm/rev, t=0.2 mm. Curve 1- pz (Mo); 2 - Py (yV), 3 - pz (W).

158

Rz• ~m ,------------------,

/ill

Fig. 4. The dependence of surface roughness versus the heating temperature: S=0.05 mmlrev, t=0.2 mm. Curve 1 - Mo, V=60 m/min; 2 - W, V=80 m/min.

defonnation development up to some critical level. At a heating temperature higher than 200-300

°e, ( 500 °e for Mo, 600 °e for W), a continuous chip is fonned. Smooth grooves without tear­

outs and a net of microcracks are fonned on the surface. Roughness of the treated surface

decreases from Rz = 20-40 f.Ull to 1.0-5.5 mm. Some typical results are presented in Fig. 4.

Heating of zirconium hydride higher than 300 °e eliminates shears of the material when

the cutter enters the workpiece; there are fewer big tear-outs on the surface, plus decreased depth

and number of microcracks.

3. 3. THERMO-PLASTIC DEFORMATION OF THE MATERIALS

Local laser heating leads to an accumulation of thenno-plastic defonnations in the surface

layer of tough materials with high defonnation strengthening. Microhardness of samples treated

by the laser increases by 55% at a depth of 0.2 mm; this corresponds to a plastic defonnation

increase (according to the stress-strain diagram) of 15%. An estimation, for example of one of

the nickel alloys (EI 826), has showed that local plastic defonnation of the metal in the cutting

zone, when placing the cutting edge directly under .the thennally strengthened layer, leads to a

decrease of plastic defonnation expenses by 20% in comparison with unifonn strengthening. A

decrease in the tangential component of the cutting force by 30% was observed in the

experiments.

3.4. STRUCTURAL AND PHASE CHANGES OF THE MATERIALS

Under the influence of laser radiation in some materials being treated, structural and phase

changes, which change workability characteristics, may take place. Therefore, hardening and

159

annealing, depending on the initial state, can be perfonned on carbonic steels. Laser hardening of

the steel Y8 during LAM leads to a 2 times increase in the cutting force. However, in this case

treatment quality is considerably improved, and the turning groove depth becomes 2.5 /lffi

instead of 5.5 /lffi. The experiments demonstrate the distinct properties of this steel, with

preliminary hardened up to HRC 62 during LAM. Laser heating was perfonned at distances 1= 3,

5, and 8 mm from the cutting edge. When heating power grew to 1200 W and the distance 1

increased, one could observe: the decrease of cutting forces, the fonnation of a continuous chip,

and improvements in the treatment quality. These effects can be explained by the decay of

martensite in austenite under the influence of laser radiation. When 1= 3, and 5 mm, phase

transfonnations. may have no time to take place and only a small decrease ofPz, from 200 N to

150 N, is observed. When 1= 8 mm, Pz is as low as 40 N.

Preliminary laser annealing of ceramic materials (e.g., Si3N4, SiC), by changing the

strength characteristics of the layers being removed, make it possible to increase diamond tool

lifetime and treatment efficiency. Experiments on ZrC, SiC show that strength characteristics

decrease as a result of structural changes in the surface layer. We observed a 1.7-2.2 times

increase of lifetime of a cutter from superhard materials.

Laser remelting (both preliminary and during the LAM process) of the materials being

treated can also improve workability. Many materials contain hard large-sized inclusions

intensifying abrasive wear of the cutting tool. While melting, these impurities can dissolve.

Because of the rapid cooling in the process of crystallization after laser remelting, the size of

impurities can considerably decrease and their influence on cutter wear will be less strong. The

results of metallographic observation of remelting zones of the alloy Udimet 700, led us to

suppose that this factor promotes increase of the tool life time, when turning this alloy with laser

heating.

In the case of turning alloys of aluminium with silicon (the silicon content being higher

then 60%), because of high strength and hardness of the material grains, destruction is of a brittle

character. Big shears and tear-outs are fanned on the treated surface. High silicon content

intensifies cutter top wear and, as consequence, one cannot maintain the fonn of the treated

surface. Curvature of the surface exceeds 0.5 mm in a length of 100 mm. Laser heating of the

layer being cut, up to a temperature higher than the eutectic melting temperature (T= 577 QC),

leads to the melting of eutectic on alloy grain boundaries. Thus, shift and chip fonnation take

place mainly on grain boundaries. There are no tear-outs larger than the grain sizes. Cutting

forces and specific contact loads on the cutter decrease by 40%. Increase in the cutter life time

makes it possible to obtain curvatures of the treated surface of only 0.035-0.060 mm in a length

of lOOmm.

160

3.5. CREATION OF CONTINUITY DEFECTS

It is possible to decrease the force characteristics of mechanical treatment by creating

continuity defects in the cut layer, e.g., creation of thermal cracks, and partial evaporation of the

layer being cut. During zirconium hydride LAM, even when laser power is 100 W, formation of

evaporated craters and microcracks takes place on the surface being treated. Cutting forces

decrease by 70-75%. The optimum power range of laser radiation, providing the improvement of

treatment quality, has been determined. When P < 50 W, when the cutting depth equals 0.2 mm,

material heating is not sufficient and LAM does not differ practically from routine treatment.

When P > 220 W, cracks madc by laser heating fall outside the boundaries of the layer being cut.

In our experiments a Nd:YAG laser with frequency 400 Hz was used for preliminary creation of

holes (up to 0.5 mm) in the layer being cut (alloy Ti-6Al-4V). When the cutting velocity was

equal to 0.3 mis, the distance between the holes was about 0.8 mm. When the cutting depth t was

equal to 0.5 mm, a decrease of the cutting force by 30-45% was observed, and by 20% when

t=l mm.

3.6. CHANGES IN THE CHIP-FORMATION MECHANISM

The addition of laser-made holes near and parallel to the cutting edge, or melting of the

material being treated in the chip root from the back side of the tool surface, changes the chip­

formation mechanism. During chip formation, breaking away of the material takes place, in

contrast to routing cutting, where successive shift of the material occurs. As a result, the WOlX

necessary for material removal and the cutting forces can be reduced about one order of

magnitude.

Planing down of the plate from Inconel 718 alloy of 0 = 0.8 mm thickness with chip-root

melting gave the following results. The component Pz decreased up to 180 N. The ultimate

cutting speed increased from 22 mmls to 92 mm/s. Further increase in the speed is possible, but

the limited heating power P = 500 W did not allow us to achieve melting of chip roots when v> 92mm/s.

3.7. CHANGE OF FRICTION CONDITIONS

It is known that with an increase in the contact temperature (of the cutting surfaces with

the material being treated), the friction coefficient while cutting increases at first and then

decreases;the temperature of maximum friction is 400-600 °C when steels are treated by hard-

161

alloy cutters. Experimental results show that during routine cutting the average contact

temperature is close to the mentioned range, and during LAM it is as a rule about 800°C (Fig.5).

~DD ~::::2 ~ __ T!

::: I! I I I I 2S .fD 7.f TOO l.f 3D 73 V, m/min

Fig.5. The dependence of the cutting temperature of tungsten (a) and molybdenum (b) versus cutting speed for the different temperatures of additional heating; S=0.05 mm/rev, t=0.2 mm. Fig. a. Curve 1 -q=300 K; 2 - 1600; 3 - 1700; 4 - 1800 K. Fig. b. Curve 1 - q=300 K; 2 - 1200; 3 - 1300; 4 - 1400; 5 -1500; 6 - 1600; 7 - 1700 K.

t, min t. r------------, . M 170£1

I I T-'

,1./",'

.fll 7.f l/lll v, m/min

Fig. 6. The dependence of tool durability (T - solid line curves) and the length of the path of the tool until it grows dull (L - dashed curve) for the "cold" turning of tungsten (1) and turning with the additional heating qh = 1800 K (2) versus the cutting rate.

9M,K M~D·r--------------,

Fig.7. The dependence of surface tern perature of Mo in the middle of the heating spot for various powers of irradiation; V=60 m/min. Curve I - 1 kW, 2 -2, 3 - 3 kW.

162

One can therefore suppose that during LAM the coefficient of friction between the

material being treated and the tool material is lower than during routine cutting. As an example,

experiments on W-CO pair abrasion demonstrated three times decrease in abrasion, with rising

temperatures up to 400°C. In addition, during LAM specific contact loads on the cutter are

lower, than during routine cutting, because of the decrease in cutting forces and the increase of

the contact area of the chip with the cutter. As a result, during LAM specific friction forces and

intensity of tool wear decrease (Fig.6).

3. 8. EFFECT OF LASER PLASMAS

Due to a high value of energy density flux in laser action, a plasma cloud is formed over

the surface being treated. Saturation of the workpiece surface by the elements from the plasma

cloud creates conditions for embrittlement of the metal and decreases of plastic deformation

work during cutting. Losses of strength and embrittlement may also be caused by plasma

radiation. The laser plasma changes temporal, spatial and energy parameters of the heating

process. In our experiments self-excited oscillations of the surface temperature with an amplitude

up to 800°C, in a frequency range 1-20 kHz, were observed for a wide class of materials. The

general tendency of the oscillation amplitude increase with laser power is illustrated by Fig. 7.

As a consequence, quickly alternating thermal stresses appear in the material. Considering that

the frequencies mentioned are close to the frequencies of the cutter self-excited oscillations. it is

possible to match frequencies and to perform treatment under resonance conditions; this could

improve workability.

4. Conclusions

The experimental results show that in optimum conditions LAM significantly increases the

quality of machining: roughness of the treated surface. depth and rate of cold working decrease;

and shears and microcracks disappear. It is possible to perform: interrupted cutting of materials

with increased strength characteristics; turning of soft and/or thin-walled materials; finishing of

materials in which structural and phase changes below the cut line are not desirable.

Improvements in machining quality can be obtained even using lover power lasers. Generally.

the productivity of machining can be increased up to 4-6 times.

163

From the scientific point of view LAM can be used also as a powerful tool to analyse the

mechanical properties of materials at high temperatures.

Taking into account the general tendency of price decreases in laser equipment, the

appearance of powerful CW Nd:YAG lasers, etc; and on the other hand the increasing

requirements for higher quality and productivity of machining, the application of new materials

(oftenly difficult to machine), etc; it is possible to expect that eventually LAM will find its own

place among the other widely used energy beam technologies.

References

1. Reznikov, A.N. (1981): "Thennophysics of Machining Processes of Metals". Moscow, Mashinostroenie.

2. Stroshkov, A.N., Tesler, Sh.L.and Shabashov, S.P. (1977): "Cutting of Preheated Low­Machinability Materials". Moscow, Machinostroenie.

3. Bass, M., Beck, D. and Copley S.M. (1978): "Laser Assisted Machining", SPIE Vol. 164, Fourth European Electro-Optics Conference, Sira, October 1978, Utrecht, Netherlands, p. 233-240.

4. Jau, B. (1981): "Laser assisted machining", Ph.D. Dissertation, Chairman Prof. Copley S., Center of Laser Studies and Department of Materials Science, University of Southern California.

5. Tipnis, V.A., Ravignani, G.L. and Mantel, SJ. (1981): "Economic feasibility of laser-assisted machining (LAM)", Proc. 9th. NAMRC., p. 547-552.

6. Rajagopal, S. (1982): "Laser-assisted machining of tough materials", Laser Focus, March 1982, p.49-54.

7. Copley, S.M., Bass, M., Jau, B. and Wallace R. (1983): "Shaping Materials with the Laser", in (ed) Laser Material Processing, M. Bass, North-Holland Publishing Company, Amsterdam, p.299-336.

8. Gavryushenko, B.S., Okorokov, L.V., Rykalin, N.N., Uglov, A.A. and Smurov, I.Yu. (1985): "Laser-assisted metal cutting". Fizika i Khimiya Obrabotki Materialov, Vol. 19, No.2, pA-7.

9. Bashenko, V.V., Deich, A.Sh. and Karhin, V.A. (1987): "Laser Assisted Turning", in (ed) Laser Technology, A. Kanapenas, Vol. 3, Vilnus, p.79-85.

10. Bashenko, V.V., Deich, A.Sh., Karhin, V.A., Rimmer, I.S. and Safarevich, S.S. (1988): "Laser heating of the surface of the workpiece in finish machining", Fizika i Khimiya Obrabotki Materialov, Vol. 22, No.5, p. 114-120.

11. Marot, G., Fan, L.J., Tarrats, A., Cohen, P. and Longuemard, J.P. (1991): "Workpiece material­laser interaction and laser-assisted machining", Annals of the ClRP, Vol. 40, No.1, p.91-94.

HEAT PROCESSES IN PULSED LASER ACTION

I.Yu. SMUROV Ecole Nationale d'lngenieurs de Saint-Etienne 58, rue Jean Parat 42023 Saint-Etienne Cedex 2, France

A.M.LASHIN Istituto di Tecnologie Industriali e Automazione, via A. Ampere, 56 20131 Milano'/taly

ABS1RACT. Results of numerical simulation of pulsed laser action on metals are presented. The main feature of our model is the consideration of the movement of two phase fronts - melting and evaporation. Besides heating and melting, cooling and solidification are also taken into account, and this permits an analysis of the entire thermal cycle. One of the advantages of the present approach is the opportunity to analyze the dynamics of phase front velocities on the basis of an exact determination of the position of the phase boundaries. In the case of constant energy flux the transient periods of surface temperature, evaporation front velocity, melt thickness and their corresponding quasi-stationary values are determined. The influence of the shape of laser pulses on the dynamics of heat processes is considered. The evolution of surface temperature during pulsed laser action obtained by numerical simulation is compared with the experimental results. The influence of pulse interruption on heat process dynamics, and particularly on cooling-solidification rates, is simulated. The influence of the spike structure of laser pulses, corresponding to the cases of free and periodic spiking,is analyzed as it affects the evolution of thermal processes. The two-dimensional problem of melting-solidification in pulse laser action taking into account weak surface evaporation is considered.

1. Introduction

The processing of metallic materials by concentrated energy flows of differing physical

nature (laser, electron beam, plasma flows, concentrated solar radiation etc), specifically

hardening, welding, cutting, alloying, cladding, and glazing, constitutes a well known and

widely used technology [1-4]. Because of the great differences in technological requirements for

165

s. Martellucci et al. (eds.), Laser Applications/or Mechanical Industry, 165-206. © 1993 Kluwer Academic Publishers.

166

different processes, a wide range of operating parameters are used. For example, materials

remelting (or thennocycling) in a solar furnace can have a duration of about ten minutes under

the action of a constant energy density flow of nearly qo=103 W cm-2 [5-7]. The technology of

laser surface alloying calls for a unifonn layer of melt and therefore a virtual absence of

evaporation, which implies a comparatively low energy density flow and long interaction time

(qo=5·104 -105 W cm-2 , At -10-3 - 10-2 s) [4, 8-11]. The requirements for surface glazing and

treatment of coatings are not far from those mentioned above [12, 13]. On the other hand, for

drilling one needs extensive evaporation in order to prevent the fonnation of a thick melt layer,

but also it is necessary to minimize the heat affected zone. This means a comparatively high

energy density flow and short interaction time (qo=107 - 108 W cm-2, At -10-4 - 10-6 s) [1-3,

14].

At present, experimental investigations of the dynamics of physical - chemical processes

during pulsed energy flow action are based mainly on high-speed pyrometry, high-speed

photorecording, laser interferometry or similar methods [15-18]. By these means it is possible to

obtain precise infonnation about surface phenomena and the removal of the products of

interaction of energy flow with material (vapour and plasma flows, droplets of melt etc), but no

infonnation is derived concerning the interior of the irradiated material. Analysis of the results

following the end of energy action (usually on the basis of different cross sections) mainly can

help to reconstruct integral parameters, but not the dynamics of the process. That is why a

numerical simulation provides a unique opportunity to obtain infonnation that later can be used

to optimize the technological parameters of materials processing.

2. Mathematical model

The calculation of temperature fields during pulsed, pulse-periodic and continuous action

of energy flows on the basis of linear mathematical models is a comparatively simple task (for

example, [19,20D. The main difficulties appear during the simulation of phase transfonnations of

melting (solidification) and evaporation. The corresponding mathematical models are essentially

nonlinear. Analytical and numerical simulation of melting-solidification phenomena on the basis

of Stefan-type boundary conditions (or similar) are widely used in different branches of science

and technology [1-4].

In many cases, the destruction of materials under the action of energy flows is evaluated

using an ablation model [21-23] which assumes the removal of material from the zone of action

167

of heat source. From a mathematical point of view the ablation model can be considered among

the simplest approaches for analyzing the phase transfonnation. This model has its own field of

applications; from a fonnal point of view it can be used to detennine the position of the melt

boundary, mainly in three cases: (1) when practically all of the melt is removed from the wne of

action by one or another mechanism (flow of gas, reactive pressure of evaporation, etc); (2) the

temperature of the melt is close to unifonn (because of intensive remixing), so the absorbed

energy is used mainly for heating and melting the solid body and less for overheating of the melt

layer; and (3) the ablation model can be used to analyze the initial stage of melting when the

melt layer is thin enough (a few microns). From the mathematical point of view, the latter case is

equivalent to a constant temperature gradient in the melt, i.e., a the linear temperature

distribution. Consequently, the velocity of the melting boundary coincides with that of the

ablation front (both obtained by the Biot variation method) and the linear approximation is

accurate for small values of t-tm (tm - starting time for melting). Furthennore, the velocity of

melting becomes lower than the velocity of ablation; the difference increases with time, as the

melt thickness increases. This reduces the energy flux reaching the phase transition boundary.

One more qualitative difference needs to be pointed out: in the action of constant power density

on a semi-infinite body the ablation rate reaches a steady-state value; on the contrary, the melting

velocity attains a maximum value and thereafter monotonically decreases.

The simplest approach to numerical simulation of melting is the so-called enthalpy method

on a fixed grid. The phase change is assumed to occur over a certain temperature range and the

associated latent heat effect is handled by suitably increasing the specific heat in this range

[24,25]. More precise infonnation about the dynamics of a phase front (including its velocity)

can be obtained on the basis of the explicit tracking of phase boundary position with the help of a

moving grid [26,27].

The modeling of melting phenomenon without taking into account surface evaporation,

the most energy-demanding process, has a relatively narrow field of application. It is limited by

maximum values of surface temperature, or in other words (using technological parameters) by

energy density flow and duration of action [28]. It is well known that in the case of a one­

dimensional heating model (without surface heat losses), the action of constant energy density

flux causes an unlimited temperature growth; the same applies for a one-dimensional melting

model. Heat loss calculations (convective or radiative or both) from the irradiated surface leads

to the appearance of a steady-state surface temperature, whose value as a rule is outside the range

of practical applications.

The movement of an evaporation front under the action of high energy fluxes has been

intensively investigated during the last thirty years, since the appearance of the first

168

mathematical models as for example [29]. Until the present day these investigations could not

be considered as complete.

A number of works were devoted to the simulation of the movement of the evaporation

front under the action of laser irradiation on metals. As a rule they used the so called "heat

model" approach [30, 31] whose features are: a) the temperature fields in the material during

the action of relatively low energy density fluxes is considered to be independent of the

interaction of laser beam (energy flow) with the evaporated material; b) the determination of the

melting front position without use of the Stefan type boundary condition, but only on the basis of

the corresponding position of the melting point isotherm Tm; c) description of the evaporation

heat losses in the boundary conditions (on the irradiated surface) by a comparatively simple

expression whose form is close to the Herz-Knudsen law of evaporation.

An alternative to this approach is the detailed consideration of evaporation kinetics based

on the theory of the Knudsen layer [32, 33]. It allows us to consider the influence of gas­

dynamic disturbances in the flow of evaporated material and to obtain some new results using

considerably more sophisticated mathematical models.

For modeling the heat processes of concentrated energy flow action on metals it is

necessary to join both the processes of melting and surface evaporation in the frame of a single

mathematical model: this leads to the situation where we must model two phase boundaries

moving at different velocities. Such a model makes it possible to determine the dynamics of

melt thickness over a wide range of values of energy density flow and duration of action. This

information (and the corresponding temperature distribution) is necessary to optimize the

technological parameters of a number of processes including alloying, cladding, and glazing.

The mathematical model proposed includes the processes of heating, melting, evaporation,

cooling and solidification under the irradiation of arbitrary energy flow on a metal slab [4, 34-

38]. It is assumed that the energy flow is absorbed on the irradiated surface; convection and

radiation mechanisms of heat losses from both the sides of the slab are considered; melting

(solidification) is determined by a classical Stefan-boundary condition.

Energy flows of differing physical natures each exhibit their own peculiarities. For the

case oflaser action, in some cases it is necessary to take into account the variation of absorptivity

due to temperature changes and due to chemical reactions (mainly oxidation) [4. 39, 40]. The

increase of acceleration voltage in electron beam treatment transforms the effective heat source

from surface to volume [1,2]. One of the limits of the application of the present model is the

maximum value of energy density flux, above which it is necessary to consider the interaction of

energy flow with the products of the destruction of the target (vapours, plasma cloud, droplets of

melt, products of condensation of vapour phase etc). It is possible to neglect convective heat

169

transfer in the melt in comparison with conduction, for the range of parameters discussed in

detail in Refs,[4, 38]. The possibility of using a classical parabolic type equation for the

description of high frequency temperature oscillations is discussed in Refs,[38, 41].

From our point of view thc present model can be considered, on the one hand, as one of

the simplest to take into account both melting and evaporation phenomena; on the other hand, it

can be used over a comparatively wide range of energy density flow and pulse duration values

[42,43].

The mathematical model used can be written in the following form:

dSl dt

A, aTl - A, aT2 _ L dS2 1 ax - 2 ax P2 m dt '

T( x, t=O) = TO

x=Sl (t)

(1)

where: Tl (x,t) = temperature of liquid phase; T2(x,t) = temperature of solid phase; x, t =

distance and time, respectively; S 1 (t),S2(t) = positions of evaporation and melting phase

boundaries, respectively; qO(t) = absorbed energy density flux; al,2 = thermal diffusivities of

liquid and solid phase, respectively; A,1,2 = thermal conductivities of liquid and solid phase,

170

respectively; Pl,2 = densities of liquid and solid phase. respectively; Lv = specific heat of

evaporation; Lm = latent heat of melting; a g,f = convection heat loss coefficients of irradiated

and rear surfaces of the slab; £ g,f = emissivities of the environment near irradiated and rear

surfaces of the slab; £ 1,2 = emissivities of irradiated and rear surfaces of the slab; Tm =

temperature of melting; tm = starting time for melting; L =thickness of the slab; TO = initial

temperature. Constants V. , T. are detennined by Herz-Knudsen's law of evaporation:

V. = Y exp v • p, [L] 2Pl(21tk/m) Tv (kIm)

T-~ • - (kim)

Where: k = 1.38 10 -23 JIK, the Boltzmann constant; m is the atomic mass of the slab material;

and Tv is the boiling temperature corresponding to the pressure Pv [43.44].

Equation (1) are solved numerically; the phase change fronts are tracked continuously and

the latent heat release is treated as a moving boundary condition. In both the regions of liquid

and solid phases the moving unifonn grids consist of a fixed number of points. Each grid point

moves with a different velocity. The Crank-Nicolson technique of various derivatives is used

[45-47].

Calculations were made for titanium and steel slabs of 1 mm thickness. In all the cases

examined below the slab can be considered as a semi-infinite body.

3. Constant energy density flux

Let us now consider the action of aconstant energy density flow on a slab with thennal

properties close to those of titanium. In all cases calculations were continued up to the time

when values of surface temperature and velocities of both phase boundaries reached their

corresponding quasi-stationary values.

The velocity of the evaporation front monotonically increases with time, approcbing a

constant, time-independent value. The velocity of the melting boundary. first quickly increases.

attains a maximum value, and then more slowly decreases to the same value of quasi­

stationary velocity as for the evaporation front (Fig. I). The reason for this unusual type of

melting velocity behaviour is an increase in melt overheating (relative to the melting point) on

the one hand, and an increase in melt thickness on the other. The ratio of the extreme value of

melting velocity to the corresponding steady-state quantity is in the range of approximately 5 to

8 for an energy density flux range loS - 108 W cm-2. Starting from the time when both phase

15

10

5

600

400

200

20

12

4

o

-1 V [em 5 ]

T/Tm

10 20

171

30 t ems]

Fig.I. Simulation of heat processes in the action of a constant energy density flux qo= 105 W/cm2 on titanium. The figure illustrates the time dependence of: melting (1) and evaporation (2) front velocities; positions of melting (la) and evaporation (2a) fronts; melt thickness (3), surface temperature taking into account evaporation phenomenon (4) and neglecting evaporation (5).

172

fronts move at the same velocity, the melt thickness stops increasing and reaches its quasi­

stationary value.

The comparison of surface temperature dynamics, with and without taking evaporation

into consideration, reveals not only quantitative but also qualitative differences. The

temperature of the evaporating surface quickly reaches its steady-state value, that is of the order

of the melting point temperature. On the other hand, neglecting evaporation leads to an unlimited

increase in the calculated temperature of the irradiated surface temperature, which at the

conclusion of energy flow action can be ten times higher than the melting temperature.

Quasi-stationary values of melt thickness and irradiated surface temperature versus

absorbed energy density flow are presented in Fig.2. This kind of dependence is important to

analyse, because in the case of concentrated energy flux (laser, electron beam, plasma) action on

metals, the energy density flow varies on some orders of magnitude (usually from 103 to 106 W

cm-2). A first rough estimation can be made on the basis of the above mentioned curves. The

melt thickness rapidly decreases with the increasing of qO values (from the value 120 J.UI1 that

corresponds to an energy density flow of 105 W cm-2, to 0.3 J.UI1 that corresponds to 108 W

cm-2, i. e. some 400 times). A comparatively weak dependence of surface temperature on

energy density flow (for the same energy density flow range, from 2.S Tm to STm ) is the

result of extensive evaporation, which prevents an unlimited temperature increase as occurs

in the case of simulation of melting without evaporation. It is important to note that above a

value of approximately qo=106 W cm-2 the evaporation products interact with the laser beam

and affect the heat flow within the metal [1-4]. A value of qo=106 W cm-2 is the limit of the

present model (1) relative to the value of energy density flow. For higher values of qO, the

results are interesting from a phenomenological point of view.

The system (1) has an accurate analytical solution only for the condition that

vst=dS l/dt=dS2Idt:

dS = _~ In A2(Tm -To}/a2 +pLm V A2(Tm -To}/a2 -AI(TI-Tm}al +p(Lv +Lm}

dS = S2 - Sl

173

Detennination of the transient period, Le., the period for which the transient solution

varies by m% with respect to the corresponding quasi-stationary value, is of practical interest.

The results presented in Fig.2 correspond to the value m=5%. As qO increases, the transient

120

100

80

as [J.lm ] TIT m

60

6

40

4

20

2

-2 10

-4 10

t [5] -6

10

-8 10

5 6 7 8 10 10 10 10

q [W em -2]

Fig.2. Dependence on absorbed constant energy density flux of: quasi-stationary values of melt thickness (1), surface temperature (2); transient periods of surface temperature (3), evaporation front velocity (4), melt thickness (5). The material is titanium.

174

period decreases, primarily due to the increase in phase boundary velocities, and to a decrease in

the heat propagation length (into the bulk of metal). On a logarithmic scale, these functions

vary linearly as a function of qO. By studying the melt thickness transient period in the well­

known form tst =K3lvst2, where Vst is the corresponding quasi-stationary velocity of phase

boundaries and a is the thermal diffusivity, one can obtain constant value K=1.6 (for the case of

Ti). The simulation shows that it is necessary to distinguish three different transient periods: the

smallest for the surface temperature, the largest for the melt thickness and melting front velocity,

and an intermediate period for the evaporation front velocity. This can be explained by the fact

that the quasi - stationary temperature field that determines the velocities of phase boundaries is

initially formed near the irradiated surface and only later propagates some distance into the

bulk metal.

4. Pulsed energy flux

Although it is possible to study the dynamics of heating and melting when one uses a

constant value of qO, no information about cooling and solidification after the end of energy

application can be obtained. To study cooling and solidification one can consider the effects of a

single energy pulse. Heat processes resulting from the application of rectangular energy pulses

(whose energy and duration correspond to pulsed laser action) have been considered [4, 36].

4.1. INFLUENCE OF TIIE SHAPE OF LASER PULSES

The influence of the shape of energy pulses on heat process dynamics is considered for

triangular pulses which have the same energy input, temporal duration and maximum value of

energy density flux. The only difference is the position of the extreme value of energy density

flow qrnax, which leads to different rates of increase (and decrease) of energy flow. In this

numerical experiment the pulse duration is 4 ms, qmax=105 to 5·105 W cm-2, and energy per

pulse E= qrnax /2 (energy per square cm).

Fig. 3 shows the results of the simulation of heat processes for a triangular pulse with a

maximum value of energy flow density qruax=5.105 W ~m-2 , which is reached 1 ms after the

action begins. These results are noticeably different from the case of constant energy density

flux. The evaporation front velocity closely follows the temporal profile of the energy pulse. In

175

this case, the transient period of the velocity is comparatively short, tst=O.1 ms; that is why the

velocity change follows the comparatively slow variations of energy density flow.

The velocity of the melting front shows a complex pattern. First, there is a sharp increase

in velocity up to 13.8 cm/s. Next, there is a sharp decrease, and the velocity drops by almost one

half. Third, there is a slow increase to 8.5 cm/s. Finally there is a monotonic decrease to negative

values, corresponding to the transition from melting to solidification; thus, the velocity of

melting is a nonmonotonic function of time with two peak values. The first peak is caused by the

increase of surface temperature and the thickening of the melt layer, which increases the thermal

.c -1

V [em 5 ] 0

12 ~ n(JI

8 3 1

N

4 4

2

0

-4

250 5, .15, [~m] 200

150

100

50

0

3

2

o 2 3 4 5 6 t [ms]

Fig.3. Time dependence of: melting (1) and evaporation (2) front velocities; positions of melting (la) and evaporation (2a) fronts; melt thickness (3), surface temperature (4). Shape of energy pulse:: dashed curve. Material: Ti.

176

resistance. The second peak is a direct result of the interaction of the two phase boundaries: the

velocity of motion of the evaporation front increases up to values exceeding the melting velocity,

which causes the melt layer to become thinner. This is the reason for an increase of heat flux

reaching the melting phase boundary and therefore for an increase of melting velocity. The

maximum thickness of the melt layer was 84 flIIl, and it was reached approximately 0.5 ms after

the radiation impulse ended, when dS2Idt=O. This is the result of the heat stored in the melt.

In the case of a rectangular energy pulse the second maximum of the melting velocity does

not appear [4, 36]; because there is no extreme for the evaporation front velocity, the latter

monotonically increases during pulse action and even for qo=106 W cm-2 is smaller than the

melting front velocity.

When the maximum value of power density is reduced to qmax=2.5.loS W cm-2 with the

duration and shape of the pulse remaining the same, the second, lesser peak of velocity of

melting disappears since the velocity of evaporation does not exceed that of melting. Also, the

similarity between the evaporation front velocity and the shape of the energy pulse disappears if

Qrnax<105 W cm-2 because the transient period increases.

Fig. 4 shows the results of a simulation of the application of a triangular pulse with the

same maximum value of qmax but occurring 3 ms after the pulse has started. The shape of the

curve is the inverse of that in Fig. 3. The above mentioned regularities are qualitatively the same

as in the previous case, but it is possible to note some quantitative differences. First, maximum

value of melting velocity decreases to 9.6 cm/s. Second, the second extreme of the curve of

melting velocity can be seen more distinctly (its value is close to the first extreme). Third, the

maximum thickness of the melt layer (67 flIIl) is considerably smaller than in the previous case.

Fourth, the melt thickness is characterized by the existence of a visible minimum (41 flIIl) when

q- qmax (when evaporation is more extensive). Finally, decreasing energy flow, while keeping

the rest of the parameters constant, leads to changes that are similar to those in the previous pulse

shape.

The evolution of heat flow corresponding to an energy pulse in the shape of an equal­

sided triangle was considered. Pulse duration was equal to 4 ms and qmax=105 - 5·loS W

cm-2. For this case, the velocities and positions of phase fronts fell between the cases shown in

Figs. 3 and 4.

The dynamics of surface temperature depend on the shape of the energy pulse to a lesser

extent than the velocity of motion of the phase fronts. For all triangular pulses with the same

duration and qmax energy per pulse this dependence is weak. After an initial sharp increase, the

temperature in all cases studied reached approximately 5000 K, with the same maximum value of

5190 K reached when the energy flow density qmax equals 5· loS W cm-2. After the pulse ends,

177

the surface temperature drops sharply to values near the melting point of the metal, and it

remains at this level until the end of solidification. When qmax decreases, there appear

differences in surface temperature dynamics.

It should be noted that parameters such as the depth of the propagation of the melting front

and the thickness of the evaporated layer depend only weakly on the shape of the energy pulse,

provided that it's duration, qmax and energy per pulse remain the same. However, differences

sharply increase when qmax decreases to the threshold of melting.

.D

-\ V [em s 1 0

12 ~ (')U1

3 8 I

..JV

4

2

0

-4

250 5, ~5, [Ilml 200

150

100 3

50

0

T/T m

3 4

2

o 2 3 4 5 6 t [msl

Fig.4. Time dependence of: melting (1) and evaporation (2) front velocities: positions of melting (Ia) and evaporation (2a) fronts; melt thickness (3), surface temperature (4). Shape of energy pulse _ dashed curve. Material - Ti.

178

From these results, it seems possible to make some practical recommendations for

materials treatment by concentrated energy sources. To produce a larger thickness of the melt

layer, it is better to use pulses for which the maximum energy flow occurs near the beginning of

the pulse. In this case more energy is used to heat the solid body and therefore to initiate the

beginning of melting. This type of pulse shape can be optimum, for example, in laser alloying. In

contrast, for material removal, it is better to use pulses for which Qrnax occurs near the end of

the pulse, when extensive evaporation can force out the melt. This type of energy pulse can be

used, for example, in drilling.

4.2. THE EFFECTS OF REALISTIC LASER PULSE SHAPES

For some applications, such as pulsed laser treatment, it is necessary to improve the

accuracy of the calculations by considering the real shape of the laser pulse, because the

results can differ significantly from the results obtained for rectangular and triangular pulses.

Fig. 5 present the results of laser irradiation of a titanium slab with a pulse shape which

depends on time as follows: q=at(b-t) (Qrnax=s·l05 W cm-2, O<t<b).

The heat flow is qualitatively the same as for triangular pulses. In addition, the dependence

of the evaporation front velocity on time closely approximates the shape of the laser pulse,

because the corresponding transient period is less than the pulse duration. The melting

velocity curve again has two maxima, the first at the beginning of melting and the second close

to the maximum of the evaporation front velocity. The increase of the evaporation front

velocity up to values larger than the melting front velocity leads to a decrease in melt thickness

in the middle of laser pulse.

After the end of the energy pulse the melting still continues (Vm>O), and only later does

solidification commence (Vm<O). This post-pulse melting is due to overheating of the melt, and

its duration depends mainly on the overheating rate and the thickness of the melt, i.e., on the

energy stored in the melt. The solidification velocity curve is also characterized by such

extreme dependence. If the maximum values of melting velocity increase nearly

proportionally with q, a different behaviour (weak dependence on energy flux) is observed for

the extreme values of solidification velocity. The increase of solidification rate with time is the

result of a decrease of temperature gradient in the melt after the end of laser pulse. After the

surface temperature decreases nearly to the melting point (the temperature gradient in the melt is

practically absent), the most important phenomenon is the amount of energy which appeared

during solidification of some volume due to latent heat of melting. This is the reason for a

comparatively weak dependence of the extreme value of solidification velocity on energy flux

179

.0

16 V [em s-~ 5 0(.11

".. ....... I , ~

I \ n

12 I \ 4 3

I

/ \ ~

\ I \ 3 I \ 8 \ \ 2 1

4

0

-4

300 5, .1.5, [J,lm]

200

100

0

2.6

2.2

1.8

1.4

1.0 2 3 4 5 6 t [ms]

Fig.5. Time dependence of: melting (1) and evaporation (2) front velocities; positions of melting (1a) and evaporation (2a) fronts; melt thickness (3), surface temperature (4). Shape of laser pulse: dashed curve. Material: Ti.

180

density, because it is detennined mainly by the properties of the metal (in particular Lm). Later,

the absolute value of the solidification front velocity will decrease because of the decrease of

temperature gradient in the solid phase near the solidification front. During the final stage of

solidification, the most important heat loss is heat transfer into the bulk of metal, and this

depends mainly on the properties of the metal.

The initial value of the melting velocity is zero, Vm(tm)=O, so all the melting curves

start from the V=O axis. In contrast, the values of solidification velocity, corresponding to the

end of solidification, usually are not equal to zero and the curves do not reach the V=O axis.

In some cases, when the heat losses from the surface of the melt are strong enough,

solidification can also start from the surface and two solidification fronts move towards each

other because of heat transfer into the bulk of metal and surface heat losses. The appearance of

the second solidification front is most typical of the final stage of solidification.

The evolution of heat processes corresponding to the same shape of laser pulse, but with

Qmax five times smaller, is qualitatively the same as in Fig.5, except for the disappearance of

the second melting velocity extreme. This is the result of weak evaporation (whose velocity is

much less than the corresponding melting front velocity). The shape of evaporation front

velocity differs from the laser pulse shape because the transient period for the smaller energy

flux is larger than or comparable to the pulse duration.

The increase of energy density flow up to 106 W cm-2 for the same shape of laser pulse

leads to an increase of the second extreme of the melting front velocity up to values larger than

for qmax=5.105 W cm-2. Forthe time interval 0.8 to 3.5 msec the velocities of both phase fronts

are almost the same (and also follow the shape of energy pulse), and the melt thickness is small

(- 25 J.UI1) and practically independent of time. In contrast to the constant energy density flux

case the velocities of the phase fronts are not constant, but vary in time over a wide range of

values.

The comparison of the cooling rates corresponding to laser pulses of the same "bell­

shape", with the same duration of 4 ms, and with Qmax=105• 5.105• 106 W cm-2 , respectively,

reveal unusual results: the larger the energy input, the smaller the surface temperature after the

end of laser pulse. Actually, the surface temperature reaches the melting point at 8. 5.5 and 5 ms

for qmax=105; 5·105; 106 W cm-2 , respectively. In order to explain this phenomenon it is

necessary to remember that for larger energy fluxes, the melt thickness decreases (see Fig.2). The

values of surface temperature during the period close to the end of the pulse are limited mainly

by evaporation and are not far from one another (nearly 2 T m). Also, the higher the temperature,

the higher the heat losses from the surface. At the bottom of the melt pool, the temperature is

equal to the melting point; consequentially, the relaxation of temperature distribution obtained

181

during laser action is detennined mainly by melt thickness. The larger the length of the heat

affected zone (since the temperature on both sides is fixed), the smaller is the temperature

gradient and the smaller the cooling rate.

The spatial distribution of temperature (for the case of the same "bell-shaped" laser pulse

as in Fig. 5, but with duration 1 ms, qmax=105 W cm-2 ) is presented in Fig. 6. The irradiated

material, steel, has a melting temperature Tm=1808 K. Because of the small melt thickness, the

temperature distributions approximate straight lines. It should be noted that because of the

comparatively small energy input, cooling of the surface starts at 0.75 ms and solidification at

0.86 ms, both of which are during the laser pulse. In this case the curve of melting front velocity

has only one very gently sloping maximum.

It is possible to distinguish two different regimes in the dynamics of the melting -

solidification front. The first regime, corresponding to a comparatively high energy input, is

characterized by the early appearance of the melt tm « ~t, by continuation of melting after the

end of the laser pulse ts > ~t, and by the comparatively long lifetime of the melt (up to about

three times as long as the pulse duration [4, 36]). Here dt=pulse duration, tm=starting time for

melting, and ts = starting time for solidification, i.e., dS2(ts)/dt=0.

2300

2050 ~

cu L :::J 1800 .oJ IV L cu 0..

E 1550 cu

I-

10 20 30 40

D1stance [ Ilm ]

Fig.6. Spatial temperature distribution in a steel slab during pulsed laser action. Curve 1 corresponds to the time 0.5 ms; 2 - 0.6; 3 - 0.75; 4 - 0.9; 5 - 1.0; 6 - 1.01 ms.

182

The second regime, corresponding to comparatively low energy input, is characterised by

the late appearance of the melt tm '" at, by the beginning of solidification during laser action

ts < at (up to the end of solidification tsol before the end of laser pulse tsol < at), and by the

TIT m

1.6 2

1.4

1.2

1.0

2 4

1.8

1.6

1.4

1.2

2 4

a .c

0

:(UI n 3,

3.....}\l

2

-3 t [ 10 s1

b o :(UI n 3 , .....}\l

6

4

2

-3 t [ 10 s1

Fig.7. Evolution of the surface temperature of a Ti slab during pulsed laser action. Curve 1 - shape of laser pulse; 2-4 results of numerical simulation corresponding to the absorptivity values 1.0; 0.8; and 0.45 respectively; 5 - experimental results; a - laser pulse energy 1.2 J; b - 3.0 J.

183

comparatively short lifetime of the melt (up to the absence of melting).

The comparison of calculated and experimental results is presented in Fig. 7. The shape of

the surface temperature curves agrees both qualitatively and quantitatively with the results of

pyrometric measurements [15, 16]. The shape of the laser pulse corresponds to the Russian

tlKvant-16t1 Nd:YAG laser. There is good agreement with the experimental results, especially for

the case of larger energy inputs (Fig. 7b) which result in extensive evaporation, and thus limit the

temperature rise. The differences in the lower temterature ranges (Fig. 7a) is the result of

neglecting the temperature dependences of the absorptivity and thermal properties of Ti.

4.3. INFLUENCE OF PULSE INTERRUPTION

The influence of the laser pulse shape on heat process dynamics is evident. It could be one

of the most influencing factors in, for example, laser glazing, which appears at high cooling

rates. It could be useful to modify some typical parameters of laser action, such as pulse duration

and the rate of decrease of energy flux at the end of the pulse, to obtain a high cooling rate of the

irradiated material. One of the simplest possibilities is to interrupt the laser pulse by using on

opto-electronic or mechanical device (for example, a shutter, a perforated rotating disk, etc). The

aim of this paragraph is to study the influence of pulse interruption on heat process dynamics.

The basic shape of the laser pulse under consideration is parabolic, given by the following

formula: q(t)::;;at(.:1t-t), where M ::;; pulse duration and a:::;4qmaxl.:1t2, being the maximum flux

reached at t= .:1t/2. We consider laser pulses with qmax equal to 105 and 106 W cm-2, .:1t = 1

ms. These parameters are typical for Nd:YAG lasers. We consider three different times tl when

the laser pulse may be interrupted, equal to 1/4,1/2, and 3/4 of the laser pulse duration. These

conditions of laser pulse interruption are chosen because they correspond to different rates of

change of the flux: a)- increasing energy flux; b)- the maximum value of energy flux; and c)- the

same value of energy flux as in the a) case, but decreasing.

4.3.1. Surface temperature

The surface temperature versus time is presented in Fig. 8, where the upper solid line

curve corresponds to qroax equals 106 W cm-2 and the lower to 105 W cm-2. A comparatively

large.energy density flow has been chosen so that we can observe the influence of evaporation on

cooling rates following pulse interruption.

The shape of the lower curve almost duplicates the shape of the incident flux. The upper

curve has a different shape, with a weakly varying maximum temperature: it is limited by the

184

~ 2800 =>

-<

"' bJ2300 .. '" ~ ~

1eoo

1300

800

\3 \

2

0.215 0.15 0.715 I. 25 1.!5

TIME (t.lS!:C)

Fig. 8. Time dependence of irradiated surface temperature. The solid curves (1-qmax=106W cm-2; 2-105 W cm-2) correspond to uninterrupted parabolic pulses. Dashed curves correspond to the pulse interruption at: 0.25 msec (3),0.5 msec (4,5), 0.75 msec (6,7). Curves 5,7 are for the ijmax=105 W cm-2 pulse.

evaporation phenomenon due to the higher incident flux (106 W cm-2 instead of 105). 1he

dashed curves correspond to laser pulse interruption after the above mentioned time periods

(with the exception of ti=0.25 msec for the lower energy pulse, where there is no melting).

During the cooling portions of both curves it is possible to see "steps" at a temperature

equal to the melting point (in the present case 1808 K), corresponding to solidification. 1he

duration of these "steps" depends on the interruption time: it is larger without interruption, and

decreases monotonously with earlier times of interruption.

4.3.2. Melt thickness

The dependence on time of melt thickness is given in Fig. 9 : the upper curve represents

the action of a 'lmax=106 W cm-2 laser pulse, the lower loS W cm-2. As in Fig. 8, the solid

curves correspond to uninterrupted pulses. The comparatively lower values of the upper solid

line curve are the result of intensi ve evaporation, which decreases the melt thickness. That is why

the melt thickness values, despite a ten times difference in energy flux, are comparable.

For interrupted pulses, we observe a slight increase in the melt thickness after the

interruption of energy action. This may be explained as follows: As soon as the interruption

occurs, the evaporation rate decreases almost to zero. At the same time, melting continues after

interruption, because of the large temperature gradient in the liquid phase.

2~

u

~

'" 20

'" .., z ~

u - 1 ~

'" ~

10

0.215 o.~

, , 2

" , , , , , , ,

\ ' \ '6 , ,

\ \ \ \ 1 \

0.715

\ \ \ \

\ \ \ \

MELT THICKNf;;SS

1.28

TIME (MSEC)

185

Fig. 9. Time dependence of melt thickness. Solid line curves (1-Qmax=106 W cm-2; 2-105 W cm-2) correspond to uninterrupted parabolic pulses. Dashed curves correspond to pulse interruption after: 0.25 msec (3),0.5 msec (4,5), 0.75 msec (6,7). Curves 5,7 are for the qmax=105 W cm-2 pulse.

4.3.3. Phase/ront velocities

Let us consider first the action of the pulse with Qrnax=105 W cm-2. For the uninterrupted

pulse the solidification starts during the pulse and at a relatively long time (0.15 msec) before its

end (Fig.lO). Generally, the thicter the melted layer, the longer will be the solidification period.

-0.' :s

-0.2.5

-0.35

" " " " "2 " I' 1\

" " "

, , \ , " \~,

3

IME 1(MSEC

Fig. 10. Time dependence of melting/solidification front velocity for a parabolic laser pulse (qmax=105 W cm-2). The solid curve corresponds to the uninterrupted pulse, while the dashed curves correspond to pulse interruption after: 0.5 msec (curve 2), 0.75 msec (3).

186

_ 0.2

'" '-'" >-.-- 0.1 .., o ~

~

>

-0,1

-0.2

o.~a I

I

\4 I \ I I .~

II

S" " I!

o.p I I I I ,6 I

I I I 'J

II II

711 \I

o.~. 1. • \ \ I I \

~\"E \ MS£C

'8 I

I , -

Fig. 11. Time dependence of melting/solidification (curve 2) and evaporation (1) front velocities for a parabolic laser pulse (Qrnax=l06 W cm-2). Solid curves correspond to the uninterrupted pulse, dashed curves correspond to pulse interruption after: 0.25, 0.5, 0.75 msec. Curves 3, 5, 7 correspond to evaporation; curves 4, 6, 8 to melting/solidification.

The increase of the solidification rate when the interruption time is 0.5 msec in comparison with

0.75 msec is mainly due to the small amount of energy accumulated in the melt: this layer is

only slightly overheated (under the melting point) when ti equals 0.5 msec.

The weak influence of the interruption time on solidification phenomena (pulse with

-~

......

" w

~ a:

0.2:5 o.a

Fig. 12. Time dependence of surface heating/cooling rate. Solid curve: qmax=105 W cm-2, uninterrupted pulse. Dashed curve: pulse interruption in 0.75 msec.

187

Qrnax=106 W cm-2), is due to the only minor changes in the melt layer parameters (thickness,

temperature distribution) during the time period considered (Fig. I 1). The initial conditions of

solidification are essentially identical. The solidification velocity would only be increased with a

more reduced interruption time (for example, close to the melting time tm).

4.3.4. Temperature rate

In contrast to the weak dependence of the solidification velocity on the conditions of pulse

interruption, the surface cooling rate is strongly influenced by the interruption time.

In Fig. 12, the sharp decrease of the heating rate in the middle of the pulse (about 0.5

msec) corresponds to the beginning of melting; it is explained by phase transformation, and by

the increase in the emissivity from solid to liquid phase (from 0.45 to 0.65). It is possible to see

that cooling (negative values) starts at three quarters of the pulse duration, as previously

mentioned. Zero values of cooling rate after I msec correspond to the final phase of

solidification, when the melt temperature is close to the melting point. The sharp decrease later

indicates the end of solidification.

In the case of the higher energy pulse, the heating cooling rate is close to zero during the

long period in the middle of the pulse which corresponds to small variations in the surface

temperature (see Fig. 13). When an interruption occurs, the cooling rate reaches a high value: the

shorter the interruption time, the higher the cooling rate.

w I-00{ a:

I( 0: ~," 'i " " I, I·

" ,,3

~ , , , ,

....... 0 .:7~' , ' " " I' :4

TI E (MSEC)

Fig. 13. Time dependence of surface heating/cooling rate. Solid curve: Qrnax=106 W cm-2, uninterrupted pulse. Dashed curves: pulse interruption after 0.25 msec (curve 2); 0.5 msec (3); 0.75 msec (4).

188

To explain this, it is necessary to keep in mind that even for the same value of surface

temperature the heat affected zones are different: the shorter the interruption time, the smaller the

heat affected zone, and the larger is the heat flux inside the bulk of the metal.

On the basis of the foregoing analysis, it is possible to recommend for pulsed laser glazing

a lower energy density flow, with an interruption before the beginning of the cooling period.

4.3.5. Spatial temperature distribution

Curves 1-3 (Fig. 14) show the spatial temperature distribution during the period of pulse

action, curves 4-6 after pulse interruption. The surface temperature decreases strongly after the

interruption: 0.01 msec after interruption, the decrease is by one-half. Later on, in contrast, the

temperature in the melt is practically uniform (for curve 6 the temperature difference is less than

10 K) and does not vary in time. It is possible to distinguish three typical periods:

- (1) During laser action the temperature distribution can be represented by practically

straight lines; the temperature gradient is roughly determined by the absorbed energy flux (qabs)

and the heat conductivity (A.) : dT/dx=qabs fA.;

- (2) This is followed by a short transition period; the surface temperature decreases

rapidly because of surface heat losses (including evaporation); the spatial distribution inside the

metal varies slightly; and as a result the temperature in the melt is represented by a "parabolic"

type curve (generally, if surface heat losses are strong enough, the temperature distribution in the

melt can have an extreme).

TEMPERATURE DISTRIBUTION

10 20 30

distanco (mic)

Fig. 14. Spatial temperature distribution for the interrupted (ti=0.75 msec) laser pulse. Curve 1 corresponds to the time 0.5 ms; 2 - 0.58; 3 - 0.748; 4 - 0.76; 5 - 0.788; 6 - 0.8 ms.

189

TEMPERATURE GRADIENT IN LIQUID PHASE

;::- 20 ~

'" " 2 < '" I '. ,', ~ I

I \\, '"

, , , , \ \ 0 , ~

, , , , , , \ ' , , I I , , , , , , ' ,

10 , , ' , :3 :4 , :S , ,

i , : ,

6i , , , , , , , , , , , , , , , : : ,

i , , l , , , , , , , i , , , , , ,

0.5

TIME (MSEC)

Fig. 15. Time dependence of the average temperature gradient in the melt. Solid curves (l-qmax=105; 2-106 Wcm-2) correspond to uninterrupted parabolic pulses. Dashed curves correspond to pulse interruption after: 0.25 msec (3),0.5 msec (4), 0.75 msec (5, 6). Curve 6 is for the qmax=105 W cm-2 pulse.

- (3) The final period occurs a short time after the end of energy action; it describes the

greatest part of the solidification process; the temperature distribution in the melt is practically

uniform and equal to the melting point.

Let us remark that the negative surface heat flux at the beginning of cooling (second stage)

reaches a value of 5.102 W cm-2. After the completion of solidification there occurs an ordinary

cooling of the solid body.

4.3.6. Temperature gradient in the melt

It was shown above that during practically the entire period of existence of the melt, its

temperature distribution can be represented by a straight line. Thus the temperature gradient in

the melt is expected to vary only slightly during laser action, and to equal zero during the

solidification period (Fig.15). For technological applications the average temperature gradient

through the whole melt thickness (I1T/l1x= (T(x=Sl(t), t) - Tm)/(S2-SI» could be more

important than just the gradient at the surface.

4.4. MAXIMUM MELT THICKNESS AND LIFETIME OF THE MELT

Taking into account the preceding considerations, we can address the behaviour of

190

maximum melt thickness for a given (constant) energy input into the material, for different

energy density fluxes and corresponding values of pulse duration. Two asymptotic cases are

easily considered: the first, corresponding to low values of energy flux densities and long pulse

duration, and the second, exactly opposite, for high densities of energy flux and short pulse

duration. In the first case, a large part of the energy will be spent on material heating (tm - M)

up to the disappearance of melting. In the second case, strong evaporation will be the most

important process leading to heat loss and since Lv»Lm only a comparatively small part of the

energy input will go into melting. Therefore, since both asymptotic cases, as a function of qo ,

shown a small melt thickness then for same intermediate value of go there must exist a

maximum.

Let us consider the action of a pulse with energy equal to 10 Joules, which is focused in a

spot with radius about 1 mm. These values of pulse energy and of focal spot size are typical for

pulsed laser treatment. The typical value of absorption for steel at the wavelength 1.06 ~ is

about 30%. Thus the value of absorbed energy flow is approximately 100 J/cm2. This value of

}C.

10 3

10- 1

THICKNESS OF THE HELT

(.\..

i \ i \ I \ f .

I \ /. '2\ ! /~-,\ ! 3 / '\

! .--. / \. /1 ""- I ", . I "\,

Y· '/'-. \' I r ""-. '~,

.1 ! "'-.'" '~" / i " .------. '-..

104 10 5 10 6 107

DENSITY Of ENERGY FLUX, W/cm2

10- 2 10. 3 10. 4 10- 5

108

10-G

PULSE DURA TJON. sec ~----------------------- ~~~~------------------------~

Fig. 16. ~xim~m values of me~t thickness ~roduced by laser pulse versus density of energy flux and pulse duratIOn (fixed energy denSity 100 J/cm ). Curve 1 - Ti; 2 - steel; 3 - Zr02.

191

energy input is assumed in Eqs. (1) for the action of laser pulses with different constant energy

flux densities qo and pulse durations ~t, but in all cases qoi' ~ti=l00 J/cm2 (for the one

dimensional model it is better to operate with energy density rather tha =1On with total energy).

The results this of calculation are presented in Fig. 16; for convenience two scales are used: one

for energy flux density and the other for the corresponding pulse duration.

For some technological applications, for example alloying from the gas or the liquid

phase, it is important to know the duration of the melt on the irradiated surface. The penetration

of admixtures from the gas phase occurs during this period. This time is equal to (tso}-tm), where

tsol is the time at the end of solidification (the disappearance of the liquid phase). The

calculations show that, as in the case of the melt thickness, there exists a maximum value of the

melt lifetime. In the case of pulse-periodic action, it is easier to find optimum value of duty cycle

when the characteristic time is normalised with respect to the pulse duration. The extreme value

of melt lifetime normalised to the pulse duration (tso}-tm)/~t, is about 2 for steel and about 3

for titanium. As qo increases, the curves tend to unity because tm ",0 (tm - qo-2) and tsol "'~t.

5. Spiked structure of laser pulses

For the above calculations, the real shape of laser pulse corresponding to the case of

chaotic (free) oscillation is approximated by a smooth curve. As a rule, the same approximation

is made for typical numerical simulations of heat processes of laser action. In practice, however,

a radiation pulse with duration about 10-3 s consists of a sequence of spikes with duration about

10-6 s and a duty ratio (pulse period to pulse duration ratio) of approximately 2 to 5 [2,48].

Also, the amplitudes of the spikes are not constant. In a spiked pulse with chaotic generation,

one can characterize the front by a certain rise rate, and observe some regularity in amplitude

decrease towards the end of the pulse. In practice, it is important to understand when it is

necessary to consider this real spike structure and when it appears possible to approximate its

shape by a smooth curve. In part, this is a question of the reproducibility of laser pulse action in

cases of chaotic generation.

5.1. LASER PULSES WITH CHAOTIC GENERATION

The results of a numerical simulation of spiked pulsed laser action are presented in Fig.17.

The model pulse consists of a base (which represents the average value of energy density flow)

192

plus pulse-periodic oscillations, represented by parabolic functions, that are added to the base.

These pulse-periodic oscillations are used to model the spike structure. The pulse shape is

characterized by a more rapid rate of increase of the average (base) values of energy density

flux, then by a decrease of average values near the pulse end. For the period in the middle of

the pulse its average value is constant. All these features are typical of the chaotically generated

pulse. The fine structure of the model pulse consists of 20 spikes, whose energy density flux

values oscillate in the range 2.6.106 W cm-2 , which is twice as small as the base value in the

middle of the pulse. The period between spikes is l.8·10-5 s, and the end of laser pulse is at

7.5

N I

E 5.0 u

~ 11>

0 2.5

0'

2 3 3.00

2.50 25

E 2.00 20

l- E ...... IS ::I. I- 1.50

10 In

1.00 <l

0.50 5

2 3 o -4

t [ 10 51

50

I en E 25 U

>

2 3 -4

t [ 10 51

Fig. 17. Simulation of heat processes due to a spiked laser pulse with chaotic generation. Time dependence of: density of energy flow (1), surface temperature (2), melt thickness (3), melting (4) and evaporation (5) front velocities.

193

3.6·10-4 s. Of course, this model pulse differs from the real one, for two reasons. On one hand,

the amplitude of model oscillations is less than that of a real laser pulse. On the other hand, the

duration of the model pulse is much less than the few milliseconds which is typical of

chaotic generation. These two factors compensate for one another, because the more spikes one

has, the smaller the oscillation amplitude. This results in a decreased influence of spike structure

on heat process dynamics. Also, it is possible to consider this model structure of laser pulse as

representing not individual spikes, but groups of spikes. In practice, the largest nonmonotone

changes of energy flux during pulsed laser action are caused by groups of spikes.

In this model, the metal subjected to laser irradiation is a steel slab, with thickness 1 mm,

with the following heat transfer constants: a = 5.5 10 -6 m2/s; A. = 29 W/mK; Lm = 2.7 . 105

J/Kg; Lv = 7.1.106 J/Kg; Tm = 1808 K; Tv = 2750 K; To = 300 K; e = 0.42.

It is possible to note from Fig. 17 that weak oscillations of surface temperature are not

accompanied by melt thickness oscillations. The melt depth monotonically increases during the

energy pulse. This corresponds to a melting velocity behaviour in which oscillations strongly

decrease in time until their practical disappearance. The oscillations of evaporation front

velocity, which is determined by the surface temperature, copy the structure of energy flow. It is

possible to infer that, for the assumed model shape of the laser pulse, its spike structure is not

significant. For the temperature distribution in liquid and solid phases, and for the

movement of phase fronts, the most essential factor is the average energy input, which is

determined mainly by the base values but not by the spike structure. As the melt thickness

increases and overheating (over Tm) occurs, the influence of any regular or irregular

oscillations of energy flux with varying amplitude is not very important for the movement of the

melting boundary, if the amount of energy in a single spike (or in a group of spikes, or in the

general case the variation of energy input) is much smaller than the energy stored in the melt.

Heat accumulation in the melt depends on a number of different factors: density of energy

flux, duration of action, heat transfer coefficients, surface heat losses, etc. As a rule, at the

beginning of laser pulse action, when the melt just appears and its overheating is comparatively

small, the influence of spike structure on heat processes is more significant than in the

middle of pulse action. It is easy to see this in the decrease of the amplitude of oscillations of

melting velocity during pulse action.

To verify the foregoing conclusions, oscillations with the same amplitude were considered

for a pulse which had a base amplitude one-half that of the previous case. In this case the

oscillation amplitude was equal to the maximum base value, i.e., 2.6 .105 W cm-2. The

behaviour of surface temperature, melt thickness and phase boundary velocities was

practically the same as before. There are weak oscillations of surface temperature, a decreasing

194

amplitude of melting velocity oscillations, and a monotonic increase in melt depth. The

quantitative differences are as follows. First, the oscillation amplitude of the melting velocity

is twice as large as in Fig. 17, but the values of the velocity are always positive, so that the melt

thickness never decreases. Also, the value of the evaporation front velocity is approximately one­

half the velocity of the previous case.

When we consider pulses with the same base value as in Fig. 17, but with spikes which are

two times larger (5.2 . 105 W cm-2), we obtain practically the same results. Increasing the

absolute energy density leads to an increase in the evaporation velocity, whose maximum values

nearly reach the melting velocity curve. During almost all the pulse action the oscillations of

melting and evaporation front velocities are out of phase, in some cases opposite in phase. This

is the effect of the lag time between the absorption of energy on the surface of melt (which

determines the surface temperature variation and therefore the evaporation rate) and its

propagation through the liquid phase up to the melting front.

The analysis reveals the influence of the ratio of spike amplitude to the base value. The

larger the ratio, the longer the period of melt thickness oscillations caused by the spike

structure. In other words, the larger the ratio, the stronger the influence of spike structure on heat

process dynamics. From the experimental (technological) point of view one of the "dangerous"

situations can arise when melting appears during the final period of laser pulse action. If the

amount of heat accumulated in the melt is comparable with the energy per single spike, then

as a consequence of chaotic spikes, the results of laser treatment may be non-reproducible. The conditions of this regime can be written in the fonn (t-tm)/~t« 1 (tm :=:; t :=:; ~t, where ~t = pulse

duration, tm = starting time for melting). For the action of constant energy density flow on a

semi-infinite body tm=(Tm -TO)2 ).21t/4a(Aq)2, where a = temperature diffusivity, A =

absorptivity of radiation.

For a comparatively small energy input, solidification of the melt layer can begin during

pulse action (as was discussed above, see Figs. 10-13). As a result, the melt thickness and its

overheating decrease at the end of laser action, and the influence of spike structure on the

dynamics of heat processes will increase. Later, it was noted in pyrometric measurements of

surface temperature that strong oscillations were observed mainly during the final period of pulse

action [15]. The conditions of this (second) regime can be written in the fonn ts<At,where ts =

starting time for solidification dS2(ts)/dt=O. The above mentioned conclusions mainly concern

relatively low values of energy density flow, when evaporation is weak.

Let us consider the influence of higher values of energy density flow on heat processes

caused by the spike structure of laser pulses. Results of numerical simulation of laser pulse

action with the model spike structure are presented in Fig. 18 (qmax= 2.106 W cm-2). The

195

oscillation amplitude coincides with the middle base value. In this case it is possible to see weak

oscillations of the melt thickness with a frequency which is close to the frequency of energy

flow. The reason for these oscillations is the increase of evaporation front velocity up to

values larger than the corresponding ones of melting boundary velocity. The decrease of the

oscillation amplitude of surface temperature in the middle of pulse action is the result of the

limitation of maximum surface temperature values by evaporation, and the increase in its

minimum values by heat accumulation near the irradiated surface. As a consequence, the largest

oscillation amplitude of surface temperature is realised at the beginning and at the end of laser

pulse action.

20 N I

E u ;t

10 OJ> 0

a

2 3

240

2.00 20

f-E 160 15 E

"-f- ::l

1.20 10 If)

0.60 <I

040 5

0 o -4 2 3 t [ 10 51

50

'(I)

E U 25

>

0

-4 2 3 t [ 10 sl

Fig. 18. Simulation of heat processes due to the action of spiked laser pulses with chaotic generation. Time dependence of: density of energy flow (I), surface temperature (2), melt thickness (3), melting (4) and evaporation (5) front velocities.

196

The significant influence of evaporation on heat process evolution demonstrates that not

only the ratio of the spike amplitude to the base value (in other words to the average energy

input, which determines the heat accumulation near the irradiated surface) is significant, but

also the energy content of a single spike (or a group of spikes), is important because it can cause

extensive evaporation. This phenomenon may be of great importance for melt movement and its

removal from the molten pool due to reactive pressure. The periodic action on the melt by the

pressure pulses induced by surface evaporation may be the reason for its removal from the zone

of laser action. The chaotic spike structure of a laser pulse is not significant for melting front

movement if the average energy input is not significantly changed, but in the case of extensive

evaporation caused by a single spike (or a group of spikes), the spike structure of the laser pulse

may lead to uncontrolled movement (removal) of the melt.

5.2. LASER PULSE WITH ORDERED GENERATION

Another type of laser pulse spike structure corresponds to so-called ordered generation.

This pulse is a sequence of individual spikes with a duty ratio around 2 to 5 and a total

duration of about 10-3 s, similar to the case of chaotic generation. However, the amplitude of the

individual spikes and the time interval between successive spikes remains constant during a

considerable part of the pulse [2, 48]. The results of heat process simulation are presented in

Fig. 19. The real spike structure of a laser pulse with ordered generation is approximated by a

periodic energy flow (the period is approximately 2.10-5 s). The main difference from the

previous case is the decrease of the energy flux to zero during each period of oscillation. For the

assumed pulse structure 80% of the energy input per period corresponds to an interval of only

7.10-6 s (the duty cycle equals 2.5). Twenty energy spikes have been assumed during the period

of laser pulse action. We should also mention that we have taken the energy input for the first

and the last spike to be one-half as much as for all the rest. The calculations were continued up to

time 3.64 ·10-4 s, so that after the end of laser pulse one more time interval equal to the period of

energy flow oscillations was considered. For this simulation, the irradiated metal (steel) has the

same properties as in the previous case.

The dynamics of the surface temperature are characterized by oscillatory behaviour similar

to the structure of the laser pulse, except when the curve crosses the melting point temperature.

The evolution of melt thickness in time is characterized by a different behaviour at the

beginning of pulse action and at the end (the last three spikes). The strong oscillations (first

stage) are converted to a practically linear dependence of melt thickness (second stage). For the

initial period, during the action of the seventh, eighth and ninth pulses, melting is followed by

197

complete solidification. For the final stage, during the action of the last spikes, the increase of

melt thickness vs. time is determined mainly by the average energy input and depends only

weakly on the spike structure of the laser pulse. The dependence of melting (solidification)

velocity vs. time reveals these regularities more distinctly. The initial increase of amplitude of

the velocity oscillations during the first stage, is followed by their decrease during the second

stage. Since the average energy input and surface temperature are not large enough for extensive

evaporation, the largest velocity values are less than 1 cm/s.

N I

E 42 u

1Il~ o

VVVvvvuVvv vv Jvvvv 2 3

-4 t [10 sl

1.8 r---------------------,

-I

1.6

1.4

E 1.2

t:: 1.0

I- 0.8

0.6

0.4

0.2

2

5

E ::1.

0.0 1---____ tA-u...~t..._ _ __+-----__+----'O -4

2 3 t [10 sl

III

E u

2 3 -4 t [10 sl

Fig. 19. Simulation of heat processes produced by a spiked laser pulse with regular generation. Time dependence of: density of energy flow (I), surface temperature (2), melt thickness (3), melting (4) and evaporation (5) front velocities.

198

Considering the same structure of laser pulse as in Fig. 19, but with qmax=106 W cm-2,

it is possible to see oscillations of the melting (solidification) velocity, as if they have been

truncated from the left, as compared with the plot in Fig. 19. In this case, the increase in the

absolute values of solidification velocity with the number of pulses, (which corresponds to the

first stage) disappears and the largest absolute values are reached during the early melting stage.

The amplitude of oscillations of melting velocity decreases down to a value of about 2.5 em/s

and during some periods the melt thickness increases practically monotonically. Upon an

increase in the maximum values of surface temperature up to 1.9 Tm, the velocity of

evaporation increases also, and its maximum value (11 cm/s) becomes higher than the

corresponding value of melting velocity, (Le., about 8 cm/s). Because of this, the oscillations in

melt thickness reappear.

In contrast, decreasing qmax values lead to a smooth monotonic increase of oscillation

amplitude of the melting (solidification) velocity. This is later followed by a smooth monotonic

decrease. As a result, the plot of velocity is quasi-symmetric with respect to the oscillations with

the largest amplitude.

Let us briefly analyze the problem of stability of the laser irradiation. Usually this problem

concerning pulsed laser action (or pulse - periodic action) is considered relative to the parameters

of the entire pulse action (energy per pulse, pulse duration); or relative to the action of a number

of pulses, for example, 100 - 1000 pulses (for the pulse- periodic case). Using this approach, it is

not possible to model accurately the sharp unstable changes of energy flux during comparatively

short time periods, especially when the variation of total energy input is on the order of a few

percent. This situation has been analyzed for the case of spiked laser pulses corresponding to

ordered generation. The results are presented in Fig. 20 (material - steel). Previously, different

pulse shapes were considered, but all these cases dealt with fairly regular structures (or regular

variation of the structure). In the present case, the sharp increase of energy input during the

period of action of a single spike is considered. Before and after it, there are regular variations of

the energy flow. It is possible to see that variables such as the surface temperature and

evaporation rate (which is determined by surface temperature) maintain their regular oscillations

just after the end of an abrupt change of energy flux. This the results from comparatively short

transient periods. The relaxation of melting front velocity also occupies all of the next period of

energy flow oscillations. The main result is a sharp decrease in melt thickness because of the

increasing evaporation rate. Later, this decrease is partially compensated by an increase in

melting velocity (as a result of melt thickness decrease). It seems that after increasing the pulse

duration up to nearly a millisecond. it will not be possible to distinguish the results of unstable

change of energy flow. This conclusion will be correct if the reactive pressure caused by

199

evaporation is not strong enough to remove the melt from the zone of action. On the basis of the

present analysis, it is possible to make a preliminary conclusion that it is necessary to control the

extreme deviation of energy flow from its average value, probably by joint adjustment of the

energy per pulse and pulse duration. Considering these two alternatives, the most dangerous is

the increase of energy flow, because its decrease (or temporary interruption) will not be

distinguished later.

1.5

'" 'E u

~ 1.0

10 0

0.5 C"

2 2.40

200

~ 1.60 IS E :1.

"- 1.20 l-

10 ~ 0.80

0.40 5

0 o -4 2 3 t [10 51

4

50

I en

E 0 u

>

-so

2 3 t [ 10 -4

51

Fig. 20. Simulation of inSlability during the action of a spiked laser pulse with ordered generation. Time dependence of: density of energy flow (1), surface temperature (2), melt thickness (3), melting (4) and evaporation (5) front velocities.

200

6. The two - dimensional melting-solidification problem

The proposed 2D model is an extension of the above mentioned ID model, but with some

simplifications. The principal simplification is that evaporation is assumed to be weak enough to

neglect the movement of the evaporation phase boundary, but enough strong to be taken into

account as a heat loss. Also, the free surface of the melt is assumed to be flat. From the physical

point of view this corresponds to the action of comparatively low energy density flux, when

surface evaporation is weak, the amount of evaporated mass is small, and the reactive pressure of

evaporation does not create too mach curvature in the free surface of the melt. In technological

applications this corresponds, for example, to pulse laser alloying (by solid state Nd:YAG -laser

sources, A. = 1.06 J.Ull) with energy density flow qo= 5·104 - 105 W cm-2 and corresponding

pulse duration 't= 1-10 ms. The mathematical model used, which assumes cylindrical symmetry,

can be written in the following form:

EQUATION OF HEAT CONDUCTION

T = T(r,z,t); t>O, O<z<L, r>O

BOUNDARYCONDnJONS

_A.(dT) =Qof(r)-Ql az z=O

V*exp(-T* / T(r, z = 0, t)) -<1 ( ) Ql 1/2 pLv + oer r, Z =0, t

T (r,z=O,t)

_A.(~T) =oe-r4(r,z=L,t); aZ z=L

( aT) =0' (lr r=O '

T(r=oo,z,t) = T(r,z=oo,t)=TO

INITIAL CONDnJON

T(r, z, t= 0) = TO

201

Here: r= radial coordinate, z - a coordinate nonnal to the surface, directed inside the material;

O(T-Tm)= delta function, which is used for the treatment of latent heat evolution; f(r) = radial

distribution of absorbed energy flux. The remaining nomenclature is the same as for the ID

model.

The computation of the heat conduction equation used discrete time intervals in a two

level backward Euler (fully implicit) scheme. For spatial intervals we employed the control­

volume finite-difference method [1,2]. A general iterative source-based method on a fixed grid

was used for the treatment of latent heat evolution [1,2]. To solve the linear algebraic equations,

the tridiagonal matrix algorithm was used. Convergence at a given time step was declared when

the residuals of enthalpy became less then 10-8.

Some of the results of heat process simulation under the action of a laser beam with

Gaussian spatial distribution q=qo exp(-kr2) on Titanium slab 1 mm thickness are presented in

Fig. 21 and 22. The duration of the laser pulse equals 7 ms, qo= 1.5·105 W/cm2, and k= 500

cm-2. The energy flux qO- ql which is expended on the processes of heating and melting of

metal decreases in time due to heat losses on evaporation. The surface temperature in the middle

of the heating zone T(r=O, z=O, t) reaches a saturation value of about 4000 K. During the initial

period oflaser action (curve zero, Fig. 21), both heat flux (a) and surface temperature (b) follow

the spatial distribution of the beam. Later, because of the significant nonlinearity of the process,

they start to differ substantially from Gaussian shape. The real spatial distribution of surface

temperature is important for a number of applications: for example, it detennines thennocapillary

convection of the melt. Thus, in order to simulate accurately the convective mass transfer in

pulse laser alloying, one need to take into account surface evaporation. After the end of laser

action, solidification begins at the edges of the pool, while melting continue at the bottom of the

pool. During the final stages of solidification, heat losses from the free surface of the melt induce

the solidification of the pool from its periphery towards the centre and from the surface down

into the bulk. The solidification front fonns a circle, and the last portion of the melt solidifies at

some depth inside the plate.

7. Conclusions

The dynamics of heat processes corresponding to the action of energy flows with a wide

range of parameters (continuous, pulses with different shapes, and laser pulses with spiked

structure) were analyzed with a model that considers both melting/solidification and evaporation.

202

Including evaporation in the model prevents the rise of temperature up to unrealistic values, and

makes it possible to consider comparatively high values of energy density flow. On the basis of

the behaviour of the phase front velocities, it is possible to analyze the dynamics of heat

processes to a very precise level.

1 N'

l.5l E ~ I

'b I - 1.0 -r->(

" I ~

01 .., 0.5 -,-.c:

e ::I I 01

I ... Q)

0. E ~ 2000 J_

! I

I I

I 293 I

I E f ~ \.

"0 I 0 50 -I 0. bll C '

~ 100-1-E . .c: \ - I fr 150 -: Q i Z

'If

(a)

0.25 0.50 0.75 1.0 1.25 1.50 1.75 radius (mm)

(b)

(e)

r

Fig. 21. Distributions of: (a) heat flux (qO- qt); (b) temperature alon~ the irradiated surface; (c) shape of melted pool. The curves correspond to the following times: 0 - 3·10- ; 1 - 0.1; 2 - 0.2; 3 - 0.5; 4 - 1; 5 -2; 6 - 4; 7 - 7 ms.

203

The principal physical results are the following: The simulation of the action of a constant

energy density flux shows that it is necessary to distinguish three different transient periods: the

shortest is characterized primarily by variations in s the surface temperature; the longest, by the

melt thickness (and melting front velocity); and an intermediate period, by variations in the

... ~ .,.. 8

>< ::s Ii: OJ <) .c

e -:!-

I bJl

-1.0

-2.0

-3.0

.§ 100

~ ~ 150 <)

Cl

radius (0101) T

0.25 0.50 0.75 1.00 1.25 1.50 1.75

T

T

Fig. 22. Distributions of: (a) heat flux (qO- qt); (b) temperature along the irradiated surface; (c) shape of melted pool. The curves correspond to the following times: 0 - 7; 1 - 8; 2 - 12; 3 - 14; 4 - 16; 5 - 18 ms.

204

evaporation front velocity. By varying the shape of the energy pulse it is possible to change the

dynamics of heat processes, and (usually) to a much smaller degree such resulting parameters as

the maximum propagation depth of the melting front, and the total thickness of the evaporated

layer. The results of the interaction of two phase fronts (two peaks of melting velocity) were

analysed for triangular and parabolic energy pulses.

In addition,the characteristic features of heat processes of pulsed laser action were discussed. It

was shown that in some cases, the larger the energy input, the smaller the temperature rise after

the end of the pulse. The existence of two different regimes in the dynamics of melting!

solidification front was shown. The results of high speed pyrometry measurements are in

agreement with the predictions of the present model. The influence of spiked laser pulses,

corresponding to chaotic and ordered generation, on the dynamic of heat processes was also

simulated. Finally, the problem of reproducibility of the results of pulsed laser action was

discussed.

To simulate the two-dimensional problem of melting-solidification in pulsed laser action,

one needs to take into account surface evaporation. This leads to a number of interesting results,

including a circular solidification front.

References

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2. Rykalin, N., Uglov, A., Zuev I., and Kokora, A. (1988), "Laser and Electron Beam Material Processing. Handbook", Moscow, Mir Publishers.

3. ECLAT' 90, (1990) Proceedings of the 3rd European Conference on Laser Treatment of Materials. Erlangen, Germany. Edited by H.W.Bergmann, R.Kupfer.

4. Uglov, A., Smurov, I., Lashin, A., and Guskov, A. (1992), "Modelling of Thermal Processes under Pulsed Laser Action on Metals", Moscow, Nauka.

5. Royere, C. (1981), 'Les applications du four solaire de 1000 kW du C.N.R.S. a'Odeillo'. Entropie, No. 97, p.147-160.

6. Ferriere, A., Flamant, G., Robert, F., Pekshev, P., Smurov, I., and Chystyi, A. (1990) 'Characterization of AS-sprayed and laser treated zirconia-based plasma sprayed coatings'. 1. de Physique, Colloque de Physique, Supplement au nO. 18, Tome 51, p.393-402.

7. Ferriere, A., Flamant, G., and Olalde, G. 'Heat treatment of thermal barriers using solar furnaces'. High Temperature Chemical Processes.

8. Draper, C.W., and Poate, I.M. (1985) 'Laser surface alloying'. International Metal Reviews. v.30, p.85-108.

9. Herziger, G., and Kreutz, E.W. (1986) Physica Scripta, T13, p.l39. 10. Uglov, A., Smurov, I., Tagirov, K., and Guskov, A. (1992), 'Simulation of unsteady-state

thermocapillary mass transfer for laser doping of metals'. lnt. 1. Heat and Mass Transfer, vol. 35, No.4, p. 783-793.

205

11. Smurov, I., Covelli, L., Tagirov, K., and Aksenov, L. (1992), 'Peculiarities of pulse laser alloying: Influence of spatial distribution of the beam'. J. Appl. Phys., vol. 71, No.7, p.3147-3158.

12. Smurov, I., Uglov, A., Krivonogov, Yu., Sturlese S., and Bartuli, C. (1992), 'Pulsed laser treatment of plasma-sprayed thermal barrier coatings: effect of pulse duration and energy input'. J. Materials Science, to be published.

13. Matteazzi, P., Miani, F., Smurov, I., and Lashin A. (1990) 'Modelling of structural transformation in laser glazing of metallic materials'. Proceedings of ECLAT90. p. 927-938.

14. Ready, J. (1978) "Industrial applications oflasers". Academic Press. 15. Rykalin, N., Uglov, A., Selishev, S., Smurov, I. et al. (1985) 'Measurement of free temperature

oscillations induced by strong energy fluxes acting on metals'. SOY. Phys. Dokl., v. 30, No.8, p. 724-725.

16. Uglov, A., Zavidey, V., Ermolayev, A., and Smurov, I. (1990) 'Equipment for the investigation of heat and optical surface phenomena during the action of concentrated energy flows'. Tepl. Vys. Temper., v. 28, No.4, p.786-792.

17. Selishev, S., Smimov, A., Toker, and G. (1989) 'Peculiarity of surface laser plasma forming in high pressure gases'. Tepl. Vys. Temper., vol. 27, No.3, pp.449-455.

18. Smurov, I., Smimov, A., Ferriere, A., and Fiamant, G. (1991) 'Surface laser plasmas in the processes of laser treatment of materials in high pressure gases'. Proceedings of the 10th International Symposium on Plasma Chemistry. Bochum, Germany.

19. Rosenthal, D. (1941) 'Mathematical theory of heat distribution during welding and cutting'. Weld. J., vol. 20, No.5.

20. Rykalin, N. (1951) "Calculations of Heat Processes of Welding". Mashgiz, Moscow. 21. Uglov, A., Smurov, I., and Andrianov, S.S. (1981) 'On calculation of the rate of unsteady

ablation'. Zhurnal Technich. Fiziki, vol. 51, No.9, p.l787-1792. 22. Uglov, A., Smurov, I., and Lokhov, Yu. N. (1981) 'Calculation of the ablation of a plate of finite

thickness'. SOY. Phys. Dokl., vol. 26, No.2, p. 231-233 23. Smurov, I. (1985) 'Heat processes in melting and ablation'. In "Action of concentrated energy

fluxes on materials". Moscow, Nauka. 24. Prakash, C., Samonds, M., and Singhal, A. (1987) 'A fixed grid numerical methodology for phase

change problems involving a moving heat source'. Int. J. Heat and Mass Transfer, vol. 30. No. 12, p. 2690-2694.

25. White, R. (1982) 'A numerical solution of the enthalpy formulation of the Stefan problem'. SIAM J. Numerical Anal., vol. 19, No.6, p. 1158-1172.

26. Lynch, D., and Sullivan, J. (1981) 'Continuously deforming finite elements for the solution of parabolic problems, with and without phase change'. Int. J. for Numerical Methods in Engineering, vol. 17, p. 81-96.

27. Dalhuijsen, A., and Segal, A. (1986) 'Comparison of finite element techniques for solidification problems'. Int. J. for Numerical Methods in Engineering, vol. 23. p. 1807-1829.

28. Uglov, A., Smurov, I., and Guskov, A. (1985) 'On the melting of metals by concentrated energy fluxes'. Fizika i Khimiya Obrabotki Materialov, No.3, p.3-8.

29. Anisimov, S.I., Bonch-Bruevich, A.M., Eliashevich, M.A. et al. (1966) 'Action of power light flows on metals'. Zhurnal Tekhnich. Fiziki, vo1.36, No.7, p.1273-1284.

30. Anisimov, S.1. (1%8) 'Evaporation of light absorbing metal'. Teplofizika Vysokikh Temperatur, vol.6, No.1, p.116-120.

31. Anisimov, S.I., Imas, Va. A., and Romanov, G.S., et al. (1970) "Interaction of High-Power Radiation with Metals". Moscow, Nauka.

32. Majukin, V.I., and Pestriakova, G.A. (1984) 'Mathematical modelling of surface evaporation processes by laser irradiation'. Dok!. AN SSSR, vol. 278, No.4, p.843-847.

33. Majukin, V.I. (1985) 'Kinetics of metal surface evaporation by laser irradiation'. In the book: "Action of concentrated energy fluxes on materials". Moscow, Nauka.

206

34. Uglov, A., Smurov, I., and Lashin, A. (1989) 'Numerical simulation of phase change interface movement in pulsed laser action'. Teplofizika Vysokikh Temperatur, vo1.27, No.1, p.87-93.

35. Smurov, I., and Lashin, A. ( 1989) Thermo-physical process modelling of pulsed energy flow action on metal plates'. In "Physics - Chemical Processes in Materials Treatment by Concentrated Energy Flows", Moscow, Nauka.

36. Lashin, A.M., Smurov, I.Yu., Uglov, A.A., MaUeazzi, P., and Tagliaferri, V. (1989) 'Heat processes in pulsed laser treatment of metallic materials'. International Journal of Heat and Technology, Vo1.7, No.2, p. 60-73.

37. Smurov, I.Yu., Uglov, A.A., Lashin, A.M., Matteazzi, P., Tagliaferri, V., and Covelli, L. (1991) 'Movement of phase boundaries of metals subjected to surface periodic energy pulsed'. J. Appl. Phys., vol. 69, No.12, p. 8031-8036.

38. Smurov, I.Yu., Uglov, A.A., Lashin, A.M., Matteazzi, P., Covelli, L., and Tagliaferri, V. (1991) 'Modelling of pulse-periodic energy flow action on metallic materials'. Int. J. of Heat and Mass Transfer, vol. 34, No. 4/5, p.961-971.

39. Rykalin, N.N., Uglov, A.A., and Smurov, I.Yu. (1982) 'Nonlinearities of laser heating of metals'. Sov. Phys. Dokl., vol. 27, No. 11, p. 970-972.

40. Rykalin, N.N., Uglov, A.A., Smurov, I.Yu., and Volkov, A.A. (1984) 'Laser heating of a metal in an oxidizing atmosphere'. Sov. Phys. Dokl., vol. 29, No.8, p. 686-688.

41. Kar A., Chan C.L., Mazumder J. (1992) 'Comparative studies on nonlinear hyperbolic and parabolic heat conduction for various boundary conditions: Analytic and numerical solutions J. Heat Transfer, vol. 114, No.1, p.14-20.

42. Kar, A., and Mazumder, J. (1990) Two- dimensional model for material damage due to melting and vaporization during laser irradiation'. J. Appl. Phys., vol. 68, No.8, p. 3884-3891.

43. Chan, C.L., and Mazumder, J. (1987) 'One- dimensional steady-state model for damage by vaporization and liquid expUlsion due to laser-material interaction', J.AppI.Phys., Vo1.62, No.l1, p.4579-4586.

44. Soga, T. (1986) 'A kinetic theory analysis of unsteady evaporation from a liquid surface with temperature change', J. of Physical Society of Japan, Vol.55, No.5, p.1556-1567.

45. Hastaoglu, M.A. (1986) 'A numerical solution to moving boundary problems - Application to melting and solidification', Int J. Heat and Mass Transfer, Vol. 29, No.3, p.495-499.

46. Benard, C., Gobin, D., and Zanoli, A .. (1986) Int. J. Heat and Mass Transfer, Vol. 29, p. 1669. 47. Blom, J.G., Sanz-Serna, J.M., and Verwer, J.G. (1988) Journal of Computational Physics,

Vo1.74, p. 191. 48. Ready, J.F. (1971) 'Effects of high power laser radiation'. New-York, Academic Press

RECENT DEVELOPMENTS TO INCREASE THE EFFICIENCY AND FLEXmILITY OF LASER MATERIALS PROCESSING

H. HUGEL, F. DAUSINGER University o/Stuttgart lnstitutfor Strahlwerkzeuge IFSW Pfaffenwaldring 43 7000 Stuttgart 80, Germany

ABSTRACT. Utilization of parallel oriented linear instead of circular polarization in cutting and welding of steel sheets yields an Increase in processing speed of up to 50 %. Normal orientation produces seam widths which are considerably larger than those obtained with any other polarization form. A device has been developed allowing one to control the polarization orientation at any point of a 3-D processing path. Welding of aluminium with 1 kW power is shown to be possible when a laser of high Deam quality is used.

1. Introduction and background

The high potential of the "beam tool" laser for industrial processing is based upon the fact

that energy can be transmitted via a massless electro-magnetic beam onto the workpiece, and that

intensity and interaction time are easily controlled. Essentially these two features yield a high

degree of flexibility both with respect to processing techniques and variety of materials. High

processing speeds and high quality - in particular, the precision which often reduces refmishing

or even makes it unnecessary - have made the laser an accepted production tool. High investment

and operating costs, however, stand in the way of a still broader substitution of classic

technologies by laser technology in industry.

Considerations of economic efficiency in laser materials processing lead to a demand for

rigorously cutting the investment cost of the beam source and of the complete installation which

undoubtedly is a right and important starting point. Nevertheless, to obtain higher oost­

effectiveness, it is also essential to take advantage of the unique characteristics of lasers in order

207

S. Martellucci et al. (eds.), Laser Applications for MechanicaL Industry, 207-218. © 1993 KLuwer Academic PubLishers.

208

to increase process efficiency and flexibility.

The intensity is the most decisive parameter for the interaction processes on the

workpiece. To achieve high enough intensity - e.g. for cutting highly reflecting materials or

welding aluminium -. the usual approach is to increase the power level of the beam source. TIle

recent availability of multikilowatt C02-lasers with high beam qUality is an important step

towards higher efficiency. since the improved focussability reduces the demand for laser power

to obtain a particular intensity.

The beam quality of a laser beam is characterized by dL 0/4. dL being the waist diameter

and 0 the (full) divergence angle. To measure and to compare this property of different sources

and. what is even more important. of beams having passed through the beam guiding and

shaping systems. a quality number k is introduced describing the ratio of the theoretically best

possible beam quality (which has a value of A. /1t) to the measured one. Typical numbers for

C02-lasers range between 0.7 and 0.2. They fall off with increasing power level of the source

due to higher modes and interaction effects between beam and optical elements. such as optical

deformation and diffraction. Since the focus diameter achievable with a focussing optic of focal

number F is given by df - F/k and the Rayleigh length (depth of focus) by ZR - F2/k. a direct

influence of beam quality on the cutting process is to be expected.

These relations indicate that. for realizing a certain value of df. the F-number can be made

the larger if k is larger. This is of considerable practical usefulness. In addition to reduced

demands for positioning. larger distances between focussing optics and workpiece facilitate the

use of clamping installations and allow one to apply new nozzle concepts. such as described

below.

In addition to the intensity. the polarization and the wavelength of the beam are deserving

of more attention. Both properties control the absorption and hence directly contribute to the

energy coupling and the process efficiency [11. Although positive effects have been reported in

the past [2. 3. 41. little practical use has been made of them. so far.

This chapter will address how attention to C02 laser beam quality and polarization can

improve the process efficiency and the flexibility of cutting and welding.

2. Cutting

Experience has shown that in laser cutting the kerf width is to a good approximation

5r---------------------------------~O~-,

108mm2

min 4

J:l x >

3

2

resonator:

V 5 mm, stable

o 2 mm, stable

o 1 mm, stable

X 2 mm, unstable t::,

t::, 1 mm, unstable t::, 00

0 0 X

~ ~

~# O~~----~---------L--------~------~ o 3 kW/mm 4

Fig.l. Normalized plot of cutting performance (conical nozzle with N2, mild steel).

209

proportional to the focus diameter of the focussed beam. For this reason, it directly results from

energy balance that at the same laser power the cutting velocity will be the higher the better the

beam is focussed. To examine this relationship, we plot the product of cutting velocity and kerf

width, as a function of the quotient of laser power and material thickness. Experimental data

obtained in the range from 200 W to 4200 W with stable resonators on three different systems, as

well as the results with an unstable resonator, show good correlation, see Fig. I [5].

Today, the gas flow necessary for melt expUlsion is normally produced in nozzles with a

conical or cylindrical exit. Such a flow field shows considerable variation in the axial direction

[6]; as a consequence, the distance from nozzle exit to workpiece can hardly be varied and, in

general, it is very short. Supersonic nozzles, however, supply gas jets having constant conditions

over long distances. Coaxial Laval nozzles have the disadvantage that their smallest cross-section

has to be adapted to the converging laser beam and therefore must be relatively large, leading to

higher gas consumption. Extra-axial configurations avoid this adverse effect and, moreover, they

do not require a window made of ZnSe or other transparent material to build up the pressure.

Their advantage of longer distance and wider working range can be used for cutting jobs,

such as deburring and trimming of forged or casted components parts, or for other applications

which have not been possible up to now, for reasons of inaccessability. Figure 2 demonstrates

the large working range of such a nozzle [7].

In addition to its advantages of accessability and tolerance, a Laval nozzle can improve the

achievable cutting velocity, as well, as may be seen in Fig. 3 [8]. The results further underline

210

'0 CD CD C. m 0> C ~

:; o X al E

25.-------------------------------------,

m/mln working range:

'0 CD CD C. m E ::J

E x al E

5

o conical 5 bar

L::. Laval 15 bar

Laval 5 bar

oL-----------L-----------~----------~ o 5 10 mm 15

distance nozzle - work piece

Fig.2. Demonstration of the large working range of an extra-axial double Laval nozzle (2 mm mild steel, PL = 3.8 kW, N2)'

6

m/mln

5

4

3

2

0 0

pol.: nozzle: * parallel. Laval

o circular, Laval

L::. parallel, conical

o circular, conical

X normal, conical

.. ' ..... ~:::::: .. X

........ o:<.>c 0 ......... 'X"

..... ~g<:.:x······ 200 400 600 600 W 1000

laser power -

Fig.3. Effect of polarization and nozzle design on cutting speed (1.5 mm mild steel, N2 ,k = 0.7 ).

gnear Xo1arization .Q£ientor (LlPOR)

"'"

LIPOR­software

211

FigA. Scheme of the linear-polarization-orientor adapted to a robot with external beam-guiding system.

the favorable influence of parallel oriented linear polarization. In order to take advantage of this

effect for plane and three-dimensional contour cutting and welding of workpieces, the plane of

polarization must be tangentially oriented at each point of the path.

At the IFSW (Institut fiir Strahlwerkzeuge), the facts discussed above led to the

development of a device for the controlled and systematic use of linear polarization in 3-

dimensional materials processing. Its main parts are a beam rotating unit consisting of three

mirrors rotating around the axis of the delivered beam, and the software allowing one to control

this unit as an additional NC-axis [9]. Fig. 4 shows the linear-polarization-orientor (LIPOR) in

its realization at the IFSW laser robot system. The LIPOR allows free choice of any desired

angle between polarization and working direction in any path. It should be emphasized that the

LIPOR works with all types of machinery, be they robots or cardanic moved devices.

3. Welding of steel sheet

The results to be discussed were achieved using the above-mentioned device together with

a 5 kW laser. The economic potential of the selective utilization of linear polarization is shown

212

"0 <D <D Q. (J)

E OJ E x a:l E

.r: 0. <D "0

C o

~ W C <D Q.

40

m/mln

30

20

10

0

rn:m parallel (3950 W) § circular (3800 W)

~ 45 degree (3950 W)WW normal (3950 W)

~ § ~ § ~ ~ ~ ~ E

cutting df : 190 Jlm

polarization

f=::

E= 1=

through-weld d f : 190 Jlm df : 360 Jlm

Fig.S. Maximum speed for cutting and through-welding of 1 mm mild steel plates dependent on polarization (cutting gas: N2' welding gas: Ar).

2,5 ~--------------------.

mm

0,5

gas: argon

o parallel (3900 W) * circular (3550 W)

<) 45 degree (4000 W) + normal (3850 W)

0.0 L-__ -'-__ -----' ___ --'-__ ----'-__ -.l

20 40 60 80 100 J/mm 120 line energy

Fig.6. Penetration depth in overlap-welding obtained with different polarization forms. Values of 2 mm correspond to through-welding.

213

in Fig. 5. Compared with circular polarization, for both processes, cutting and throughwelding of

steel plates, a remarkable increase in maximum processing velocity was obtained ranging

between 20 % and 50 %.

Some more detailed infonnation on the contribution of polarization to process efficiency is

found in Figs. 6 and 7. The values of penetration depth of overlap welds for two I mm thick

mild steel plates (St 14) as a function of the line energy (the corresponding velocity values range

up to 13 mlmin) are presented in Fig. 6. To obtain a given depth, the lowes line energy (highest

velocity) is required in the case or parallel oriented linear polarization and the highest for a

nonnal orientation. The results of circular polarization and those of a 45°-inclination of linear

polarization lie in between.

As a characteristic figure for the process efficiency, the ratio of the amount of heat

required to melt the seam volume and the energy delivered to the workpiece is plotted in Fig. 7.

With increasing line energy, each curve corresponding to a certain kind of polarization

approaches a maximum that occurs just before through-welding takes place (see also Fig. 6). A

further increase in line energy then necessarily leads to a decrease of process efficiency because

beam energy is transmitted through the keyhole. As we have already expected from Fig. 6, the

highest process efficiency is obtained with parallel and the lowest with nonnal orientation of

linearly polarized laser light.

An additional advantage of the controlled utilization of linear polarization is a

considerable increase in process flexibility for both overlap and butt-joint welding.

>. 0 c CD "0 ;;::

m III III CD 0 e Cl.

30

%

25

20

15

10

5

/ df :O.36mm

/ ~........ . gas: argon

~~,:,::::.i.:l·ct o parallel (3900 W)

* circular (3550 W)

o 45 degree (4000 W)

+ normal (3850 W) oL-____ -L ______ L-____ ~ ______ ~ ____ _J

20 40 60 80 lOa J/mm 120 line energy-

Fig. 7. Process efficiency in overlap-welding. The data correspond to those in Fig. 6.

214

CD c III a. .... c :Q, E .c. '5 ~ E III CD C7.l

1,0

mm

0,8

0,8 1 i

0,4

0,2

0,0 2

.'

....... parallel, d .-360,.tm

- normal, d. "190 pm

- parallel, d ."190pm

6 10 14 18 m/rnln 22 welding speed -

Fig.S. Demonstration of process flexibility through appropriate orientation of linear polarization in overlap-welding (2 x 1 mm mild steel, PL = 3.6 kW, AT).

In the case of overlap welding this is illustrated by Fig. 8, representing seam widths

measured in the joint plane that were obtained for different focus diameters and polarization

orientations. Since the strength of an overlap weld is mainly detennined by this value, it might

be necessary for a certain workpiece to increase it by changing the process parameters. A

common way to enhance the seam width is to increase the focus diameter. The comparison of the

regimes covered by df = 0,36 mm and df = 0,19 mm (with parallel polarization) shows that this

measure is successful, however, at the expense of reduced velocity and efficiency. Using nonnal

polarization at df= 0,19 mm, on the other hand, even broader widths are obtained at 40 % higher

velocity. For this reason, it can be concluded that perpendicularly oriented linear polarization

should be used first to obtain economical laser welding with the required broad seam widths.

Subsequently, and only if the achieved seam widths are still insufficient, an increase in focus

diameter can be adopted taken as a second step.

Butt-joint welding of sheet metals in the thickness range of about 1 mm generally make

high demands on path accuracy. To reduce the consequences of displacements and of different

joint gap widths for the seam quality (and hence for process reliability), a perpendicuiar

orientation of the polarization again provides an advantage. As seen from Fig. 9, broader joint

widths by almost a factor of two can be realized if the polarization is chosen nonnal to the

welding direction instead of parallel to it. The tolerance of this measure is underlined by the

almost unchanged width obtained within the variation of the velocity from 5 to 7 m/min and of

the focal position in a range of ±l mm.

215

0,7 mm X 0 0,6 ~ normal

0,5

.c ~

0 i3 0,4 0 t::.

0 ~ ~ X parallel E X X Ilk X en 0,3 CD (1)

E 0,2 ::l

0 5 m/mln E t::. 6 m/mln C E 0,1 X 7 m/mln

0 -1,5 -1 -<l,S 0 0,5 1,5 mm 2

focus position ---

Fig.9. Minimum seam widths in "utt-welding with different orientation of polarization ( 1 mm mild steel, l1. ;;;; 3.6 kW, df = 0.36 mm, Ar ).

4. Welding of aluminium

Due to the higher reflectivity and heat conductivity of aluminium as compared with steel,

higher intensity values are required to initiate the deep welding effect, i.e. to produce deep and

slender seams. The heat losses by conduction and the large difference between vaporization and

melting temperature in aluminium alloys tend to considerably broaden the seam widths. Hence,

an even more pronounced effect of beam quality might be expected in welding of aluminium.

The threshold intensity for deep penetration welding as well as the line energy to gain a

certain depth also depends on the chemical composition of the alloy, as indicated by Fig. 10.

lbi.s can be explained by the effect of different vapor pressures in the keyhole [10]. Since the

vapor pressure from magnesium is much higher than that of aluminium, increasing the

magnesium content results in sufficient pressure to establish a stable keyhole at a lower

temperature than for pure aluminium. This has been confirmed by other groups as well [11].

At otherwise constant parameters the penetration depth increases with beam quality as is

seen from Fig. 11. In addition, more slender seams can be achieved with higher beam quality,

especially in the range of low line energy, see Fig. 12. The data obtained with merely l.l kW

power but a quality number of 0.7 at the workpiece appear very encouraging.

So far, the results presented above cannot be generalized because it is not only the

processing speed or efficiency but, at the same time, the processing quality, as well, that counts.

216

3

mm

2 .c a. CD "0 c 0

~ CD C CD C.

0 0

Mg-content, vapor pressure t

20 40 60

¢ AIMg5Mn

t:,. AIOuMg2

'V AIMgSll

80 J/mm 100

line energy -

Fig. 10. Penetration depth depending on line energy and composition of the aluminium alloy (PL = 1.1 kW, F = 8, k = 0.7 ).

B

mm

5

4 .c a. CD '0 3 c 0

~ 2 a; c CD C.

0 0 2

beam quality:

OK· 0,3

* K· 0,18

¢ K· 0,11

AIMgSI1 p. 3,6 kW

4 6 8 10 m/mln 12

welding speed -

Fig. I 1. Effect of beam quality on processing data in welding of aluminium (AI Mg Si 1, PL = 3.6 kW, He ).

217

4

mm 0 K • 0,18 p. 3,6 kW

0 K • 0,7 P • 1,1 kW 3

0 K· 0,3 p. 3,6 kW

.<:: 0 0

D j 0

0 2 0

E 0 0 <0 0 0 0 m

0 000 0 '" 0 C <Il

0 08 0

m E 0

° ° 25 50 75 100 J/mm 125

line energy __

Fig.12. Effect of beam quality on the seam shape (AI Mg Si 1, F:=4at 3.6 kW,F:=8 at 1.1 kw, He).

The appearance of pores and hot cracks depend on the alloy and the process parameters.

Consequently, further investigations are required.

5. Conclusions

The recently obtained results presented here show that optimization of beam properties,

such as beam quality and polarization, can yield an increase in process efficiency and flexibility

and hence might lead to an even higher acceptance of laser materials processing in industry. In

particular, the controlled orientation of linear polarization can be regarded as an additional

process parameter to be used for influencing process efficiency, flexibility and reliability.

References

1. Dausinger, F. (1990): 'Lasers with different wavelength· implications for various applications', in Bergmann, H.W.(ed.), Proc. 3rd European Conf. on Laser Treatment of Materials, 1990, Sprechsaal Publ., Coburg ,voU, p.l.

2. Olson, F.O. (1980): 'Cutting with polarized laser beams', DVS-Berichte 63, p. 197. 3. Behler, K., Beyer, E., Schulz, W., Wolf, N. and Welsing, O. (1988): 'Accommodation of

polarization and shaping in laser beam welding', in Hiigel, H. (ed.), Proc. 5th Int. Conf. on Lasers in Manufacturing (LIM 5), IPS Pub!., Kempston , p. 187.

218

4. Dausinger, F.and Rudlaff, T. (1987): 'Novel transformation hardening technique exploiting Brewster absorption', in Proc. In!. Conf. on Laser Advanced Materials Processing (LAMP '87), High Temperature Society of Japan, Osaka , p. 323.

5. Schreiner-Mohr, U., Dausinger, F.and Wiedmaier, M. (1990): 'Trennen mit C02-Hochleistungslasem - Einsatz instabiler Resonatoren', Laser und Optoelektronik 22 n° 6, p. 51.

6. Fieret, J and, Ward, B.A. (1986): 'Circular and non-circular nozzle exits for supersonic gas jet assist in C02-laser cutting', in Quenzer, A. (ed.), Proc. 3rd Int.Conf. on Lasers in Manufacturing (LIM-3), IPS Publ., Kempston , p. 45.

7. Edler, R. and Berger, P. (1991): 'Vorstellung eines neuen Dusenkonzepts zum Lasertrennen', Laser und Optoelektronik 23 , nO 5, p. 55.

8. Schreiner-Mohr, U., Dausinger, F. and Hugel, H. (1992): 'New aspects of cutting with C02-lasers', Proc. ofICALEO '91, San Jose, USA.

9. Dausinger, F. and Wahl, R. (1991): " Robotergefiihrtes C02 Laserstrahlschweipen im Dunnblechbereich ", DVS-Bereichte 135, p.7.

10. Rapp, J., Glumann, C., Dausinger, F. and Hugel, H. (1992): 'Lap welding of AlMg5Mn-alloy plates with C02-lasers', ISATA 25.

11. Cieslak, M.J. and Tuerschbach, P.W. (1988): 'On the weldability, composition and hardness of pulsed and continuous Nd:YAG laser welds in Aluminium alloys 6061, 5456, 5086', Metall. Transactions B, 19B , S. 319-329.

LASER WELDING TECHNOLOGY FOR JOINING DIFFERENT SHEET METALS FOR ONE PIECE STAMPING

KAZUOAZUMA Toyota Motor Europe Marketing & Engineering Co. Production Engineering Division Hoge Wei 33A-1930 Zaventem, Belgium

KlMIKAZU lKEMOTO Body Production Engineering Div. 1, Motomachi-Cho Toyota Aichi 471.Japan

ABSTRACT. The welding conditions for stamping have been evaluated with successful results in formability and strength. The laser welded blanks made by using this new technology have been applied to actual body components with satisfactory results in material cost savings, improved body accuracy, and decreased body weight..

The typical car body is made of more than 300 pieces, with different thickness and

different treatments, according to the requirements of the part's characteristics (Fig. I). Examples

of ceiling area are shown in Fig. 2. Fig. 3 shows the cost table of a car with body: - The material

cost is 48 %;-The welding cost is 12 %; - The stamping cost is 13 %; and - The ceiling cost of

the joining point is 3 %. Thus it is very important to increase the material yield and to decrease

the spot welding, the stamping dies and the joining point for white body cost reduction.

If we can stamp these separate panels to from a one piece panel efficiently, by using

integrated sheet blanks, we can reduce the costs of material, welding, stamping and ceiling.

Table I shows the advantages and disadvantages of the conventional methods for side- member

panels. If we choose the divided type, then the cleanness, press dies cost and accuracy are not as

good as with of the one-sheet type.

Our "Laser Welded Blank" which joins different kinds of sheets before stamping (as

shown in Fig 3) has overcome the disadvantages of conventional methods.

219

S. Martellucci et al. (eds.), Laser Applications for Mechanical Industry, 219-233. © 1993 Kluwer Academic Publishers.

220

Fig. 1. Body construction and examples of ceiling area.

@Cemng cost 3%

Fig.2. Cost table of White Body.

Table 1. Advantages integrated of side member using laser welded one sheet blanks.

Compared items Divided type One-sheet type

~":=:"J "./~---··-··l

f(J1n~f-\ ... / i i '1 '.

rOO~j (I :. • J Schematics ~! f i ,..' .

L..J ' .. J L •• __ / r -::J r~--" .. --"1 .~-"- .. -.--.. - .......

CD Cleanness bad X good 0 @ Number of dies 20 dies X 4 dies 0 ® Accuracy low X high 0 @ Material yield high (65%) 0 low (40%) X

® Material flexibility selectable 0 fixed X

@ Weight light 0 heavy X

221

Blanking L-___ vv_e_ld_m __ g ____ ~1 ~I ___ S_t_a_m_p_m_g ____ ~

c?[][) .. c==-=J

Fig. 3. Integrated side member using laser welded blanks

When welded sheet metals are used for stamping, the welding characteristics must satisfy

the requirements of joint shape, high strength and high productivity. Fig. 4 shows a comparison

of the available welding methods. It is impossible to satisfy all the requirements by using Hush

butt-, Mashseam-, Plasma- and Electronic beam welding methods. Only, TIG arc and C02 laser

welding methods have the possibility of satisfying all these requirements.

Fig. 5 shows the sectional view and the hardness distribution when shear cut edges are butt

welded using C02 laser welding and TIG arc welding methods. The differences between those

methods are the following: the C02 has a smaller hardened area and no area softened by heat,

unlike the TIG arc welded bead. This is very important for press forming, because the hardened

area decreases the formability and the softened area, permits too much extension and may cause

a partial crack. Therefore, we decided to use the C02 laser welding method.

~ 1- ~ ~ -.. o..n.D ""tn ot..". ~ ~

A.<(r.C) I!:. A- I!:. I!:. I!:. n..hbuK X I!:. A- 0 X

Masbs .... X I!:. A A X

PIa<m.o ® 0 I!:. I!:. A E1td.mnll: brim 0 0 0 X I!:. Olall"rr 0 0 0 0 0

Fig. 4. Sectional view of butt welding and comparison of welding methods.

222

0--0 Laser Welding /:::.._-/:::. TIG arc Welding

300 Hardened ilrCa Heat affected zone

, _.L:r--t:. 'l:r --~

Softened area

2.0 1.0 o 1.0 2.0 Distance from center of welding (nun)

Fig. 5. Sectional view of laser and TIG arc butt welding and welded bead hardness distribution.

We studied the relationship between the welded bead characteristics and fonnability by

using laser welding test machines. Fig. 6 shows the welding test machine. After several tests, we

were able to establish suitable specifications, the operating conditions and the specimen

characteristics of the test samples (see Fig.6). Sheet metals with a thickness between 0.8 and 1.0

mm are used for test specimens, because this range is widely used for body components.

Fig. 7 shows the results of the tensile strength test. It points out the relationship between

the welded bead characteristics and the strength ratio S/So (S = strength of welded sheet metal;

So = strength of raw sheet metal). When T/To is equal to or greater than 0.7, breakage will occur

on the raw sheet metal, indicating that the welded bead has a higher strength than the raw sheet

metal.

Mirror ),5 kW CO2 laser oscillator Material Sheet

. thickness Yield Tensile Hardness poin t st rength

CJI !

i 0\ ~ I ,

I '

:'vlode: Low.grade multi Beam d iameter : 28 mm

Lens foc us distance: 12i mm

assist gas: Argon

Imm l

SPCC 0.8

SPCC 1.0

SPFC.S I 0.8

i SPFClS 1.0

SPFC60 0.8

! SPFC60 1.0

Fig. 6. Welding test machine and specimen characteristies.

d';~'mml, I I\!,mm21 d-h- l

20.0 34.5 100

197 33 .• 100

.2 49.5 150

[ ,)1.9 .)8.9 ISO

. 8.2 62.0 190

.7.8 , 61.1 190

223

1.0 I Welded line

/ 0

rJl en 0 . .;:::: 0:: ..... 0.5

.J:;

"b1:, c: CJ ....

c:n Breakage at welded portion

(l 1 1

0.4 0.6 0.8 1.0

Thickness ratio T/To

Fig. 7. Relationship between thickness ratio and tensile strength.

We also considered, the relationship between the welded bead characteristics and

fonnability. In order to evaluate the stamping fonnability of the welded bead, the liquid bulge

test was adopted as shown in Fig. 8 and 9.

Fig 8. The liquid bulge test.

u o.~ x QJ

.~

:;

oil pressure :.§

. s D,:; "2 E ~

OJ C

Fig. 9. Test sample.

Material: spec Sheet thickness: 0.8 mm

• •

• Formable area

O.:l 0.0 07 o.~ 0'1 1.0

Thickness ratio T rr 0

Fig.lD. Relationship between thickness ratio and deformation limit strain.

224

0.4

u x (]) ,...

. :l

]; '"

"@ 0.2 "" .§ co ,:: c

..Q OJ 0

0

1\ ~ .

Raw ~ •• sheet .. metal

Formable

100

Sheet thickness: 0.8 mm

rITa? 0.8

•

300

Welded portion hardness Hv

Fig. 11. Relationship between hardness and deformation limit strain.

Fig. 10 shows the test results, where the deformation limit strain is indicated at each

thickness ratio Trro. The formable area is shown with oblique lines; the curve shows that exc

drops sharply when Trro is under 0.7. This means that Trro must be 0.8 or more for practical

application.

Fig. 11 shows that deformation limit strain exc' decreases as the hardness Hv increases.

Therefore, if we want to adapt this panel for stamping, we have to decrease Hv.

Fig. 12 shows the result of the bending fatigue test. The breakage occurs on the raw sheet

metal; 20 % and 30 % pre-strain sheet metals yield the same result. This indicates that the

welded bead is stronger than the raw sheet metal.

"f Raw material

! 25 /

I

~ Pre-strain ~

'''~ direction m 0

U 0 Of> ~ J ~

" ,r--<-, o Without Pre-strain

0

'" (' ,f' ~ I /-t'" 6. 20% Pre-stram ~

101 ~

~' o 30% Pre-strain if, ~) if, ~

01 ~ US

10' 10' 10- 10;

Number of repetitive cycles

Fig 12. Durability of bending fatigue.

1.0

0.8

<)

t:: . .3 il5 e ~

'" ~ u

~ m/mm

0.1 0.15

Sheet thickness: 0.8 mm

0.2

Welding is impossible

Clearance (mm!

Fig 13. Relationship between clearance of butt welding and thickness ratio.

225

Fig. 13 shows the relationship between the welding speed and the thickness ratio. When

the clearance and the welding speed increase then the thickness ratio will decrease, as shown

here. The welding was difficult when the clearance exceeded 0.2 mm. If we want to keep TrrO at

0.8 or larger, we have to decrease the clearance below 0.15 mm at a welding speed of 4 m/min.

Fig. 14 show the relationship between the specimen's carbon content and the hardness. The

hardness increases with the carbon content, and formability requires a carbon content lower than

0.04%.

Fig. 15 shows the relationship between the welding speed and the hardness for each

specimen . The hardness will increase when the welding speed increases.

Thickness: l.0 mm

Material Carbon Yield Tensile content point strenght

400 o

SPCC 0.04 % 19.7 kg/mm2 33.4 kg/mm2

SPFC 45 0.06 r. 41.9 kg/mm2 4B.9 kghnm2

SPFC 60 0.09% 47.8 kgimm2 61.1 kg/mm2 0.02 om 0.06 0.08

Carbon content (o/c)

Fig. 14. Relationship between carbon content and hardness, along with specimen characteristics.

226

;>

:r: 400 U) U) ·lJ

] :r:

300

T

SPCC®SPF~

• SPCC 0 SI'FC45 D.

~)-----------~C~)

°1 _L'.

SPCC®SI'CC

4

Welding speed (m/min)

Fig. 15. Relationship between welding speed and hardness .

-;: ...--:--

2: U) bead U)

Q ~ O.l5mm

I!J 300 c:

I Hv. max I "':l ... '" ..c:: n \!:

~.~ OJ .>:

2001 ;; down

100

TITo;;:;:; 0.8 1.0 0 1.0

Distance from center of welding (mm)

Fig. 16. Hardness distribution of bead; the necessary conditions of welded bead are also indicated.

SGAC 30[30 t = 1.0

SGMCF t = 1.0 ,

SGAC bU:f>I) t = 1.0

Fig. 17. Integrated side member blank.

-" Total weldmg length: 2600 mm Maximum welding ll;!u",h : 500 mm

spec 28 DY t =0.8

227

0.10

E E ."

.c ·iii

'or ] '0 ;>.

~

~ TOP )'fiODlE BOTTOM

Location of coll edge

Fig. 18. A,B,C,D, and E panels and accuracy of coil edge linearity.

In summary, the gap between two substrates should be smaller than 0.5 mm, to keep the

Trro above 0.8 and to decrease Hv as much as possible (see Figs. 16 ).

Fig. 17 shows the integrated side member blank of Toyota Lexus' side-member-outer­

panel. The total welding length is 2,600 mm. It joins 5 different sheets of blanks ABC D E. It is

very difficult to join. Therefore, we need four development techniques to realise it for practical

use.

,---------------_._-

1st Step .. ~I~\S' L..lf Sell 2nd Step (@@ C .. ,lmp)

3rd Step 11~i]@ Cr,mlp)

Fig. 19. Jig specification.

228

First, to minimize the gap we used the following procedure. To satisfy the requirements,

we used a trim coil edge for the A,B,C panels and we adapted special punching dies for the D,E

panels (Fig.18). The linearity of the coil edges almost meets the requirement of minus 0.05

deviation, as shown in Fig. 18, although it is sometimes out of range on the top of the coil.

Second, we developed a special jig that can butt the edge to decrease the gap. Fig. 19

shows the action of this jig in tree steps: 1st step - The ABC D E panels are placed on the jig;

2nd step - The jig is partly lifted, and seven cams push the E D panels to the location block.

After that, six pads cramp the E and D panels; and, 3rd step - Three pads push ABC to strait

and to butt strongly to the edges of E and D panels when the jig is completely lifted. Even after

all these efforts we could not achieve under 0.1 mm gap for a realisable mass production, as

shown in Fig. 20. Thus, we decided to increase the allowable maximum gap for butt welding.

Third, filler wire feeding welding (see Fig. 21) was evaluated in order to increase the

allowable gap of butt welding. The maximum gap allowance was substantially improved as

shown in Fig. 22. This filler wire feeding brings other advantages. If we feed low carbon content

filler wire, we can decrease the hardness of this welded bead. Fig. 23 shows the relationship

between the bead hardness and the filler wires.

Fourth, we developed a welded bead inspection system to maintain bead quality as shown

in Fig. 24. This system is installed on the welding torch. The slit laser beam is placed just

0.20 ~ I Deviation Average_l range

---0.15 - -.---.-.--~- Theoretical maximum gap

0.10 ~ __ ~ ____ .. _. ___ .. __ . _______ . ____ Maximum gap for reliable mass production

0.05

o 300 400 500 600

Welding length (mm)

Fig. 20. Gap of butt welding.

~

1" ./\

Laser Beam

-----' ----Shear cut shapes

~ .... Non-filler welded bead 'j

J ~

Filler welded bead

Fig. 21. Welding with filler wire.

Raw

u

1.2

lJl

o.s

04

229

I'LlIe thidcnoss : 1.0 mal -e__ FillerweJdjng ______ quantity; 0.4 mm'1mm

._\-.......... -......... .-

Filler weId.iI\g

Non-6Der weId.iI\g

Gap of butt welding (mm)

Fig. 22. Allowable gap of filler welding.

Bead Center

Non-filler material: SPCC

> :r: ~ c:

]

~ >

200

100

-+-+--t- -

Carbon contel of filler wire

Fig. 23. Vickers hardness by carbon content of 0 L 1'-.6--'--...l.O.s--'-_....L_-'---'-0.S:---'---1'-.6-­

filler wire (0.01 % and O. 005 %. respectively). Oistenee from center of welding (mm)

Different bead detected cases Portrate treatment

~~~~~~ Hight Width Undercut Unmatch Gap EiE~~

surface

Computer analysis

Iv CD U: Undercut (j) L : Unmatch Ul W: Width @ A : Angle G) C : Center G H: Hight

Fig. 24. Welded bead detection method.

230

o Piling station

Q( ;~d. inspection - ~ation

work piece --»/// input ~/

~Oading station

welding & inspection station

CO2 Laser oscilator (3.5 KW)

Fig. 25. Middle size laser welding line.

Fig. 26. Back-door inner panel made from laser welded blank.

231

opposite the welded bead, so the CCD camera will detect its bright line of light. After the

portrate treatment, the computer analyses those 6 inspection items and will detect 5 fault types in

real time.

By using this technique, we can successfully realise the production of the integrated side

member blank.

Fig. 25 shows the equipment of step 1. It uses 4 stations, and 1 laser oscillator and it can

produce different panels by automatic change of the jigs. It is applied to the middle panels, for

example the reinforcement of a sun-roof, the inner panels of a caul, the inner-panel of a

backdoor, and so on.

The actual panel of a back-door inner panel which was made from laser welded blank is

shown in Fig.26.

Fig. 27 shows the equipment of step 2. It has 5 stations, and 2 or 4 laser oscillators and it

can also produce these panels by changing the jigs. It is applied on the larger panels, for example

the side member. The actual panel of Toyota Lexus' side member made with an integrated one­

sheet blank is shown in Fig.28.

Work piece input

CD-Loading station

0---Welding & inspection station l

I 0'

Grindinb st<lLiOI1

guide

CBN grindstone

Fig. 27. Large size laser welding line.

/' Laser beam guide ~/)T':'--v~ /'

® Piling station

Laser oscilator (5 KW) x 2 sets

or (5 KW) x 4 sets

Workpiece output

232

Fig. 28. Toyota Lexus' side member made from an integrated one-sheet blank.

Mdtel;.,1 vield-up zone: ~

Side member paneis

integrated One-sheet

Fig. 29. Application to actual body and details of actual components.

233

Fig.29 shows the step in our development of this manufacturing process.

In 1983, we began the development of laser welded beanks for material yield-up and in

1986 we started with the production process that is step 1. After some experience, we found the

way to produce the integrated one-sheet side-member panels, in 1989, represented here as step 2.

At present, this method is used for 4 different passenger vehicles.

Fig 30 shows the growth of laser welded blanks. At this time, we have 4 middle size and 4

large size laser welding lines. Every year we produce I million middle size blanks and 0.6

million side-member blanks.

'86 '87 '88 '89 '90 'nl

X 1(J3

Middle size laser welded blanks

Laser welded blanks of side member

J o L-----"86r--'S..,..7--'S.:;:8----'-'"TS9--., .... clo--'9r;-

Fig. 30. Number of laser welding lines (right): and, annual production rate of laser welded blanks (left).

LASER APPLICATIONS FOR 3-D COMPONENTS: BEAM DELIVERY SYSTEMS AND ROBOTICS

L. PERA, P. PERLO, E. RABINO Fiat Research Center Strada Torino 50 10043 Orbassano (TO).Italy

G.MARINONI COMAU Robotics Division Strada Orbassano, 20122 10092 Beinasco (TO),1taly

ABSTRACT. The greatest opportunities for high power lasers in mechanical industry will lie in those areas where ease of optical manipulation goes hand in hand with the flexibility of robotization and CAD/CAM integration: for example, in precision 3D cutting and welding. In these areas, laser technology offers many advantages, ranging from high flexibility, reduction of time to market and product differentiation to CAD/CAM integration, automation and cost reduction. When laser beams were first coupled with multiaxis robotic systems, the number of optical elements drastically increased, and most industrial setups were used without a deep understanding of optical problems. It turned out that an excess of laser power was needed for the applications. As laser robotics moved to higher power, the hidden problems emerged. Proper solutions to maintain the beam quality had to be quickly found. More recently, production systems have been developed taking into account all the optical problems from the source to the workpiece. High quality beams associated with better accuracy of the delivery system, flexibility and software capability are making possible innovative cutting and welding applications. The automotive and the aerospace industries are taking advantage of this recent evolution. COMAU under the "Progetto Finalizzato Tecnologie Eletroottiche" of the Italian National Research Council (CNR) has specifically developed a robolaser to meet the most stringent specifications for both welding and cutting. This chapter examines the requirements of a robotized laser system, describes our anthropomorphic design approach, and shows some novel applications for the automotive industry.

1. Introduction

In the twenty years following the introduction of the first high power laser in industry, the

number of systems installed has been increasing continually. A detailed analysis [1] has shown

235

S. Martellucci et at. (eds.), Laser Applications for Mechanical Industry, 235-262. © 1993 Kluwer Academic Publishers.

236

that high power laser systems now number more than 2000. This has been made possible by

developments in optical beam delivery systems (BDS), as well as by a better knowledge of the

inherent characteristics of laser processes.

The 1970s were characterized by rigid systems with simple geometry, few optical

elements and medium power sources. During the 1980s, flexible systems were introduced with

Cartesian (gantry type), cylindrical, spherical polar and articulated (anthropomorphic) geometries

in many applications. Experience with such units has spurred COMAU to develop a new

anthropomorphic robot designed especially for integration with a C02 laser.

This chapter, after discussing some basic aspects of laser propagation, describes how to

compare the performance of different beams. It continues describing those effects that contribute

to beam quality degradation through the BDS. The usual requirements for a laser system and an

overview of technological development, with particular attention to the most notable robotized

systems, are given. Typical tolerances for cutting and welding are reported for new applications

made possible by the latest generation of articulated "robolasers".

2. Beam propagation and beam quality

The description used in this section is oriented to the generation and the propagation of

industrial C02 laser beams.

A laser beam has an intensity distribution that can be sufficiently well represented in a

general form as [2]:

1 ~ (r, z) 12 where ~ (r, z) = ~ mn cmn 'I'mn (r, z)

The set of orthonormalized functions {'I'mn} is defined in a region G with dimensions of

the order of magnitude of the laser beam diameter and orthogonal to the propagation axis z. 'I'mn

are the so called transverse mode configurations. The most convenient set of modes to describe

the laser beam intensity distribution depends on laser resonator geometry. Axial or transverse

flow stable resonators are most readily represented respectively by Laguerre and Hermite modes.

Commercial sources have usually beams with a prevailing mode, i.e. a single Cmn that tends

towards unity. Longitudinal flow lasers present in the market generally have cylindrical modes

TEMoo, TEMoI *, TEM02* . Likewise [3] for transverse flow we have rectangular modes

TEMoo, TEMI5, TEM16, TEMI (2)7 .

237

In this latter case the cylindrical symmetry of the resonator is only partly broken by the

transverse flow, and Hermite modes represent the real beams less accurately than Laguerre

gaussian modes describe axial flow lasers. The substitution of rectangular with elliptical

coordinates leades to elliptical modes which better represent the beams originated in the

transverse flow sources.

The theoretical considerations for an ideal gaussian beam are essentially: - the transverse

envelope preserves its shape along the optical axis; and, - the radius of any mode in the plane (i,

z), is related to the "parent" TEMOO radius according to the equation:

i=x,y

where: woi = radius of the TEMOO mode at the waist z = zO; zr = 1tW20i / A., Rayleigh range;

and, Mi is a multiplicative factor. The envelope of any coherent laser beam (no matter how its

diameter is defined or what irradiance distribution it has) can always be represented by the

envelope of a Gaussian beam. Accordingly, to describe the laser beam diameter in free

propagation it is sufficient to determine the three unknowns Mi' woi, ZO [4]. Their theoretical

estimation, at different power levels, requires the knowledge of all resonator parameters

including flow and excitation conditions, guiding effects in the excited medium and thermal

lensing in the optical components. Much simpler is the determination of the unknowns by a

polynomial fit of the free propagation experimental diameters taken at different distances from

the laser window. We can state that in the near field the experimental procedure gives a

coefficient Mi which expresses the radio W(z)/w(z) relating the real beam, no matter whether it

is originated in an axial or transverse flow resonator, to the "parent" TEMOO' For a given source,

the three parameters are functions of both power and lasing time. For a continous wave 5.0 kW

source, due to thermal lensing, there might be a considerable variation of the beam envelope if

the power increases from 2.0 kW to 5.0 kW, Fig. I.

For pulsed lasers there is one more problem to deal with; i.e., the transient time necessary

for the beam to reach its" quasi stationary regime". While "steady state" conditions of the optical

components are usually reached in a fraction of a second, the medium and the electrodes need

much longer time. During this transient the beam may change its diameter, waist position, and

pointing direction, and often the dominant mode (or, more accurately, the irradiance distribution)

changes as well. The coefficients Cmn vary rapidly in time at the beginning of the transient until

a "quasi stationary regime is reached", Fig.2.

The residual fluctuation of the dominant mode intensity in the "quasi stationary regime"

has still strong impact on high precision cutting and high speed welding. For a pulsed laser a

further distinction should be made according to the process.

238

60r---------------------------------------------~

E 50 §. II) 40

! 6 30 !!l ] 20

.~ III 10

Fig. 1. Thermal lensing effect on a C02 CW laser source at 1 and 5 kW.

0.9 RISE II)

TIME -a MODEl ~ 0.8

c: 0.7 Q) Q)

! 0.6

.2l ~0.5 'iii :ii 0.4

'E :i 0.3

~ 0.2

~ 0.1

2

0= 1 kW 6=5kW

o

3

5 10 15 20 25 30 Distance from output window (m)

MODE3 dominant Fig.2. The intensities of individual modes vary much at the beginning of a laser pulse. After a certain time one single mode dominates but residual fl uctuations of secondary modes remain.

4 5

Oscillation time (U.A.)

Fig.3. Thermal lensing of a 5 kW axial flow source in a pulse mode. "Cold" corresponds to the ftrst second of operation; "Warm" after 180 seconds of continuous operation.

50

10

PULSED FAST AXIAL FLOW SkW

WARM

2 3 4 5 6 7 8 9 10 11 12 13 14 15

Distance from output window (m)

239

In fact, if the process lasts few seconds and a limited number of pulses are employed, then

the process is carried out in a very unstable regime. There is mode competition filling within the

single pulse and the train of pulses combined with thermal lensing of both lasing medium and

optical components. If instead the process requires the beam to be on for more than 3 to 5s in an

almost continous wave operation, then during the first 3 to 5s we have a varying beam envelope

and mode competition-filling: the "cold regime". After these first seconds both beam envelope

and mode structure are rather well defined: the "warm regime", Fig.3. This overall thermal

transient is very much related to the resonator design and stronger effects are obviously present

at high power levels.

As a first approach, thermal lensing can be thought of as the phenomenon causing marked

defocus and lower spherical aberration (transverse and axial flows), astigmatism (transverse

flows) and tilt. A partial compensation of these aberrations in some CW lasers is obtained by

appropriate output coupler and window radii of curature (the elements are separated). For pulsed

lasers the compensation would not be so effective and is usually not applied. Although

considerations in the near field give the first useful indications, for better understanding a far

field analysis is necessary. Reconsidering the idealized Gaussian beams the half angle far field

divergence is:

A.M OFF=--

1t wO

The measurement of M2 can be accomplished with a focused spot diagnostic system. At the

focal point of a lens the far field divergence becomes:

OFF = Wj(z) f

where Wi (z) is the real beam radius at the focusing unit with focal lenghth f and located at a

distance Z from the resonator window. We have then a far field measurement of M2 given by:

M2 = [i artg(~)WO] = (4: j 2Wi . 2WOi) = (41tA. j D dO) .

The parameter M2 expresses the number of times the spot envelope of the real beam is larger

than the envelope of the TEMOO beam having the same l/e2 diameter at the focusing lens (the

parent TEMOO would give the ratio M). In the contest of the Gaussian Beam treatment M2 is the

so-called beam qUality. The beam quality is often expressed in normalized units as:

K=!:---1t OFF Wo

(K=1/M2)

240

In Fig.4 measured values of the propagation parameter M for transverse and axial flow C02

lasers are reported. The data refer to measurements we made at FlAT in 1986 when a large

number of new sources were acquired [5]. The situation is essentially unchanged. From the point

of view of both BDS designers and process engineers the two experimental approaches

mentioned above, for near and far field analysis, are complementary to describe the performance

of a laser beam in either continuous or pulsed operation. The approach of the Gaussian beams

can then be adopted in many phases of the overall optomechanical design of the system. For

instance, at a determined power, can we accurately calculate the variation of both spot size and

spot position in a variable path robot such as the gantry type, Figs. 5,6. Similarly, with different

M. wOo zOo we can compute spot size variations and shifts due to the change of the power or the

continuous spot shift and its envelope variation during the first seconds of operation. To

overcome these problems at least in part the known solutions are: - Start the processes when the

beam is sufficiently stationary; - Move the workpiece instead of the beam; - Avoid variable beam

paths; - Use a larger beam to reduce its divergence; and, - Keep the beam path as short as

possible.

4r----------------------------,

... J& If

3

r:: 2 o ~ OJ

~ !2 c..

uia1Flow

oL---------L-------~ ________ ~ o 2,000 4,000 6,000

Power (Watt)

Fig.4. Near field propagation coefficient M for C02 laser source.

E .§, C 0.5

.Q

."1: <J)

8. c '8. "3 .g .f (0.5)

!!: ..c; (J)

Focallenght: 127 mm Laser parameters:

M = 2.63 w. = 6.14mm Z. = 5200mm

7 e g 10

Distance laser to lens (m)

241

12

Fig.5. Variation of the axial position of the focal point vs. the distance between laser and focusing lens.

0.4 ,-------------------------

0.3e

E g 0.36

~ 'iii

8. 0.34

rn

0.32

Focallenght: 127 mm Laser parameters: M = 2.63 wa = 6.14mm Zo = S200mm

0.3 4'-~-'--~-'--~-.L7 -~---Le-~---1.g-~---1.10---'--..---J1-1 -~~12

Distance laser to lens (m)

Fig.6. Variation in focal spot dimension, related to the distance between lens and laser source.

3. Transfer of the beam through the BDS

The purpose of this section is to introduce a methodology to integrate the Gaussian beams

analysis for a better approach to the overall BDS design. With this in mind we will use one of

the first definitions of beam quality still in wide use:

242

B.Q. = [ P.I.B.th. )112 P.I. B.exp.

where P.I.B. = power in the bucket. This parameter compares powers within a prescribed area

rather than spot sizes. The ideal theoretical beam has unifonn phase at the focusing device

system. We will compare the measured power in the central peak at the best focus (p.I.B.exp) to

that of the diffraction limited beam (p.I.B.th). Assuming that for moderately aberrated beams

the ratio of the peak irradiances is nearly that of the P.I.B values, we have:

B.Q. {~y/2 = (i.)112

where 10 and I are the far field peak irradiances of the non aberrated and aberrated beams and S is

the Strehl ratio. At the entrance of the BDS the B.Q. is that of the laser, and the wavefront has

usually not a unifonn phase; however, we can suppose that the degradation of the B.Q. due to the

passage of the beam through the BDS is given by the multiplicative factor IIYI, where 10 and I

represent the far field peak irradiances respectively of a non-aberrated beam at the entrance and

of the aberrated beam at the output of the BDS. The task of the design is then to maintain the

ratio 1/10 as high as possible.

The approximate fonnulation we adopt is similar to the one proposed by Holmes and

Avizonis [6]:

-.!.= RN Td TM Ts F 10 1+ A

The meaning of the symbols are as follows:

(a) RN = product of mirror reflectances: values are 0.90 to 0.97.

To control the highly variable quality of the mirrors used in laser systems within Fiat Research

Center, we developed a reflectometer at 10.6 J.Illl able to measure the reflectance with various

angles of incidence and polarizations [7]. The accuracy has been estimated to be below 0.2%,

giving adeguate data for the BDS designer.

Td is the trasmittance factor to account for diffractive losses over mirrors, lenses and diaphrams.

Values are Td= 0.97 to 0.99. Refering to Belland and Creen [8] we have the following cases:

243

1) a/W> 2 the beam preserves its overall charateristics; 2) 1.8 < a/W < 210ss of power in the

range 0.15+0.03%, minimum effects on diffraction; 3) 1.6 < a/W < 1.8 loss of power in the range

0.6+0.15%, the wavefront is slightly disturbed with an increase in divergence around 5%; and,

4) 1.3 < a/W < 1.6 loss of power higher than 0.6%, appreciable diffraction effects.

where (l is the absorption coefficient, P is the scattering coefficient and the subscripts refer to

molecular and aerosol phenomena. Typical values are TM = 0.998 to 0.999. Over a 10 meter

long unconditioned optical path at 10.6 J.l1ll we have losses in the range 0.10-0.20%. Even when

losses are very low, the path must be conditioned to avoid the priming of nonlinear absorption

whIch can take place in stagnating zones. The air in industrial enviroments has particles such as

smoke, liquid mists, oil fumes, etc., ranging in size from below 0.1 to above 50 J.l1ll; in addition

there are many molecular agents which show some absorption at 10.6 J.l1ll, hydrocarbons and

polyatomic molecules. Even in small concentration these cause scattering and absorption or

blooming. Very high purity nitrogen is the ideal propagation media. In some applications where

the BDS has an open reflective focusing unit the consumption of N2 is increased and considered

too costly by the users. Temperature controlled filtered air is then used.

(d) T,= [1-( 4.<\ co,otT Ts is the trasmittance factor to account for scattering losses. In it the RMS surface roughness is

in Angstroms. Typical values are 0.990 to 0.995.

(e) A=~(~)2 A is a factor to account for the broadening of the focal spot due to vibration in the BDS, jitter

and turbulence. The calculation of these effects is rather complex as they contribute on different

time scales. If the path is conditioned with a laminar flow then the residual turbulence on an

unexpanded C02 laser beam does not introduce noticeable variations. The motors, mounted on

the robot to move the articulated arms, introduce vibrations in the range of 100 Hz. Coupled with

a low but appreciable drift as a consequence of heat absorption, there is a faster beam jitter (some

Hz) originated in the resonator and in the cooling channels of the optical components. The

hypothesis is that the peak irradiance move around the centroid with zero average in less than

l/100th of a second. Values of B. Q. at the exit of an articulated robot are A=O.1 to 0.2.

244 2.2

2

• 1.8

= ~ 0

1.6 CU

0 ::J r::J' • 1.4 0 E !U 0 Q)

CD 1.2 * <) 0 •

0 0

* 0

0.8 0 0.1 0.2 0.3 0.4 0.5 0.6

Waves (0)

Fig.7 Beam quality as a function of different aberrations. *=Defocus, 0 = Astigmatism, 0 = Coma,. = Spherical aberration. B.Q. =(86%/energy in the first ring) 1/2

~

~ I/O Z ... .. z

o ... ... ... <C 2 I(

o z

1. 0 ----,-"-T"--r-~-r---,

2 " NOA ..... lIZEO 01$ rANeE

Fig.8. Relation between the normalized intensity 1/10 for three different wavefront aberration values.

245

F describes near field phase aberrations in the wavefront at the focusing unit exit to account for

thennal defonnations and overall figure errors. The control of the wavefront is the dominant task

in multimirror articulated BDS; the weight of this factor alone in the degradation of B.Q. is

expressed in Figs.7,8.

Our analysis follows Bennet, Klein and Palmer [9,10,11]. The basic ideas are: 1) in a plane

located at the entrance of the BDS and perpendicular to the beam axis, there are no phase

aberrations; 2) there is no correlation among different components and their figure errors; 3)

wavefront errors due to thennal distortions are correlated with each component and are

coherently additive; 4) all other non thennal distortions are uncorrelated with each other, or with

thennal distortions; 5) uncorrelated path fluctuations are Gaussian and random in amplitude with

zero mean value. The factor F becomes:

where:

K = 21t A

(52 = .121, RMS + V AR (AT)

N (52 = N 't2F,M 2

2 (n-l)2 2 (5 FM= 'tFM , 4 '

N NI -2 N Power fAT,M = - (FOM)M = (FOM)M' Area of the beam

It 2 N Power AT W - = --------

, - (FOM)W (FOM)W . Area of beam

N LAs = 2NOT = function of power 1

N Al +.12 +.12 . rAR =N = functlOnofpower 1 2 (2F 2

, and

246

A B c Pig.9. Real time deformation of: A) an uncooled copper mirror, B) silver coated copper mirror, C) high reflectance coated silicon mirror loaded with a 22 mm diameter 5.5 kW, TEMOI beam at 450 and S polarization.

The symbols are defined as: V AR (A) = < A2 > - < A >2 =(RMS)2 = 0 2 : the variance of the

wavefront error A; 0P, M(W) : RMS wavefront distortion caused by the errors in the figure of

the single mirror (lens/window); 'tp,M(W): errors in the figure; AT,M : expresses the thermal

expansion (perpendicular to the surface), the second one due to the radial profile of the

temperature; As: expresses the bowing caused by the axial gradient of the temperature (its

expression can be given in two different forms for clamped mirrors or for three point support

(11)); AR: accounts for the ripples caused by the rectangular COOling channels close to the

reflecting surface (it is the sum of three terms: deflection of the "plate" above the web A I, the

expansion of the web A2, and the pressure bending of the "plate" A3 [11]); (FOM)M: figure of

merit for the mirror, in units of Watt per square centimeter per centimeter; I: Intensity of the

beam; and, (FOM)W: figure of merit for the lens or window, in units of Joules per square

centimeter per centimeter.

The main conclusion is that if we neglect the thermal distorsions in a ten mirror system the

Marechal criterium (0=N'14) is reached for 'tp,M = A. / (7t 10). Then the error in the figure, once

the components are mounted, should be better then that value. Thermally induced wavefront

errors are power dependent and increase linearly with power; in terms of the far-field

performance, this implies that, inevitably, there will be a point of diminishing returns in pushing

the laser output beyond some sensible limit. Even for industrial power levels that limit is easily

reached if the components are unproperly cooled. Turbulence, thermal blooming, beam jitter,

vibration in the system and factor P itself are time dependent, so the steady state is never really

reached.

247

Instruments that allow field measurements of the effective thermomechanical distortion of

a component are powerful tools to verify the soundness of the design, the finite element

predictions and coating behaviour [12]. Figs. 9a,b,c show some measurements done at Fiat

Research Centre on commercially available uncoated cooled copper mirrors, silver coated cooled

copper mirrors and high reflectance coated silicon mirrors loaded with a 22 mm diameter, 5.5

kW power, lEMOI beam at 450 (real operation). In the case of large thermally induced

wavefront errors driven by beams with an almost Gaussian distribution, the far field peak

intensity should be described on the basis of the slope distortion rather than through the RMS

wavefront distortion. Generally speaking, power losses at the edges of the components are

associated with heavier diffraction effects but for the choice of the correct aperture one should

recognize that wavefront distortions will dominate the far field behaviour; i.e, the components

must be reasonably small. At 10.6 j.Ull and with usual industrial unexpanded beams with powers

in the range of 1 to 2 kW the values of the factor F can be better than 0.7. For powers in the

range 3 to 5 kW values above 0.5 require high reflectance mirrors or complex design of the

mirrors. For powers above 2.5 kW the tendency has been to increase the beam quality by

reducing the resonator Fresnel number (or almost equivalently the beam size). Therefore the

BDS designer has to assure, by appropriate components and mountings, that the beam quality is

maintained at higher power densities. He also has to deal with overall beam characteristics such

as jitter, vergence, thermal lensing, long and short time power instabilities, ageing of the output

coupler, induced birefringence, control of the polarization, etc. As a general rule, laser and BDS

designers have to find the compromise for the right beam size to guarantee long term

maintenance of the original performance.

4. Presentation of beam performance data

A comparison among different laser beams can be done according to several points of

view. The following can be considered an adequate approach in the industrial environment.

4.1. NEAR FIELD ANALYSIS

Free propagation curves: radius as a function of Z, power, time of continuous operation.

The data obtained by fitting are Mj Woj, Zoj . Different powers for CW operation, warm and cold

conditions for pulsed operation should be considered;

248

Mode patterns: to be studied as a function of power and time of continuous operation.

CW: maximum power and half power. Pulsed: steady state and below I sec;

Polarization: as a function of power and time: short and long time fluctuations at

maximum power;

Beam pointing stability: as a function of power and time (Measurable by several

equivalent methods). CW: short and long time at maximum power;

Power: as a function of time. CW: long time at maximum power. Pulsed: transient and

long time at maximum power.

4.2. FAR FIELD ANALYSIS

Spot size and far field divergence (or power within a prescribed bucket): function of

power, time, position of both focusing unit and measuring systems; need to specify: time of

continous operation, hypothesis of an ideal beam at the focusing unit, position of the focusing

unit, aberrations due to the focusing unit, position of the minimum measured spot, integration

time, spot/power analyzer used.

4.3. SPEEED AND QUALITY OF PROCESS ANALYSIS

Another point of view concerning beam quality comes from process engineers. Since they

are faced with factors such as speed and quality of processes, it is understandable that their

favourite method to compare laser beam performance is based on measurements such as speed of

welding, depth of penetration, keyhole shape, etc. TIlis approach has the advantage of

discriminating polarization effects shown by many treated materials and not considered in

previous beam quality definitions. The comparison can be done on the application under study

but may lead to the most controversial comparisons. As is known, speed and quality of processes

depend as well on factors such as: - direction of process; - parameters and composition of the

assistant gas; - orientation and shape of the nozzle; and, - position of the minimum spot size with

respect to the workpiece.

The data on welding speed (or cut speed), depth of penetration, quality of keyhole, have to

be presented together with the above factors and the further specification of the position of the

focusing unit. The final comparisons and interpretations should be done together with the near

field analysis. We can state that the process engineers approach is "equivalent to a far field

analysis".

249

5. Types of "Beam Delivery Systems" (BDS)

It is possible to distinguish between four laser system categories for which robotization is

involved at different levels, viz. [12]: 1) Stationary BDS and moving workpiece; 2) Coupled

BDS and workpiece movements; 3) Robot which moves the laser; 4) Moving BDS and rigidly

retained workpiece.

5.1. STATIONARY BDS AND MOVlNG WORKPIECE

A number of solutions are currently used:

(a) The beam is delivered simultaneously to different workstations through semi reflecting

mirrors (beam splitters). Representative applications include cutting operations for fabric, paper

and laminated plastics. Power fraction on each lens is PIN, where N is the number of work

stations. Workpiece is generally moved via rollers and process tolerances are not very stringent.

(b) In the "time - sharing" solution (Fig.10), one or more mirrors are used to direct the beam to

the desired work station. Process characteristics will dictate the choice between layouts. Delivery

from one station to another usually requires only a few tenths of a second. The focussing device

carries a lens or a pair of mirrors, depending on process type and power. For cutting, the

focussing device is almost always selfadapting and is generally provided with manual or

automatic raising and lowering.

(c) The workpiece may be moved by a robot, or by indexing tables and/or programmable X, Y

tables.

5.2. COUPLED BDS WORKPIECE MOVEMENTS

This solution is mainly used with gantry robots or in general with those featuring an

extendible axis. As discussed before, for these robots the problems are the beam directional

instability and divergence, together with the poor rigidity typical of telescoping extensions, limit

process accuracy and homogeneity in the work envelope. For certain processes, the combined

motion solution is necessary for robots with limited articulation and for those covering narrow

work envelopes. In both cases, workpiece movements compensate for the inherent shortcoming

of the laser robot. There are many possible combinations between BDS and workpiece

movements. Of the 5 axes which are nonnally available, some implementations use 3 (BDS) + 2

(workpieces) while others employ 4 + 1.

250

--1------1 Fig. 10. Simple beam delivering system used on most of current industrial units. The concept of time sharing is shown for the example of a three station system.

--Fig.1I. This configuration shows a laser integrated in the arm of a robot. The potentialities of this solution are high but up to now no existing industrial laser is able to withstand the accelerations and vibrations in such an environment.

251

5.3. ROBOT MOUNlED LASER

The first attempts to install a high-power C02 laser on a robot date back to the late 1970s,

while the 1980s saw the introduction of a variety of layouts with questionable success, Fig. 11.

Basically, this lack of success is to be attributed to the weight of sources and to the structure of

the resonators, which are always very sensitive to vibration and hence to acceleration. An

example of good coupling is the solution with translation and rotation relative to a vertical axis

of the laser and integrated with a motorized conduit providing an additional 3 degrees of

freedom. The solution, using lightweight sources is of considerable interest for processing small

parts. The layout is not suitable for parts of large size or complex geometry on account of the

difficulties in combining the laser's rototranslational movements with those of the conduit. The

development of new C02 resonators has led to compact and lightweight sources of up to 800 W

with good insensitivity to acceleration. These units include the so-called "sealed off' type, others

where gas is replaced at frequent intervals, though not continuosly (Le. without expensive and

bulky pumps), which make use of gas reconversion. Still other types feature new methods for gas

cooling. Nd-YAG lasers have been used in a number of robot-mounted applications where a

mean power of 300 to 400 W is sufficient. For ND-Y AG sources with higher mean powers the

problems are essentially those of the C02 laser. Systems using optical fibers (Fig. 12) remain a

valid alternative solution, at least so long as the beam quality of solid state sources remains far

from the diffraction limit. Here the articulated robot offers interesting prospects for various

coupling techniques. In terms of lasers mounted on robots the expectation for the 1990s is a

single focused spot of incoherently added beams of laser diode arrays.

Fig. 12. Coupling of Nd-Y AG with standard robots is becoming of industrial interest specially for cutting applications. Safety problems must be seriously considered if the system is located in an open industrial ground.

252

5.4. MOVING BDS AND RIGIDLY RETAINED WORKPIECE

When rating a BDS it is necessary that the type of application be taken into account. In

general tenns, it is assumed that the mobile BDS (robot) is a part of a flexible manufacturing cell

used for components which may vary widely in configuration and volume. The workpiece is

rigidly mounted on a support and it is assumed that access is possible on all sides through BDS

movements only. Where this is not possible, it is assumed that a second BDS is used. Robotized

BDS types capable of more or less complex 3 D processing are usually classified as follows: -

cartesian/gantry (variable optical path - V.O.P.); -cylindrical (V.O.P.); -spherical (V.O.P.); -polar

or "oblique" coordinate transfonnation (V.O.P.); -self-supporting articulated (pre-detennined

optical path); and, -integrated articulated (pre-determined optical path).

5.4.1 Cartesian/gantry robot

These are the best-known and most widley used robot types. The beam moves along three

mutually perpendicular axes, X, Y and Z. Two rotational degrees of freedom are usually included

in the focusing head (Fig. 13). Some setups use a floor-mounted laser, while others support the

laser overead on side pillars. Lasers moving in the X, Y plane are sometimes used as an extreme

measure in order to restrict optical path variations to the Z axis; however, the complications of

providing movements for a gas laser resonator offers advantages which are purely hypothetical.

The required number of optical elements depends both on the layout used and on beam

manipulation such as depolarization and/or expansion. In the current solutions, the number of

optical elements ranges from a minimum of 4 to a maximum of 9.

5.4.2. Cylindrical coordinate robot

From the conceptual standpoint, these robots are highly efficient, as they provide 5

degrees of freedom with a minimum of 4 mirrors for a relatively short optical path. The number

of optical elements becomes 5 if a lens is used to focus the beam, or 7 if the beam linear

polarization is in the vertical or horizontal plane. The beam is usuallv inserted from below along

the vertical rotational axis, but the solution with entry from above is equivalent. With this

structure, the major limitation is poor end articulation. The first transfer axis is accomplished by

moving the entire robot. This significantly reduces maximum possible speed through relatively

long trajectories, while coupling the movements is also difficult. The variable optical path may

require the use of a beam expander with at least two more optical elements.

253

- --- --

Fig.13. Configuration of a gantry robolaser. The optical path length continuously varies from the source to the working point during operation.

A/../"A STA'ZIONE. /t>l J..AVoR.0 N°2

Fig.14. Configuration of an articulated arm joined with a standard robot. Some similar systems were proposed in the past few years without real industrial success.

254

Fig. 15. View of the "robolaser" beam delivery system. Two reflecting mirrors are located in each joint.

I I

-~- . // .. , +- ..... , ,~" ~:2:: I '\ \ !.! ___ JlJRSO

lJ.-=-~;;,t:.~3 ,I \ ' . .' I 3450

Fig.16. Typical configuration of a working cell that includes a traditional spot welding system for tags and a "robolaser"

255

5.4.3. Spherical coordinate robots

This type has good coupling efficiency with the laser beam, as it is possible to use a

minimum of 5 optical elements for a total of 5 degrees of freedom. If the beam is focused by a

lens after the laser window, there will be 6 elements, which rise to 8 if the polarization plane is

vertical or horizontal. In this robot, the third or transfer axis does not pass through the primary

rotational axis. Aside from this, movements are very similar to those of robots with "oblique"

coordinate transformation. The major limitation on the spherical robot is its poor end

articulation. As a result, and because of the relatively small work envelope, it can only be applied

to simple geometries. On the credit side, optical path variation is low and the structure is

relatively simple.

5.4.4. Robots with "oblique" coordinate trans/ormation

Here again, coupling efficiency with the laser is high. In fact, a minimum of 4 optical

elements can be used for a total of S + 1 degrees of freedom. In existing versions, the end is

fairly bulky, but with relatively simple modifications this can become highly competitive. Given

its similarity to the spherical robot, the same considerations apply regarding the number of

optical elements and the range of applications.

5.4.5. Self-supporting articulated BDS

These types represent the first attempts to deliver the laser beam along several degrees of

freedom. They consist of an articulated conduit resembling a dentist's drill in shape (see Fig.14).

The end of the conduit is moved in space by a conventional robot for which no substantial

modifications are necessary. Because of mechanical interference between conduit and robot,

these units can be applied only to simple geometries. As the multijointed conduit is susceptible

to vibration and undesired movement in general, this solution is very sensitive to misalignment

and thus not suitable for precision work. Current layouts include some 5 mirrors, others 8 and, in

extreme cases, as many as 14 optical elements.

5.4.6. The COMAU new Smart-25 L integrated articulated robot

This COMAU proposed solution is highly efficient because the robot joints are used as

optical joints while its structure is used to conduct the beam (Fig. IS).

256

Developed jointly by COMAU and the Fiat Research Center using the Smart-50 robot as a

starting point, this type of BDS is now offered in a new form designed expecially for coupling

with a laser. Beam transfer from the resonator window to the workpoint is performed with a

minimum of seven components for a total of five degrees of freedom. Nine components are used

if the laser beam linear polarization is in the horizontal or in the vertical plane. Advantages of the

integrated articulated solution include: -no variation of the optical path; -short optical path; -laser

source is not moved; and, -high articulation at the end. The last two points are the intrinsic

advantages of the integrated articulated robots and make this the most suitable solution for

flexible manufacturing cells (Fig. 16) or for parts with complex geometry in general.

From the standpoint of laser coupling, the characteristics of the robot are as follows: -

conduit minimum inside diameter is 45 mm; -structural vibration and thermal distortion must be

extremely limited. These are the aspects which received most of the attention in the new

optomechanical design; -optical components can be replaced without repeating the alignment.

This is made possible by the use of patented supports (Fig. 17).

Fig. 17. Innovative beam deflecting mirror system with precise regulation and water cooling. Rotation of the mirror is allowed around two axes without introducing any translation of the beam.

257

i <.;) 6.Axls

I , I' Iii 1

Fig.18. Two focussing exchangeable heads: A) is for cutting processes, allows 3.75 and 5 inches focal length with autofocus and protection against collision; B) is for welding processes, allows 150 and 300 mm focal length and path-oriented wire feed.

The main characteristics are: -the optical path is sealed off. Its conditioning system

prevents damage to beam quality and guarantees excellent protection for optics; - the mirrors

have a reflectance of over 99.6% and they maintain polarization status over the entire optical

path (retarders, when used, have a typical reflectance of 98.5 %); and, - beam focusing can be

accomplished by a lens or by a parabolic mirror in relation to the application and the beam power

(Fig. 18).

Fig.l9. The Robolaser is a very flexible production tool. If needed, more than one system can be fed on a time sharing basis from a single laser source as shown in the above picture.

258

In terms of adaptability to different applications; the robolaser features: - a high level

software language to program the controller; - power and laser cycle data adaptability of various

sources, which can be stored in the robot control memory for varying process strategies;

- an optional 6th axis to orient a welding wire in a controlled path movement; - autofocus; -

predisposition to dialogue with process control sensors; and, - time sharing (Fig. 19).

The information flow between laser and robot control handled by serial bus allows:

- reduced wiring; - increased reliability; and, - reduced installation expediture.

The peripherals used by the laser and robot are: - gas distribution unit using high pressure

gas bottle storage and low pressure distribution valves and pipes; - laser and mirror cooling

system; and, - air conditioning system for the beam path having high filter capacity flow and

temperature control.

The design criteria to deliver the beam through the Smart-25 L are summarized in the

following . Beam quality is well known to be a key point in processing. Its degradation due to

poor design, of either the resonator or BDS, implies both poor quality of the process and a

diminishing return in increasing the laser output beyond some reasonable limit.

1900

Fig.20. Working volume of COMAU 5.25L robolaser. The robot can also be located on a side wall or hung from the ceiling.

Fig.21. Two structural automotive nodes designed for laser manufacturing. A 300% increase of stiffness was reached without increasing the structure weight and volume.

NODE "0"

NODE "0"

---- continuous laser welding

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Fig.22. A view of node D of Fig.21; the structure is welded both with laser and resistance welding according to needs.

260

In the Smart-25 L development special care has been taken to maximize beam transfer

efficiency. These efforts have lead to a system that allows both high quality cuts and welding at

power levels never achieved before using multiaxis BDS. The design is a substantial

improvement over its predecessor and thus should yield even greater advantages on highly

complex shapes. The good end articulation coupled with wire feed welding opens the way to a

wide range of new applications of lasers in the development of advanced aerospace and

automotive structures (Fig. 20).

6. Applications of the" Articulated Robolaser" for the automotive industry

In the automotive industry, the availability of the robolaser gives opportunities for

developing new structural designs. The capability to perform welding on 3D paths allows

subdivisions and connections impossible by conventional welding technologies. On the rear of

the car, the so called nodes "G" and "D" shown in Figs. 21. 22 are significant examples

Fig.23. Complex 3D automotive structure 90% joined with continuous laser welding.

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Fig.24. Typical profile contours of an automotive structure for butt welding. The new COMAU Robolaser has the needed precision to follow the entire path for welding both with and without wire feed.

examined to increase the stiffness of the body. A considerable increase in stiffness has been

obtained with the integration of resistance spot and continuous laser welds as compared to simple

resistance spot welds. The high flexibility and accessibility of the Articulated Robolaser is the

key factor allowing the continuous welding paths of Figs. 23, 24 inside such complex structures.

References

1. Roessler, D.M. (1989): "Laser processing: Global overview and future trends". ISATA Congress, Wiesbaden, Germany.

2. Oughstun, K. E. (1982): "On the completeness of the stationary transverse modes in an optical cavity", Opt. Comm. 42, 72.

3. Pera, L. and Perlo, P.(1989): "Industrial C02 lasers, BDS and applications update". 21st international symposium on automotive technology and automation, ISAT A Congress, Wiesbaden, Germany.

4. Perlo, P. (1986): "Propagation of a multikilowatt laser beam: experimental characterization", Spie Vol. 650.

5. Alessandretti, G. and Perlo, P. (1988): "Optical design considerations for C02 laser industrial systems", in "Laser Science and Technology", A.N. Chester, V.S. Letokhov and S. Martellucci (eds.), Plenum Press, E.M. int'! Series, vo1.35, p.l77

262

6. Holmes, A. and Avizonis, P. V. (1976): "Approximate optical system model"; Appl.Opt. 15(4), 1075.

7. Perlo, P. (1990): "Laser beam qualification and high power optical component design and testing". ATA Conference, Turin, Italy.

8. Belland, P. and Creen, J. P. (1982): "Changes in the characteristics of a gaussian beam weakly diffracted by a circular aperture", Appl. Opt. 21(2), 522.

9. Bennett, H. E. (1976): "Thermal distortion threshold for optical trains handling high pulse power", NBS Spec. Pub!. 642.

10. Klein, C. (1981): "Mirror figure of merit and material index of goodness for high power laser beam reflectors", Spie Vo1.288.

11. Palmer, J. R. (1987): "Optical distortion of multilayer coated optical components used in high power laser systems", Spie Vo!.805.

12. Bognier, A., Marinoni, G. and Perlo, P. (1990): "Robolaser", FISITA Congress, Turin, Italy.

LASER BLANK WELDING AND STAMPING OF SHEET METAL PARTS

F. A. DI PIETRO Director, Manufacturing Systems (Retired) General Motors Corporation, U.SA. 9650 South Ocean Blvd. #1510 Jensen Beach, Florida 34957 (U.SA.)

ABSTRACT. This lecture is focused upon the use of Laser Welded Steel Blanks as applied to the body components and assemblies of automotive vehicle bodies; however, equivalent applications can be applied to the appliance industries, and those industries using sheet metal components, i.e., electrical, aerospace, computer industries, etc. The integrated subjects covered include the following: 1. Description and benefits of Laser Welded{I'ailored Blanks; 2. Structural analysis of Laser Welded Blanks; 3. Benefits and impact on product design utilizing Laser Welded Blanks; 4. Facilities for Laser Welding Systems; 5. Metallurgical characteristics, strain analysis and formability of Laser Welded Blanks; and, 6. Actual applications of in-production Laser Welded Blanks in the Automotive Industry.

1. Introduction

1. 1. DESCRIPTION AND BENEFITS OF LASER WELDEDrr AILORED BLANKS [1,2]

A brief description of Laser Welded!l'ailored Blanks is given analyzing three major

categories, i.e., material savings, parts consolidation and tailored blanks for specific applications.

The advantages of tailored blanks are reviewed, including:

Weight reduction;

Minimized number of parts;

Lower tooling cost of parts;

Lower press shop costs;

Lower assembly costs;

Improved fatigue behavior;

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264

Improved crash resistance;

Reduced number of overlap joints;

Improved corrosion behavior; and,

Less final sealing operations.

2. Discussion

2.1. STRUCTURAL ANALYSIS OF LASER WELDED BLANKS

The key hole welding process [3) is described together with a cross-section of typical

carbon steel weld bead profiles on Laser Welded Blanks. In addition, (1000 x) magnification

slides depict the microstructure of the base metal, the heat affected zone (HAZ) and the fusion

zone. These (1000 x) magnification slides clearly illustrate the changes of the steel

microstructure in these three zones, they clearly depict the fine microstructure of the fusion zone

which indicates the formation of martensite; which is one reason why the weld is stronger than

the base metal.

Structural and fatigue tests are described [1) comparing both spotwelded and laser welded

steel coupons. The results of exhaustive testing clearing indicate that the tensile strength of the

laser weld is equal to or greater than the parent metal itself. In addition, comparative testing of

both equivalent spotwelded and laser welded test coupons, indicate the number of cycles to

failure is significantly greater for laser welded coupons. These tests indicated approximately

18 mm of linear laser weld is equivalent to one typical spotweld.

3. Benefits and Impact of Product Design Utilizing Laser Welded Blanks [2)

Three key reasons to use Laser Welded Blanks include:

1. Material savings;

2. Parts consolidation; and,

3. Tailored blanks.

265

3.1. MATERIAL SAVINGS

Material savings are achieved through a reduction of engineered scrap and the utilization

of lower cost materials, i.e. bare steel versus galvanized steel, where corrosion protection is not a

requirement. Engineered scrap is waste steel resulting from formability requirements, product

design, part processing, blank nesting and press and die requirements. Engineered scrap can be

reduced by the optimization of the above parameters, the use of new technologies, i.e., Laser

Blank Welding and implementation of simultaneous engineering, between Product Design and

Manufacturing Engineering; engineered scrap for a complete body-inwhite can range from 30%

to 50% of total steel requirements and therefore, represents a significant cost improvement

opportunity.

3.2. PARTS CONSOUDATION

Parts consolidation represents another tremendous opportunity to improve the product

design by varying the specified material gauges and through use of overside blanks.

Many sheet metal components in an automobile body are fabricated from .8 mm thick

steel. This relative thinness requires that the part, i.e., a door inner panel, be reinforced with a

separate reinforcement. It is welded to the door inner panel, so that sufficient structure is

achieved so that the weight of the door hinges and completed door can function properly and

meet product design specifications. The use of multi-gauge laser welded door inner panel blanks

eliminate the need for these reinforcements, and provide the required structural integrity of the

door assembly. Intensive product testing has validated this conclusion.

Laser welded over-size blanks are used when steel coil sizes are not available in the gauge

and/or coating required. A classic case is the Audi-hot galvanized floor pan laser welded oversize

blanks.

3.3. TAILORED BLANKS

Tailored blanks are laser welded blanks which are tailored for a specific requirement by

welding different steel gauges, and/or different material coatings. Typical tailored blanks

currently in production are described which include door inner panels, shock tower assemblies,

body side door rings, wheel house, motor compartment rails, etc.

In each application, the product designer has the flexibility to specify the material gauge

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thickness to meet structural requirements while optimizing weight. In addition, only the material

coating required for a specific area is specified, thereby reducing cost, while meeting corrosion

requirements.

4. Facilities for Laser Welding Systems [4]

A typical commercial laser welding facility is described including blank guillotine

shearing, blank welding, weld inspection and cleaning and the complete material handling

system requirements. The approximate cost for a production facility of this type ranges from 15

to 20 million U.S. dollars.

A key requirement is the precision of the sheared edge (cutting tolerance of less than

0.04 mm in straightness) in order to achieve required weld integrity and welding speeds of 5-10

m / min. Types of edges, admissable welding gap widths and welding speeds for different laser

powers (kw) are reviewed on individual data charts.

5. Metallurgical Charactistics, Strain Analysis and Formability of Laser Welded

Blanks [5]

Extensive testing including limiting dome height test, circle grid analysis testing, and

formability comparisons, indicate that laser welded blank formability is almost equivalent to

unwelded blanks. This conclusion applies to laser blanks of similar thickness, multi-gauge

thickness and laser blanks of different types of steel. There are special factors which must be

considered such as the location of the laser weld in relation to the areas of deep draw.

The drawing quality of the laser welded blank is excellent with minimal degradation in the

draw depth as a result of the welds presence. However, there can be a reduction of stretch before

failure. If possible, parts should not be designed with welds located in zones where extreme

stretch-type strains occur in the weld direction. Fatigue strength is also excellent with values

equal to the base metal, if proper weld parameters are used. Also, test matrix have been

developed through testing to validate and correlate necessary parameters.

6. Actual Applications of In-Production Laser Welded Blanks in the Automotive

Industry [3,4]

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There are a significant number of actual applications of in-production laser welded blanks

in use today; however due to a variety of reasons, these production applications receive very

little publication.

6.1. THYSSEN-STAHL [2,7]

Thyssen-Stahl A.G. (Duisburg, Germany) since 1985 has supplied 1.7 million oversized

laser beam welded blanks for the Audi-Series 4000 and 5000 floor pans. These blanks are 3200

mm x 1950 mm x 0.75 mm made from hot dip galvanized steel.

6.2. TOYOTA BODY [6]

Toyota body side ring has been in production in Japan since the production of the Lexus

Vehic1e.-The blank is comprised of five individual pieces with metal thickness of .8 mm and

1.0 mm as well as three different material coatings including electro galvanize and "Excelite" an­

iron electro-galvanized steel. This type of door ring construction is also being used by Toyota on

its Corolla,Tercel and several domestic models.

6.3. PRECISION BLANKING LTD [5,6]

Precision Blanking Ltd. (Armco Steel Co. Joint Venture) is laser welding scrap steel

panels to make a blank used for a U.S. truck door side beam reinforcement.

6.4. ARMCO STEEL [5]

Armco Steel fabricates a multi-gauge blank for a center pillar inner panel for a U.S. auto

manufacturer. Originally the blank was made of 1.0 mm blank, but required a heavy

reinforcement to strengthen and stiffen the upper pillar where the safety belt anchorage is

attached. In addition to the extra operations required to produce the center pillar inner this way,

more fit-up and spot welding problems were created when the triple-steel thickness of the inner,

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the reinforcement and the outer were spotwelded together.

The new laser welded blank of differential thicknesses 1.8 mm upper and 1.0 mm lower

eliminated the need for the heavy reinforcement. The benefits included a weight savings of

0.7 kgm per vehicle, cost savings from elimination of dies for reinforcement plus elimination of

assembly tooling and welding operations. In addition to the foregoing in-production applications,

there are many laser blank weld applications currently being planned or in use by U.S. and

foreign automotive manufacturers; these include the following:

Multi-gauge door inner blanks;

Multi-gauge shock tower blanks;

Multi-gauge body door side ring blanks;

Multi-gauge floor pan blanks;

Multi-gauge upper motor compartment rail blanks;

Multi-gauge frame rail blanks;

Multi-gauge wheel house with integrated reinforcement; and,

Multi-gauge wheel house with integrated shock tower.

7. Conclusion

The foregoing represents only a preliminary list of known forthcoming applications. As

this new technology becomes more accepted with the Product Design community, a wide range

of new applications can be anticipated. The integration of process capability with the new

flexibility of design alternatives will result in significant cost savings, weight savings and total

product optimization.

The laser application challenge involves four steps to be implemented:

1. Information;

2. Understanding;

3. Commitment; and,

4. Action.

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References

1. Uddin, N. M. (1986): "Proceedings of S.P.I.E." Vol. 668, 260. 2. Prange,W. and Lobring, V. (1988): "Laser to the Work Place", IMTS, Chicago,Ill. U. S.A. 3. Di Pietro, F.A.(1990): "Proceedings of the ATA", ATA, Torino, Italy. 4. SeIige, A. and Prange, W. (1987): "SAE Congress" Detroit,Mich. 5. Neiheisel, G. and Cary, R. (1991): "Industrial Laser Review" p. 6. 6. Irving, B. (1991): "Welding Journal", p.42. 7. Uddin, M. M. (1991): "Industrial Laser Review", p. 11.

ROBOTIC LASER WELDING SYSTEMS IN AUTOMOTIVE OPERATIONS

F. A. DI PIETRO Director, Manufacturing Systems (Retired) General Motors Corporation, V.SA. 9650 South Ocean Blvd. #1510 Jensen Beach, Florida 34957 V.SA.

ABSTRACT. The primary focus of this paper will be upon the application of Robotic Laser Welding on automotive sheet metal components, sub-assemblies and the complete body itself. Also reviewed is the world's first production on-line application of Robotic Laser Welding of the roof panel to quarter panel joint on a complete body in the U.S.A. The major subjects covered include the following: Overview of laser welding vs. resistance spotwelding; Advantages of laser welding vs. resistance spotwelding; Design flexibility of laser welded joints; Requirements for successful laser welding systems; Advancements in robotic laser systems both C02 and Nd:YAG; Applications of in-production laser welded assemblies; and, Laser welding challenges of the future.

1. Introduction

1.1. OVERVIEW OF LASER WELDING, INCLUDING KEY HOLE WELDING [1,4]

In order to establish a basic understanding of laser welding fundamentals, a brief summary

of the laser spectrum, types of lasers and laser material processing systems are reviewed. Laser

fundamentals and a conceptual view of the laser process is outlined, including the several

requirements for laser action, i.e., the lasing medium, an optical resonator and a means of

excitation to create photons resulting in a beam of parallel, monochromatic and coherent light of

very great intensity, i.e. the laser beam.

Laser welding process parameters may be sub-divided into two groups, i.e., "Independent"

process parameters and "Dependent" process parameters.

Independent process parameters include: laser power, beam diameter, depth of focus, beam

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mode, beam divergence, and spot size. Whereas, dependent process parameters include material

reflectivity, welding speeds, depth of penetration, shield gas, weld design,

microstructure/strength of joint and fit-up or joint tolerance.

A brief description follows for the key-hole welding process used primarily for welding

body sheet metal.

2. Discussion

2.l. ADVANTAGES OF LASER WELDING VS. RESISTANCESPOTWELDING [2]

Since the inception of sheet metal for automotive bodies, the basic joining process. has

been continues to be resistance spotwelding, namely a two sided joining process. Virtually all

automobile bodies being produced today are designed around the requirements of the

spotwelding process. The advent of robotic laser welding portends a revolutionary change in the

basic product design of automotive bodies of the future.

In order to understand the fundamentals of this dramatic change,a detailed comparative

analysis between laser welding and resistance spotwelding is outlined. The fundamental

advantages of laser welding are summarized and described:

Consistent weld integrity;

Single sided access;

Reduced weld flanges;

Increased design flexibility;

Reduced heat distortion;

High speed progress;

Increased structural strength; and,

Non-contact process.

3. Design Flexibility of Laser Welded Joints [5]

Today's automotive body product design and joining processes are severely constrained

273

due to the limitations and two-sided requirements of the spotwelding process. These limitations

virtually dictate how the body components are joined together and, in many cases, the product

design configuration.

The advent of the robotic laser welding process opens up new horizons for both the

product designer and the manufacturing engineer. A complete new family of innovative joint

designs are outlined which provide new capabilities for material processing, and design freedom.

A key requirement for implementation of these new capabilities is simultaneous

engineering between the product designer and the manufacturing engineer, to gain information

about the potential and limitations of this new technology. Together with a common

understanding, and a joint commitment, they will achieve a successful revolution in how the

bodies of the future will be designed and processed.

4. Requirements for Successful Laser Welding Systems [3]

The key factors for laser welding sheet metal are outlined. These include metal fit-up,

coated materials, joint design, flexible and efficient laser beam delivery, system reliability and

maintenance, process inspection and very importantly for widespread use in the automotive

industry: system cost The use of Computer Aid Design (CAD), coupled with Computed Aided

Manufacturing (CAM) systems are reviewed as key elements of successful laser welding

systems.

5. Advancements in Robotic Laser Systems Both for C02 and Nd:YAG

The implementation of robotic laser welding of automotive body components and bodies

has been significantly enhanced in recent years, due to the dramatic progress in effective flexible

beam delivery systems, for production use.

The first "on-line" production body laser welding system instailed in 1985 at Linden, New

Jersey by Prima-Industries is described. Following this experience, General Motors Corporation,

in concert with GM-FANUC, developed the first robot exclusively designed for laser beam

manipulation, the GMP-LIDO Robot.

274

A complete summary of robotic developments for C02 laser beam delivery is outlined for

all major robotic manufacturers.

Recent developments for fiber-optic beam delivery of the Nd:YAG laser beam are outlined

for present power levels up to 2 kw. The tremendous new potential for robotic laser welding and

cutting are reviewed with this exciting new development.

6. Applications ofIn-Production Laser Welded Assemblies [5)

In-production use of robotic laser welding is being implemented by most of the major

automotive manufacturers world-wide. However, for a variety of reasons, the developments have

not received widespread publication.

This report outlines speciflc on-line production laser welded assemblies including:

Shelf sub-assembly;

Quarter inner sub-assembly;

Radiator support sub-assembly; and,

Miscellaneous sheet metal sub-assemblies.

A brief rewiew of existing laser work cells is outlined including overall layout, system elements

and operation, together with new concepts of quick-change tooling changeovers for laser welding

tooling systems.

7. Conclusions

7.1. LASER WELDING CHALLENGES OF THE FUTURE [2)

The implementation of robotic laser welding is at a very early state of development and

thereby provides tremendous opportunities for both product design and laser process innovations.

A detailed review of laser system equipment challenges are outlined regarding performance

targets and a system target cost of $250,000.00 for an equivalent 3 kw system, less tooling.

The concepts for future robotic laser cells are outlined to meet high volume production

output. In addition, laser processing needs are outlined in the areas of cost reduction, both capital

275

and process, reliability, serviceability, process control and education.

Future trends are summarized including laser technology trends for compact multi­

kilowatt C02 and multi-kilowatt Nd:YAG lasers. Future system developments are outlined for

sophisticated time and power sharing as well as diagnostics. Final future trends in the area of

applications are discussed regarding the refinement of current processes (joint designs and

optimized parameters) and the acceptance of new processing.

References

1. Uddin, N.M. (1990): "Laser to the Work Piece", IMTS. 2. Di Pietro, F.A. (1989): "Industrial Laser Review", Oct. Issue. 3. Vasilash, G.S. (1988): "Production Magazine", Nov. Issue. 4. Roessler, D. (1986/1987): "Industrial Laser Hand Book". 5. Di Pietro, F.A. (1990): "Proceedings of the ATA", Torino, Italy.

ROBOTIC LASER CUTTING SYSTEMS

F. A. DI PIETRO Director, Manufacturing Systems (Retired) General Motors Corporation, U.SA. 9650 South Ocean Blvd. #1510 Jensen Beach, Florida 34957 (U.SA.)

ABSTRACT. This lecture covers the thermal processes of C02 laser cutting, the laser qualities most important for cutting, and the mechanics of the cutting process. In addition, a comparative analysis of alternative thermal cutting processes, and alternative beam manipulation systems, are reviewed including new developments in Nd: Y AG laser cutting systems with fiber optic delivery coupled to a robot. Laser cutting advantages are reviewed and actual applications of in-production laser cutting systems are discussed. The major specific subjects covered include the following: History and types of C02 lasers for cutting; Description of laser cutting processes and related parameters; A comparative analysis of laser and other thermal cutting processes; Analysis of laser beam delivery and motion systems benefits associated with Laser Cutting Systems; New developments in Nd: Y AG robotic cutting systems with fiber optic beam delivery; and, Actual applications of in-production Laser Cutting Systems.

1. Introduction

1.1. mSTORY AND TYPES OF C02 LASERS FOR CUTTING

Albert Einstein, in 1917, announced his "Quantum Theory" proposing that "light with

energy of a particular frequency could simulate electrons to emit radiant energy as light of that

same frequency." The theory became the basis for laser technology; in 1960, Dr. T. H. Maiman

invented the first laser and in 1966 the first industrial laser was built.

A brief explanation of the laser spectrum, and (T.E.M.) Tranverse Electromagnetic Mode,

Le., the cross sectional shape of the working laser beam is covered with particular emphasis on

the TEMOO gaussian-curve mode that is best-collimated and produces the smallest spot size of

high power density for cutting.

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The laser qualities most important for cutting are reviewed including TEMoo mode, small

raw laser beam diameter <15 mm with low divergence <.2 mrad divergence, circular beam

polarization and accurate power shaping and modulation control. Various pulsing forms are

employed in order to regulate the output power with respect to time.

2. Discussion

2.1. DESCRIPTION OF LASER CUTIING PROCESSES AND RELATED PARAMETERS [2]

When properly executed, the laser cutting process produces excellent edge qUality. It can

be summarized as follows:

Edge accuracies to ± 0.02 mm;

Edge quality better than 3.5, umR(a);

Burr from dross less than 0.12 mm; and,

Corrosion resistance equal to drilling or piercing.

The family of industrial lasers are outlined with specific detail regarding three types of C02

lasers, i.e., slow flow, cross (transverse flow) and fast axial flow lasers; trend charts indicate the

volume of3 axis C02 laser installations since 1985-1990 between the U.S., Europe and Japan.

Parameters relative to the laser material inter-action are reviewed including thermal

conduction, melting, vaporization and plasma production occuring at various levels of light

intensity (w/cm2). The impact of power and power density, as it relates to focal length and depth

of focus are reviewed; energy density is the critical factor (millions of watts per square

centimeter).

The laser cutting process and the construction of the laser cutting head and the high speed

hole cutting end-effec are described in detail. The laser cutting process itself involves focusing a

beam of coherent light, concentrated to an extremely small spot size (less than 0.20 mm) to melt

or vaporize the sheet metal. A fine coaxial stream of air, oxygen or an inert gas is used to blow

thrugh the material creating an exothermic reaction augmenting the laser's energy. Additionally,

the gas flow serves to blow the melted material away from the cut, helping to produce a high

quality, dross-less cut. When combined with smooth movements of the beam over the work.

piece, accurate cuts with relatively small kerf widths and a minimum heat affected zone will

result.

279

The components of a complete laser cutting system consist of: (1) the high voltage power

supply; (2) the laser resonator chamber with its associated pumps, mirrors, lasing chambers and

related cooling facilities; (3) the gas control system; (4) a beam delivery system and focusing

nozzle; (5) the work piece handling facilities; (6) the machine motion system; (7) the laser and

machine controls; and, (8) the capacitance feedback nozzle to maintain the proper focal distance.

In addition, protective enclosures and fume exhaust facilities are often included in the overall

system.

3. Comparative Analysis of Laser and Other Cutting Processes [3]

A variety of alternative cutting and trimming methods are reviewed. The selection depends

upon factors such as part configuration, size, thickness, material accuracy requirements,

complexity of trim edge, effect of head on the trim edge, and of course the number of parts to be

cut. System costs and flexibility requirements are additional factors which must be considered.

The laser fits well into this spectrum of processes because it has some characteristics and

capabilities not found in other competing processes.

The laser cutting process is compared in detail with alternative processes such as plasma­

arc and the EDM process; comparing both the advantages and disadvantages of each system.

4. Analysis of Laser Beam Delivery and Motion Systems

The major systems for laser beam delivery and motion systems can be categorized into

two types: CNC Systems and Robotic Systems. A variety of CNC Systems have been developed

with moving overhead gantry systems for laser beam delivery coupled with three to five axis or

more capabilities. These systems also incorporate a variety of handling methods for the work

piece. The alternative system is a robotic system for manipulating the laser beam in relation to

the work piece. Both systems are reviewed in detail for both C02 and Nd:Y AG Laser Systems.

280

5. Benefits Associated with Laser Cutting

Laser cutting has many advantages over alternative processing, leading improvements in

cost, quality and responsiveness including just· in-time manufacturing systems.

The advantages of laser cutting include the following:

Minimal Kerf width (0.025" to 0.100");

Elimination/reduction of secondary operations;

Elimination of tool storage costs;

Short lead times -maximum responsiveness;

No tool set-up time: total flexibility;

No mechanical distortion from tool contact;

Minimum heat affected zone;

No tool wear or replacement;

Smooth cutting edge at high speeds;

Cut has parallel sides (less than 2% edge taper);

Complex contoured profiles can be cut;

Can start/stop cuts anywhere; and,

Computer capable for teaching and changes.

Laser cutting world wide is the most widely used process of all laser technologies, at this time.

6. New Developments in Nd:YAG Robotic Cutting Systems with Fiber Optic Beam

Delivery [I]

Nd:Y AG lasers offer processing rates and capabilities that are competitive with the

industry standard C02 laser systems. These systems afford the added benefit of flexible of

flexible fiber optic beam delivery and, coupled with the robot, they offer cost/effective flexible

system capabilities. There are several advantages of Nd:YAG lasers over C02 lasers. The 1.06

/lID wavelength of Nd:Y AG lasers is absorbed in metals more than the 10.6 /lm wavelength

produced by C02 lasers. For multiple-beam delivery requirements, Nd:YAG lasers can also be

energy-shared by dividing the beam and sending portions of the energy simultaneously down

several fiber optic systems. Alternatively, the Nd:YAG laser beam can be time shared by

281

shuttling the full power of the beam consecutively down several fibers. Finally, Nd:YAG lasers

are solid state machines that have demonstrated reliability with production records of 98+

percent uptime.

Fiber optic delivery is one of the greatest advantages of Nd:YAG lasers have over C02

lasers. The use of robots for beam delivery manipulation in automotive and aerospace

applications is a significant motivation for the development of (Nd:YAG) laser systems.

The Nd:YAG laser operating parameters are discussed together with fiber optic delivery

systems. The coupling of the robotic capability with the Nd:Y AG laser beam fiber optic delivery

offers new capabilities which are only now being developed for many industrial applications.

7. Conclusions

7.1. ACTUAL APPLICATIONS OF IN-PRODUCTION LASER CUTTING SYSTEMS

Several in-production examples of robotic laser cutting are discussed including cutting of

floor pans for automotive applications using a 500 W Nd:YAG laser coupled with robot and

fiber optic delivery system; similar applications are described for both Japanese, European and

American automotive manufacturers.

Another application of these systems is for "Post Assembly Cutting" on completed

automobile bodies. This application portends tremendous benefits for the automotive industry to

implement "Just-in-Time" manufacturing systems and related benefits.

Finally, laser challenges of the future are analyzed in tenns of infonnation understanding,

commitment and action.

282

References

1. Bransch, H. and Kugler, T. (1990): "Proceedings of International Conference of of Lasers '90" -pp325-330.

2. Daly, D. (1990): "Lasers to the Work Piece" - INTS 90, Chicago, Ill. 3. Simpson,G. and Churin, T. (1990): "Focus on Lasers" - April Issue. 4. Bransch, A. (1991): "Photonics Spectra" Magazine, pp 107-111.

DEVELOPMENT OF LASER MATERIAL PROCESSING IN ROMANIA

I. I. FARC;AS Laser Department, Institute of Atomic Physics p.o.B. MG-6 76900 Bucharest-Magurele, Romania

ABSTRACf. C02 and solid-state lasers have been developed at the Institute of Atomic Physics In Bucharest-Romania since the late 60's. Early attempts to use them in material processing came soon afterwards. In the last ten years, DC excited lasers of 500W and 1.5kW and 2-D laser cutting units have been built. Some results concerning laser cutting and hardening of metals and other materials are shown. A laser robot of the gantry type has been installed and is now under testing.

1. Historic and general remarks

Laser research started at the Institute of Atomic Physics (lAP) soon after the first lasers

were reported. In 1962, a He-Ne laser was operated, emitting on the 1150 nm line [l]. In the

following years, most of the lasers known at the time were built and operated, such as C02, Nd

glass, He-Ne, Ar+, Dye. Specific applications have been developed, as may be seen in Table 1.

Table 1 Laser applications.

C02 Nd-glass; Nd-Y AG He-Ne Dye (pumped by N2 laser)

Material processing, medicine Material processing, medicine Holography, interferometric measuring instruments, alignment devices Monitoring of atmospheric pollutants

The first C02 laser was operated at lAP in 1967 at a power level of lOOW . 15 laser units

have been built and sold to different customers encompassing educational, research and

industrial institutions. In 1970 a 600 W laser was built and used for the first experiments in

material processing, a starting point for a steady activity.

283

S. Martellucci et al. (eds.), Laser Applications for Mechanical Industry, 283-290. © 1993 Kluwer Academic Publishers.

284

2. C02 laser development

In the section a short survey of presently available C02 lasers is presented.

Research in the field yielded various types of C02 lasers covering a power range from a

few watts to over 1 kW.

For the lower part of the power range, an all-glass, sealed-off C02 laser has been

developed. According to the discharge length, the output power varies from about 3 W up to over

100 W [2]. The mirrors are attached to the discharge tube, as may be seen in the longitudinal

section shown in Fig 1. Lifetime of this device exceeds 2000 hours. Such laser tubes of 10-20 W

are used in surgical C02 laser units. About 10 such units,built in our department, are now in use

in different hospitals in Romania. Laser tubes of 50·100 W are used in research concerning

material processing, photochemistry, biology and resistor manufacturing.

For mid-range power, a folded-type laser was developed. It is a 4 parallel tube laser, with

slow rate gas flow [3]. A sketch of this laser is shown in Fig 2. Operating in the TEMOl mode, it

gives over 500 W of output power.

At the upper end of power range is the gas transport laser [4]. As it can be seen in the

cross-section shown in Fig 3, the laser is contained in a cylindrical enclosure. A sequence of

tangential ventilators fixed on a rotating shaft moves .the gas along the outer wall. Gas excitation

occurs in a DC discharge, as in the case of the previously discussed laser. It develops between a

single cathode and a segmented anode, each segment being individually ballasted. Both the

specially provided heat exchanger and the water cooled enclosure wall contribute to the gas

cooling. The laser can operate in multi·mode or single-mode beam configuration, according to

the resonator design, the output power being lower in the latter case. In multimode operation, the

rated output power is 1.2 kW but a maximum output power of 1.5 kW can be attained.

/JVr,rE,,-f> O/SCHAHGE V(JI (l1'ft!" A TU« I<

/ r _ / lit, r / 11'- L 1 ---( 0-=.- ~-..:::rr-

(lff \. \ ) -~

\ CQ4l/pG K J.-9CKU

Fig. l. Sealed-off laser tube structure. OM - output mirror; EM - end mirror; A - anode; K - cathode

(J45 UII.9(/,fT

285

Fig. 2. Schematic diagram of the 500 W folded C02 laser. A - anode; K - cathodes; M - mirrors; P­prisms.

Significant parameters are: Power stability = +/- 7%; Divergence = 5 mrad; Size = 1920 mm (L),

860 mm (W), 1650 mm (H); and, Gas consumption = C02 - 15l/h, N2 - 100 l/h; He - 1501/h. In

an experiment with two such lasers operating in tandem, an output power of 2.4 kW has been

obtained.

2

8

7

Fig. 3. Cross section through the gas-transport C02 laser with cylindrical geometry. 1 - metallic cylinder; 2 - insulating plate; 3 - shaft; 4 - fan; 5 - heat exchanger; 6 - anode assembly; 7 -auxiliary anode; 8 - cathode; 9 - electrical discharge; 10 - copper tubes; 11 - metallic cover; 12, 13 - baffles.

286

Fig. 4. Cutting speed dependence on thickness for different materials.

6

I

rt/wt;(' : .w It! RXq/ ~1?",( 6J mm

GIlS : 02 -/or l'7t'k4-

~- ~mt'Ar4 , J'I'O',;.,4o.l".r 'k~ 2 O'#QJ'~o' &Tt't'l :1 l'O',6o/) .r1t'~1 4 J>1('~t9./0'.r.r 5 WDOO'

6*m~

Fig. 5. Cutting speed dependence on laser power and plate thickness for carbon steel.

287

3. Laser material processing experiments

The last two types of lasers have been developed mainly to be used in material processing.

Experiments have been perfonned in order to establish the processing capabilities of these lasers.

The cutting perfonnance of the folded laser has been investigated and resulting data are shown in

Figs. 4 and 5 [4]. In Fig. 4 the cutting speed variation for different materials is shown as a

function of laser power. The cutting speed of carbon steel sheets of different thicknesses versus

laser power is represented in diagrams in Fig. 5. The gas transport laser has been used in cutting

and hardening experiments, the results being shown in Figs. 6 and 7. The diagram in Fig. 6

shows the depth of the hardened layer variation for three types of steel versus laser power, at a

spot displacement speed of 1.5 cm/s. For the same steels, the hardness variation in the depth of

the hardened layer ia shown in Fig. 7.

Fig. 6. Hardened layer depth dependence on laser power for three steel types.

/lY

{},! -

tim IJ/JIJ

Fig. 7. Depth variation of the hardness for three steel types.

• o 9f,.IIot'rAJ'Id" .:l t?i('-i5 K IlLtoO V::/,scm,4

o tJ,l 0,$ .l(mtn)

288

4. Laser processing units

The C02 lasers discussed above have been developed for material processing purposes.

Developing processing units therefore appears as a natural prolongation of this activity . Because

this was beyond the capabilities of our Institute, other specialised organisations have joined in. A

joint venture of the LA.P. and the Mechanical Engineering and Research Institute in Bucharest

has resulted in two C02 laser cutting machines, each being able to perform 2-dimensional

operations.

The first one uses the folded 500 W C02 laser and can be seen in Fig. 8. This machine

operates with a fixed beam, the processed sheet being moved through the focus of the beam. The

laser is placed above the moving table, while its supply unit is plugged into the body of the

machine. The processed sheet is fixed onto the moving table, which is driven by DC electric

motors in the X and Y directions through ball-screw shafts. The table can be moved over a range

of 1100 mm (X) and 800 mm (Y) with a continuously variable and programmable speed of up to

8 m/min. The focussing head can move along the Z axis over a range ofl50 mm. Its height above

the processed sheet is automatically maintained by a closed loop system using a capacitive

sensor. A CNC system controls the table displacements and speeds along the X-Y axes, the beam

shutter, the flow of the assistance gas and the lowering and lifting of the focussing head at the

beginning and end of cutting sequences. The contouring precision is better then +/- 0.1 mm/m.

Two such units have now been built; one of them is in operation in a bus plant, the second is to

be installed in a tractor plant.

Fig. 8. 2-D cutting unit with folded 500 W C02 laser.

289

Fig. 9. 2-D cutting unit with 1.2 kW gas-transport C02 laser.

The second model of cutting unit, developed to operate with the gas transport laser, can be

seen in Fig. 9. It uses a different concept: the bending mirror is displaced along the laser beam

direction (Y-axis), while the processed sheet is moved along the X axis. The laser is placed

beside the processing table on a common chassis. A similar CNC control system as in the

previously described case is used, and also a similar adjusting system for the focussing head

position. The main characteristics are: Displacement range - along the X axis 2500 mm , along

the Y axis 1500 mm; Speed range - 0.1- 8 m/min, continuously adjustable and programmable

over both axes; Focussing head positioning range - over Z axis 400 mm; Contouring precision

+/- 0.1 mm/m; Laser rated power - 1200 W; Mode pattern - TEM OJ; and, Beam divergence < 5

mrad.

Fig. 10. Gantry type laser robot.

290

The first such cutting unit is now under testing and it is intended to be used to cut large

pieces in metal plates up to 10 mm in thickness.

Very recently a gantry type robot has been purchased from ROPOS Comp., Baia Mare,

Romania and was installed in our laboratory. It is a modified version of a welding robot whose

operating head has been redesigned in order to handle a laser beam. Four bending mirrors are

used to transport the laser beam over a variable distance ranging between 2 and 8 m. The robot

itself provides 5 degrees of freedom: 3 traslations (along X,Y,Z axes) and two rotations, the first

around the Z axis, the second around any direction perpendicular to the Z axis. If needed, a third

ratation can be provided by a special workpiece holder. An overall view of this unit appears in

Fig. 1 O. The main characteristics of the robot are summarised in Table 2.

Table 2. Characteristics of the robot.

Displacement Speed

X-translation 3200 mm 0.23 mls Y -translation 2300 mm 0.23 mls Z-translation 450 mm 0.23 mls a-rotation +1- 154 deg 90 deg/s ~-rotation +1- 110 deg 90 deg/s clamping device rotation +1- 360 deg 20 deg/s

Repeatability precision better than 0.1 mm has bean obtained. This robot is intended to be

provided with a gas transport laser of 1 - 1.5 kW output power, operating in single mode.

The main purpuse of this activity is to build an experimental facility to be used in the

laboratory for material processing experiments. In case of successful operation, a job-shop may

be set up.

References

1. Agarbiceanu, I., Agafitei, A., Blanaru, L., Ionescu-Palas, N., Popescu, I.M., Vasiliu, V. and Velculescu, V. G. (1963): "Contribution to the study of gas lasers" - Report at the 3-d Congress of Q.E., Paris.

2. Axinte, C., Farcas, 1., Gutu, I.and Draganescu, V.(198l): "Laseri cu C02 inchisi" Studii si Cerc. Fiz, Vol. 33, Nr. 5, p 506.

3. Draganescu, V., Farcas, 1., Gutu, 1. and Axinte, C. (1986): "C02 lasers and their applications for unconventional material processing" Rev. Roum. de Phys. Tome 31, Nr. 6, pp 579-587.

4. Axinte, C., Comaniciu, N., Draganescu, V., Fargas, 1. and Gutu, 1. (1978): "Gas transport C02 laser with cylindrical geometry" Rev. Roum. de Phys. Tome 23, Nr. 5, pp 447-456.

OTHER APPLICATIONS: STEREOLITHOGRAPHY, SURFACE HARDENING,

3-D SURFACE PATTERNING, MANUFACTURING AND REPAIRING

THE ROLE OF THE LASER IN RAPID PROTOTYPING

L. PERA, G. MARINSEK FIAT Research Center Strada Torino SO 10043 Orbassano (TO), Italy

ABS1RACT. Modem industry is faced with the need to improve or update products in a short time to be able to defend or increase its market position. A new field of technology is recently emerging and the laser plays a fundamental role in its success. This technology, named "Rapid Prototyping", allows the construction of component models in a minimal amount of time. The objects are created, not by material removal as in traditional CAM technologies, but by increase of material, layer by layer or point by point. In some cases the laser selectively solidifies or binds solid particles of different kinds of materials; in other cases the laser hardens any of a wide family of liquid photopolymers. Components made with these techniques show good surface quality, accuracy and suitable mechanical properties. A rapid evolution is expected in the very near future, as in the case of copy machines. Technical experts in mechanical industries, to reach their objectives of cost, quality and short time to market, will need not only a 2D printer connected to their CAD systems, but, let us call it, a 3D printer that generates real components, to quickly evaluate the characteristics of their products.

1. Introduction

The market challenges modem industries. On one hand, industries must produce more

variants of more complex products, with additional design and engineering effort. On the other

hand they have to face decreased product lifetime and shorter delivery times (Fig.1 ) [1] . Time,

in terms of required delivery time and time management, is essential to stay on the cutting edge

in today's world markets. Rapid Prototyping (RP) refers to a range of new technologies for

making three-dimensional prototype parts in hours or even minutes. In comparison with

conventional techniques, RP allows to more rapidly realize a component, ready for the evaluation

of accuracy, fitness, mechanical behaviour. In this aspect RP can be seen more as a management

293

S. Martellucci et al. (eds.), Laser Applications/or Mechanical Industry, 293-303. © 1993 Kluwer Academic Publishers.

294

decision tool in the design and product development process rather than an actual machining

operation. The first RP technique was born less than five years ago, and the main factor in its

success was the laser. In fact, RP gains significant advantages from such laser characteristics as

flexibility, velocity, automation, and dosable delivery of energy. This chapter aims to draw an

up-to-date picture of RP technologies and to understand the laser's role in their evolution.

~ _______ N_u_~_cr_O_f_V~_.M_ffi ________ ~1 ~1 ________ ~_oo_ucr_l_ire_ti_~ ______ ... 1

1970 1980 1990 1970 1980 1990

Product complexity Delivery time I

1970 1980 1990 1970 1980 1990

Fig.I. Market requirements: number of variants; product lifetime; product complexity; delivery time.

295

2. A new tool for design and production devepolment: Rapid Prototyping

Up to now, manufacturing techniques have been classifled in two groups, according to

how the component is generated: a) Fonning processes to defonn bulk material into the required

shape with no material added or removed. It can be done "in solid state" (forging, stamping,

drawing, extruding, ... ) or "in liquid or semiliquid state" (casting, injection moulding, ... ); and, b)

material removal processes to remove excess material in conventional machining (turning,

milling, grinding, .. ) or in non-traditional machining (EDM, ECM, ... ). The new group of

techniques collected under the name of Rapid Prototyping (RP) are quite different from the two

conventional methods, because the shape is obtained by adding and not by removing material.

More than 25 different RP systems are present today on the international scene, but fewer than

10 of these are industrially robust. Some other systems are limited to subcontracting and

servicing using in house equipment, while many others are still in a development phase [2 - 5].

Two classifications are possible for RP techniques. The first (Table 1) relates to the initial state

of material (liquid, powder, solid), the second (Table 2) to the building technique.

Raw material can be generally used in the following three different states:

SQlid - Several processes ("Laminated Object Manufacturing" from Helisys and Hydronetics)

glue foil by foil one on top of each other in order to produce the required shape. Semi­

polymerized plastic foils can be bound together by further photopolymerization. Although

industrial parts have been already made by these methods, it is a general opinion that they could

be further improved. Building up the model can be done directly in a 3D space or in successive

2D layers. The given classification is however rather theoretical. As a matter of fact, all the

available systems are based on 2D logic, where the CAD model is sliced in layers. Even if a

direct 3D technique does not require the complete creation of the lower part of a component with

a consequent larger flexibility in complexity of shapes, programming considerations make layer

by layer execution easier, faster and less expensive. The layer is processed either all at once (one

layer at a time) or by scanning the surface point by point. This scanning can be discrete or

continuous.

~ - The sintering of powder is obtained via laser melting of the contact areas of grains

(DTM's "Selective Laser Sintering") or glueing them together by selectively adding a binder or a

glue (,,3D Printing" of MIT). A future alternative could be to start with a two component powder

and selectively activate their binding. While Selective Laser Sintering is ready for industrial

applications, 3D printing is still in the research phase.

1imlli1- This category includes most of the industrialized systems. They are based on liquid

photopolymer solidification under ultraviolet light. The source of energy is laser (3D Sys, EOS,

296

Table 1.

SOLID

POWDER

LIQUID

RP techniques classification related to the state of material.

GLUEING SHEETS

1 component

1 component + binder

2 components

MELTING + SOLIDIFICATION

LIQUID POLYMERIZATION

-Laminated Object Manifacturing (Helisys, Hydronetics) -Selective Laser Sintering (DTM, Hydronetics, Westinghouse)

-3D Printing (MIT)

-Shape Melting (Badcock & Wilcox, Jero)

-Fused Deposition Manufacturing (Stratasys. Perception System)

-Ballistic Particle Manufacturing (Percetion System, Automated Dynamics Co) LIGHT SINGLE W A VELENGHT LAMPS -Stereolithography (Light Sculpturing Cubital)

LASER BEAM -Stereolithography (3D Sys, Quadrax, Grapp, Du Pont, Laser Fare, Sony/D-Mec, Mitsui E&S, Mitsubishi/CMET, EOS)

HOLOGRAPHY -Holographic Interference Solid (Quadtec) LIGHT TWO WAVELENGTHS -Beam Interference Solid (Grapp, Battelle, Formigraphics) HEAT -Termal Polymer (Grapp)

Mitsubishi/CMET) or lamps (Cubital). This kind of application goes under the name of

"Stereolithography" from the first developed system, that is, from 3D Sys.

Melting+Resolidification - Systems based on melting, deposition and resolidification of material

allow one to use plastics, resins, waxes, metals; the best results regarding this class are those

obtained with plastic and wax wires (Stratasys). Here the filament is melted passing through an

extruding heated nozzle. This process is called "Fused Deposition Modeling".

297

Table 2. RP techniques classification related to the building technique

Surface by surface -Holographic Interference Solidification DISCRETE -Beam Interference Solidification

3D Point by point -Ballistic Particle Mfg CONTINUOUS -Fused Deposition Mfg -Shape Melting

Surface by surface -Laminated Object Mfg -Stereolithography DISCRETE -Stereolithography

2D -Thermal Polymerization (Layer) -Selective Laser Sintering

Point by point -3D Printing CONTINUOUS -Fused Deposition Mfg -Shape Melting -3D Printing

3. How the CAD file becomes a prototype

It was previously stated that up to now all the "RP" Teclmiques work on a layer by layer

basis. Thus a general procedure for obtaining a model from a CAD file can be identified. In Fig.2

a typical process path is presented. The first requirement is a CAD model of the component. It

can be generated from any of the commercial CAD systems present on the market. CAD CAPP CAM-PROCRAMMING Pm desian - pin orieRlatiun • slicin,

_ de£ign of supports . - scan path generation • conven to S1t. format

A

from A -t

POST-PROCesSING - POll curini • suppon removal • c\eanins

OVEN

Fig.2. Typical scenario of software and machining operations for a "Rapid Prototyping" process.

298

However, solid programming is preferable to surface programming since less time is required in

defining part orientation and in trimming coupled geometrical figures. It is worthy of note that

for some RP processes it is necessary to model some kind of supports for the overhanging parts.

The second step is to convert the CAD model to an STL file. The STL interface, developed by

3D Sys for the stereolithographic process, has already become a standard adopted by all suppliers

of RP machines. This PHIGS-based format approximates the part's outer and inner surfaces by a

set of triangular flat patches. From this file it is possible to slice the component model at

distances equal to the layer thickness. The sofware is not general, but it relates to the specific RP

system. The new file (SLI) pilots the movements of the "RP" machine. After the component is

obtained, in general some post processing is required. This phase is largely dependent on the RP

technique and it varies from a simple cleaning of the prototype to a finishing mechanical

operation or to a post curing for further solidification of the part.

4. A closer look at RP laser processes

In order to understand how much the laser is involved in RP techniques, the following

Table 3 presents a list of the principal laser-oriented RP systems together with their

characteristics.

Lasers in RP are mainly used in Stereolithography, where they dominate the scene. As

this application was the first to appear on the market, it also provides at present the best results,

especially as regards accuracy. The "stereolithography" process is described in Fig.3. The core of

the process is the hardening of a liquid photopolymer by an ultraviolet laser. The laser triggers a

chemical chain reaction (photopolymerization) which changes the liquid to a solid. The extent of

hardening (curing) is controlled by the amount of energy imparted to the resin and the result is a

hard polymer, usually similar to plexiglass [6]. The laser draws all the sections of the component

at the surface of the liquid. The solidification follows exactly the drawing in the xy plane (liquid

plane), while it is limited in the z direction to a fraction of a millimeter. This layer automatically

sticks to the previous one. The component is built on a platform. At the beginning the platform is

placed below the liquid surface of about one layer thickness. After the laser scanning and

solidifying phase, this platform is lowered a distance equal again to one layer thickness; the

liquid flows in to cover and the last solidified layer. The process is repeated until the model is

complete. A support structure is sometimes added to the model in order to prevent that

overhanging parts from floating or dropping down in the liquid. These supports are designed

Tab

le 3

. L

aser

ori

ente

d R

P s

yste

ms.

Bea

m

3D

Dup

ont

Qua

drax

in

terf

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Sys

tem

so

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tion

M

ater

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Pho

topo

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topo

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The

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(UV

) (U

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phot

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ates

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Mac

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ax s

ize

300x

300x

300

S08

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30

Sx30

Sx30

S 30

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5

Sta

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20

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(mod

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830x

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500

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He-

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100-

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yes

yes

yes

none

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(15

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400

mm

/s

10

00

0m

m/s

3 s

1 -

2 s

Hel

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Foi

ls

-pap

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-fa

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flm

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no

381

mm

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DT

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cs

(AB

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PV

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PC, n

ylon

) -W

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s

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380

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2

10.6

flil

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no

no (

exc

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ics)

1016

mm

/s

0.2

-2

s tv

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) \C

)

300

with software, usually with the CAD itself, and are processed and built together with the object.

After curing they are mechanically removed. An experienced user is able to design this support

structure in order to remove it easily without damaging the component. The laser beam is

delivered to the liquid layer by mirrors or optical fibers and focusing lenses. In order to maintain

the same level of curing effect allover the working table it is of great importance that no change

occurs in the spot size and energy density. Optical fibers are more suitable for the job, but with

their use the scanning has to be performed by xy plotters that are slower in comparison with

oscillating mirrors.

The new field for application of lasers in RP is Selective Laser Sintering. As shown In

FigA, the basic of processing is quite similar to Stereolithography. The main differences between

the two. all in favour of the sintering approach, are the lack of supports and the possibility of the

greatest diversity of materials. After the solidification of each layer the deposit of another layer

takes place. However, melting powder is more difficult than curing photopolymers. In fact, heat

diffusion depends on the material and it is more troublesome to define the right amount of power

needed for the process. Thus two different problems arise, the first being related to the degree of

sintering of the powder and the second to the minimum section width. As a matter of fact,

models from ceramic material can be obtained only in a green state and a post oven treatment is

I j computer controlled moving mirrors j;!!!!!!~:!!!!! optical system --- LASER

j j

liquid or powder external supply ~ oscillating beam

Fig.3. Scheme of Stereolithography.

301

required, and special metallic powders need to be developed to obtain an acceptable mechanical

behaviour. Up to now the best results have been with plastics and waxes.

In Laminated Object Manufacturing (LOM) the laser is again the core of the process. Parts

produced by LOM consist of a stack of foils cut to shape and glued together. Each foil is first

glued to the stack, before the part section contour is cut out by a laser beam (Fig.5). The velocity

and the focus of the laser beam is adjusted so that the cutting depth exactly corresponds to the

thickness of the foil and the underlying layer is not damaged. Virtually any material can be

applied: cellulose (paper), metals, plastics, fabrics, composites. This process uses glue coated

foils, but in the future a heat sensitive glue will be applied. In this way, the same laser source

will first activate the glue only in the cross section that must adhere together, and afterwards the

laser will cut the contour of the part at a lower speed.

Besides Stereolithography, Selective Laser Sintering and LOM (Laminated Object

Manufacturing), another laser RP is Beam Interference Solidification. This technique applies to a

transparent liquid plastic (photo-sensitive monomer) contained in a transparent vat. Part building

occurs by a point by point solidification of the liquid at the intersection of two laser beams with

different wavelengths (Fig.6). All liquid hit by beam I is excited to a reversible metastable state

that polymerizes only upon impact of the light of beam 2. This process, developed by Battelle

and Fonnigraphic among others in the sixties, did not find industrial applications because of

some problems like excessive beam absorption inside the liquid, shading effects, and beam

intersection uneveness due to diffraction caused by local temperature variation or solidification.

Mirron ~ -'''h

Scanning C02 Laser Beam in t ... , .--1 :,~ I'

I

! ; I

i

... , -:;;: ..... . r~'" ~ .;-.. ~,

: ~

, " ~ Powder Cartridge " . f«ding System

PO¥I Cylinder And Powder Bed

Fig. 4. Scheme of Selective Laser Sintering.

302

Laser

x-v positioning ---~~i~~~~)1~~ device Heated roller

layer periphery ---I':;:::.,.;-'HttJ;;i±l

and croSShaIC:lh'-_-f,~~III'~~~~ Part· block -

~~~--.

Computer

Disk drives

Material -----Ill; ribbon

Rewinding roll - Unwinding roll

Fig.S. Scheme of Laminated Object Manufacturing.

Two photon N photopolymer ... :;:

r-----~~~--~--~

. - . . . "-'"=~'T-'

: ~ ;; ... ~'. ~1..;~ :~ )i(.;:: .. ·~;tt; ;::~·!~·;~!./:'.!::~; .~~:~;~;;:~; . . :';; ~'!~:~.~~~:~;

r,'

:·f . ' .

. ; . .. ' .......

Fig 6. Scheme of Beam Interference Solidification.

5. Conclusions

It is by no means necessary that RP processing involves the use of the laser. However, as

shown previously, more than 75% of RP techniques take advantage of laser characteristics. This

field is open to rapid evolution in the next years and laser developments will be crucial for an

important step forward.

303

The future needs of laser RP are: a) Build larger components; b) Build components from a

greater variety of materials, including metals and ceramics; and, c) Improve dimensional and

mechanical characteristics.

To reach these goals from a laser point of view, we are in need of: I) Fast moving beam

delivery systems able to speed-up the process while maintaining beam quality on a larger

working volume and at higher power (up to 1000 W at 10.6 jlIIl); 2) Better laser beam quality, to

allow a focal spot volume with regular energy density distribution; and, 3) Small sources to

allow the development of compact systems more compatible with modem offices, rather than the

traditional manufacturing environment.

References

1. Mieritz, B. (1992): "Layer Manufacturing Technique as a Management Tool in Design and Product Development" - COMETT Project Proposal Strand C.

2. Kruth, J. P. (1991): "Material Increase Manufacturing by Rapid Prototyping Techniques" Annals of the CIRP VoI.40/2/ pp.603-614.

3. Lindsay, K. F. (1990): "Rapid Prototyping shapes up as Low Cost Modeling Alternative" Modern Plastic International pp.75-78.

4. Ashley, S. (1991): "Rapid Prototyping Systems" Mechanical Engineering pp.34-43 5. Carts, Y. A. (1990): "Lasers prove Integral to Desktop Manufacturing" Industrial Laser Review

pp.17-18. 6. Cabrera, M. (1991): "Industrial Applications of Stereolithography" EOS GmbH.

LASER SURFACE MELTING OF BEARING STEELS

R. COLA<;O, R. VILAR Instituto Superior Tecnico and CEMUL Material Engineering Department Av. Rovisco Pais 1096 Lisboa, Portugal.

ABSTRACT. Since the mid-1970s, several authors have reported the ability of laser surface melting (ALSM) to improve the tribological properties of tool steels. This improvement seems to be due to the toughness combined with the high hardness achieved by grain refinement and dissolution of large carbides in a surface layer. However, this improvement in the characteristics of tool steels depends on controling their microstructure and, therefore, their solidification path, by controlling the chemical composition and the laser treatment parameters (power density and dwell time). In this chapter we report some results from a microstructural study of AISI 440C tool steel for bearing applications, after treatment by LSM with various parameters.

1. Introduction

Heavy duty bearings are generally fabricated from heat treated tool steels. Thcy are

submitted to intensive rolling wear, a situation where the prevailing wear mechanism is surface

fatigue [l]. It is well known that the rolling wear resistance of high-alloy bearing steels depends

critically on their microstructure, increasing with decreasing volume fraction of carbides [2] and

with increasing hardness of the matrix [3]. The influence of retained austenite on the wear

resistance remains controversial, some authors finding that austenite decreases the wear

resistance of the material [4], whereas others encountered the opposite behavior [5], which

probably means that there exists an optimum content of retained austenite for maximizing the

wear resistance. The mean size and the size distribution of carbide particles also have an

important effect on the surface fatigue resistance of the material. In martensitic steels with

similar matrix hardnesses, the surface fatigue resistance decreases with increasing mean carbides

305

S. Martellucci et al. (eds.), Laser Applications for Mechanical Industry, 305-314. © 1993 Kluwer Academic Publishers.

306

a) b)

Fig. 1. Microstructure of the material in the as recived condition. a) Annealed; b) Quenched and tempered.

size. Large carbide particles have a particularly detrimental influence since they nucleate cracks

and promote materials removal by pitting, because they are prone to fracture under repeated load.

Since LSM leads to extremely fine microstructures with a carbides fraction lower than in

conventionally processed steels, it has a great potential as a tool for producing superior bearing

materials [6-9]. However, to achieve this goal, the austenite and, for some compositions, 0 ferrite

retained in the matrix in a metastable condition must be eliminated. Strutt and co-workers (8)

showed that 0 ferrite can be eliminated by reheating to a suitable temperature. Also, retained

austenite can be transformed by tempering [7, 8, 9], a treatment that provides the additional

advantage of further improving the hardness, by inducing secondary hardening. In this paper, we

present preliminary results of a study that aims to develop novel processing techniques to

optimize the microstructure of 440C martensitic stainless steel for heavy duty bearing

applications.

2_ Experimental

The steel used in this study is a commercially available 440C martensitic stainless steel,

with two different initial conditions: annealed and tempered after quenching. It contains I % C

and 17% Cr. The dimensions of the samples were 25x15xl5 mm. Two surface finishings were

applied: fine grinding and polishing down to 1 pm diamond. Irradiances of 400 and 500 W/mm2

307

Fig. 2. Microstructure of the transverse cross-section of a multiple pass laser melted sample.

and dwell times between 0.05 and 0.18 s were used for the experiments. The laser techniques and

the methods of microstructural analysis have been described elsewhere [10, Ill.

3. Results and discussion

In the annealed condition, the steel presents a spheroidal structure consisting of particles of

M23C6 and M7C3 carbides, dispersed in a matrix of ferrite (Fig la). After hardening and

tempering the matrix is fonned by strongly etched martensite, whereas the carbide particles of

M23C6 remain unetched and stand prominently (Fig Ib). Fig. 2 shows a typical microstructure

of the transverse cross-section of multiple pass laser melted samples. It presents four distinct

zones, depending on the maximum temperature attained during the laser treatment: the base

material (A), a heat affected zone (B), a transition zone (C) and a dendritic zone CD). In the base

material the maximum temperature remained below the austenitising temperature and the

material was not affected by the laser treatment. In the heat affected zone only solid-state phase

transfonnations occurred. Near the base material, ferrite transfonned into austenite, which

dissolved some carbides. Austenite transfonned into martensite upon cooling. The microstructure

of this region consists of martensite, undissolved carbides and, eventually, retained austenite.

The transition zone corresponds to a region of the sample where the maximum temperature lies

in the L+"( or L+b two-phase spaces of the C-Cr-Fe phase diagram. Since carbon lowers the

308

Fig. 3. Microstructure of the transition zone.

Fig. 4. Dendritic structure in the central region.

Fig. 5. Microstructure of a sample treated at 4 mmls near the fusion line.

309

melting point of the steel, localized melting started around large carbide particles that were

partially dissolved during heating, creating large local concentrations of carbon and alloying

elements in the austenite. The resulting microstructure is shown in Fig. 3. Globular regions with

a cellular structure, which resulted from the solidification of the liquid, are dispersed in an

unmelted martensitic matrix. In the central region, after complete melting, the material solidified

in a dendritic structure, grown epitaxially on the substrate. Its microstructure appears as shown in

Fig. 4. The deeply etched dendrites seem to be formed by a single phase, containing a dispersion

of small carbides. The dendrites are surrounded by an extremely fine lamellar structure,

constituted by a slightly etched carbide and, probably, retained austenite caused by the

monovariant eutetic reaction L -> "I + M7C3 [12]. X-ray diffractograms of laser melted samples revealed that the alloy is formed mainly by retained austenite. The relative intensity of (200)"1

diffraction peak is much higher than expected, denoting a strong solidification texture. The other

peaks have very small amplitudes and correspond to martensite and M3C. In some samples,

traces of M7C3 were observed. The volume fractions of retained austenite, carbides and

martensite are about 90%, 6% and 4%, respectively. In some samples treated with 4 mmls

scanning speed, the dendrites near the fusion line present a lightly etched core surrounded by a

deeply etched phase (Fig. 5). The phases present cannot be conclusively identified based on the

microscopy results alone. However, the comparison between the motphology of this structure

and those found in similar steels by Hegge et al. [13] suggest that an incomplete peritetic

reaction takes place and, therefore, the core is formed by metastable S ferrite and the periphery

by metastable "I. This kind of microstructure is observed in a layer with a thickness of about 30

mm near the fusion line. Closer to the surface the dendrites are uniformly etched and the

microstructure is similar to that shown in Fig 4. The hardness results obtained on samples treated

with an irradiance of 405 W/mm2 and an interaction time of 0.05 s, shown in Fig. 6, are typical

of the microhardness profiles found in laser treated 440C steel. The hardness of the melted wne

(5OOHV) is low when compared with the hardness of the conventionally treated steel (6ooHV)

[13] but larger than the hardness of the annealed steel (300HV). Besides, it does not depend on

the previous structure or on the depth. The hardness of the heat affected wne varies between 600

and 650HV. The low hardness of the laser treated material is explained by its extremely high

proportion of retained austenite. Therefore, tempering treatments were used to transform the

austenite, improving the hardness of the material. The study was performed on samples

previously treated with an irradiance of 496 W/mm2 and an interaction time of 0.08 s, which

were submitted to two-stage tempering treatments at temperatures between 500 and 700 0C.

The evolution of the microstructure during tempering is better understood by comparing

the X-ray diffractograms obtained from those samples. The X-ray diffractogram prior to

3\0

Fig 6. Hardness variation as a function of depth for different initial conditions.

~ 60 ~ Cii !;. 40

~ '0 c: 20 1!l .:

Laser melted

o o ~

0 ..

If

22 24 26 28

Diffraction angle (0)

(a)

800

700 N c:i > 600 ~ III III Q) 500 c:

"E os

.r: ~ 400 0

~ 300

200

30 32

100r----r---.----rr---r---,----,

.g 80 :J

CD ,~ 60 iii Cii !;. 40

~ 'iii c: '" 20

E

Tempered at 640°C

~ 0 .. ~

0'" 0

i If'"

Diffraction angle (0)

(e)

32

0 200 400 600 800 1000

Distance to the surface (!Lm)

Q)

,~ 60 iii ~ ~ 'iii c:

40

2 20 c:

20 22

o o

o~

~ .. ~

24 26 28

Diffraction angle (0)

(b)

o o

"'~

30 32

100~--~--~----~--~--~----'

60

40

20

Tempered at 700 °C

0'

0: ,I • 'I ~ !!

NI! ~ ~ g (oj 0= ! N :

~~ ~C) 0: o~ i 1 0 ~ u.. 0 i i M LL

._"-"- i L, ~ .. _.,...I\-oL-~~~~~~~~~~~~~

20 22 24 26 28 30 32

Diffraction angle n (d)

Fig. 7. Evolution of X-ray diffractograms with tempering,

311

tempering shows intense peaks of austenite, Fe3C and martensite (Fig. 7a). Traces of M7C3

appeared in some samples. Still, this carbide was not detected in the present sample. Tempering

at 500 °C does not modify the structure noticeably (Fig Sa). Conversely, after tempering at 600

°C, the intensity of the austenite diffraction peaks decreases remarkably and those of martensite

and M3C increase (Fig 7b). The proportion of austenite is now about 30%. The micrograph of

this sample showed in some areas an intense precipitation of fine carbide particles within the

dendrites (Fig Sb). Increasing the tempering temperature up to 640°C completely transforms the

austenite (Fig. 7c).Body-centered cubic a ferrite is now the main component of the structure.

Also, the relative intensity of M3C decreased and diffraction peaks of M7C3 appear for the first

time. The precipitation within the dendrites is now generalized (Fig. Sc).

a) b)

c) d)

Fig. S. Evolution of the microstructure with tempering. a) Tempered at 500°C; b) Tempered at 600°C: c) tempered at 640°C; d) tempered at 700°C.

312

As the tempering temperature increases up to 700°C, a further decrease in the amount of

M3C occurs, this phase being replaced by M7C3 (Fig 7d). The main structural modification is a

coarsening of the carbides and the fragmentation and spheroidization of the interdendritic carbide

network (Fig. 8d). The lattice parameters of austenite and ferrite decrease with the tempering

temperature. Since the lattice parameter of both phases does not depend noticeably on their

chromium content, this variation denotes the progressive decrease of their carbon content as

carbides precipitate. Also, this evolution affects the austenite, suggesting that precipitation

occurs within this phase or at the grain boundaries, before it transforms, probably to ferrite and

M7C3' The lattice parameter of a ferrite does not vary noticebly at temperatures higher than 640

°C, suggesting that the evolution of this phase is complete at that temperature. The variation of

the hardness as a function of the tempering temperature after LSM and after qucnching from

1040 °C [14] is despicted in Fig. 9. As a result of the surface melting the secondary hardness

peak is shifted to a higher tempering temperature: after conventional quenching the hardness

peak appears at about 500 °C and after LSM appears at about 600°C. Similar results has already

been reported by Peng et al. [15] and Rayment et al. [9] and is probably due to the fact that the

as-solidified austenite is more stable because of its supersaturation in alloying elements.

However, a deeper understanding of the phenomena will require a more detailed metallurgical

study.

700

650

C\I 600 ci > ~ 550 Cfl Cfl (I) 500 .:::

"0

iii 450 I

400

350 0

o Laser melted

III! Quenched from 1040 °c

100 200 300 400 500 600 700 800 Tempering temperature (oG)

Fig. 9. Variation of the hardness with the tempering temperature.

313

4. Conclusions

1. After laser surface melting AISI 440C martensitic steel presents a dendritic structure

consisting of retained austenite and small proportions of martensite and M3C.

2. The melted zone is characterized by a relatively low hardness of about 500 HV.

3. Retained austenite can transform into martensite by tempering at temperatures about

600-650 °C. This heat treatment induces secondary hardening, due to the precipitation of M7C3'

which progressively replaces M3C. The appearance of M7C3 coincides with the maximum

hardening, which occurs at 600 °C.

4. After the LSM the secondary hardening peak is shifted to a temperature about 100°C

higher than after conventional quenching.

References

1. Hertjen, D. J. and Jarvis, R. A. (1983): "Rolling Element Bearings". in Jones, M. H. and Scott, D. (Eds.), "Industrial Tribology", Elsevier, Amsterdam, 132-183.

2. Zum Gahr, K. H. (1977). The influence of thermal treatments on abrasive wear resistance of tool steels. Z. Metallkde. 68: 783-792.

3. Kalousek, J., Fegredo, D. M. and Laufer, E. E. (1985). "The wear resistance and worn metallography of pearlite, bainite and tempered martensite rail steel microstructures of high hardness". Wear 105:199-222.

4. Rice, S. L. (1987). "Pitting resistance of some high temperature carburized cases". SAE Paper N. 780773: 1-8.

5. Cheng, L. C., Wu, T. B.,and Hu, C. T. (1988). "The Role of Microstructural Features in Abrasive Wear ofa D-2 Tool Steel". J. Mater. Sci. 23: 1610-1614.

6. Kusinski, J. (1988). "Laser Melting ofT! High-Speed Tool Steel". Metal!. Trans. 19A: 377-382. 7. Ahman, L. (1984). "Microstructure and its Effect on Toughness and Wear Resistance of Laser

Surface Melted and Post Heat Treated High Speed Steel". Metal!. Trans. 15A: 1829-1835. 8. Kim, Y. W., Strutt, P. R. and Nowotny, H. (1979). "Laser Melting and Heat Treatment of M2

Tool Steel: A Microstructural Characterization". Metal!. Trans. lOA: 881-886. 9. Rayment, J. J. and Cantor, B. (1981). "The As-quenched Microstructure and Tempering Behavior

of Rapidly Solidified Tungsten Steels". Metal!. Trans. 12A: 1557-1567. 10. Vilar, R., Conde, O. and Colin, D. (1990). "Laser Surface Melting of AISI 420 Tool Steel".In

"Surface Modification Technologies III", Sudarshan, T. S. and Bhat, D. G. (Eds.) TMS, Warrendale, 343-358.

11. Vilar, R. , Cola~o, R. , Colin, D. and Conde, O. (1990). "Laser Surface Treatment of a High Chromium Martensitic Steel", In Bergmann, H.W. and Kupfer, R. (Eds.), ECLAT 90, Proc. 3rd Eur. Conf. Laser Treatment of Materials, AWT-Arbeitsgemeinschaft Warmebehandlung Werkstofftechnik, FRG, 377-388.

12. Vilar, R. and Colaco, R., "Laser Surface Melting of Martensitic Stainless Steel", LAMP'92, Proc. 2nd International Conference on Laser Advanced Materials Processing, Nagaoka. Japan, 1992. (to be published).

314

13. Hegge, H.G .• De Beurs. H .• Noordhuis. J. and De Hosson. J. Th. M.(1990). "Tempering of Steel During Laser Treatment". Metall. Trans. 21A: 987-995.

14. Heat Treating of Stainless Steels. (1981). Metals Handbook, Vol. 4, American Society for Metals. Metals Park, OH .• p. 634.

15. Peng, Q.F .• Shi. Z .• Hancock. I.M. and Bloyce. A. (1990). "Energy Beam surface Treatment of Tool Steels and Their Wear". Key Engineeri Materials. Vols 46 & 47: 229-224.

LASER BEAM LITHOGRAPHY FOR 3-D SURFACE PATTERNING

C.ARNONE, C.GIACONIA, C.PACE Department of Electrical Engineering University oj Palerrno Viale delle Scienze 90128 Palermo, Italy

S.BONURA, M.GRECO CRES - Center for Electronic Research in Sicily Via Regione Siciliana, 49 90046 Monreale, Italy

ABS1RACT. A low power laser processing unit, for microlithographic applications on non-planar surfaces, is described. By combining proper laser beam handling, micropositioning, software control and surface coating techniques, a 5-axis robotic system for laser writing has been set up. Light from a He-Cd laser source is fiber-delivered to a writing head, which moves around a resist coated solid object. After exposure, traditional wet processing can be applied. The unit is capable of patterning metal films deposited on samples up to a size of 5Ox50xlOO mm, with 5 micrometer spatial resolution. An application in 3-D circuit fabrication is presented.

1. Introduction

Lithography is commonly understood to be the planar process in which a large beam of

visible or UV light transfers a 2-D pattern from a mask to a resist coated substrate. Although this

is an already well-established technology in microelectronics, different alternate patterning

techniques have recently been tried, many of them based on focused laser writing. This technique

can be used both for so-called in situ processing and for resist exposure [1]. In the first case,

direct deposition of patterned films or localized etching is achieved, actually eliminating all

phases of the microlithographic process. In the second case, microlithography still occurs, but

instead of illuminating large areas of the substrate through a mask - a "parallel" process - a

localized exposure along the desired path - a "serial" process - is used. While laser writing on

315

S. MartelLucci et a/. (eds.), Laser Applications for Mechanical Industry, 315-320. © 1993 Kluwer Academic Publishers.

316

2-D surfaces is convenient only for specific tasks in a market dominated by conventional

lithography, it can have a major role for processing non planar surfaces, where no mask can be

used. The complete system for laser writing on resist, described in this work, represents a first

generation of 3-D laser microtools. It has been set up in order to investigate the possibilities that

such a system can offer to microelectronics, micro-optics and micromechanics.

2. System setup

The laser writing setup is sketched in Fig. 1. A He-Cd laser, (60 mW, 442 run) is the

M2

Fig. I. Hardware setup.

317

writing source; the beam goes through a mechanical shutter (S) and 6 selectable filters mounted

on a wheel (W). It is launched into the optical fiber (F) by means of a 50 mm focal length lens

(Ll). The weak beam transmitted by the dielectric 45 0 mirror (MI) is used for power monitoring.

Another beam, from a I mW He-Ne laser, is used as a pointing light, through M2, MI and Ll.

Easy and independent optimization of beam-to-fiber coupling is achieved by micrometric

regulation of M2 and of the dichroic splitter Ml. Guiding of the two laser beams to the writing

head is obtained by a standard 50/125 multimode fiber, I m long. Both fiber ends are terminated

with FC connectors. The exit end is imaged on the specimen surface by an adjustable 2X

reduction optics made of a collimating lens (L2) and a microscope objective (0) with 0.25 NA.

The writing spot diameter is 25 ~m. The surface is illuminated by yellow light from a double

bundle of optical fibers (FB). Its image is reflected by the beam-splitter B and focused on the

CCD sensor of a lightweight TV camera. A sharp-cutoff yellow filter YF stops the intense blue

laser light reflected from the surface. A 5-axis robotic system determines the position of the

writing spot in the work field: 3 axis are devoted to the movement of the sample and 2 axis to

positioning the writing head. The sample can be translated in the horizontal X-Y plane and

rotated (CP) around an axis normal to that plane. The writing head can be raised along the Z axis

and rotated (8) around an axis normal to the X-Z plane. The translation stages are equipped with

stepping motors, and allow 1 ~m resolution for X, Y and Z axes and 1/1000 or 1/1000°

respectively for 8 and cP axis. The overall positioning accuracy can be estimated as around 5 ~

m, taking into account mechanical backlash and elastic yielding. A manual stage for fine

adjustment of X-Y position and tilt (MFPl) is placed between the cP stage and the sample holder,

for calibration purposes. Likewise, a manual fine positioning X-Z stage MFP2, mounted between

the cP stage and the writing head, allows for positioning of the beam focus on the head rotation

axis. The laser beam is constantly kept orthogonal to the surface of the sanlple. In order to

properly match the programmed movements with the surface, suitable procedures have been

developed for precisely setting the absolute initial position. Safety mechanical interlocks are

provided for all possible movements of the laser head and the sample.

3. Control software

An important step towards 3-D microliiliography is the development of a computer

program for the simultaneous automatic control of the 5 axis robotic system. In order to

318

optimize the performance of the machine, computer processing of the drawing data has been split

in two parts: a pre-processor and a driver (Fig.2). The task of the pre-processor is the translation

of the drawing data from a standard compact format - such as CIF or AutoCad DXF - to a low

level format (IRC) suitable for being executed by the driver. The IRC (Incremental Robotic

Code) format has been designed for being processed in real time by fast software routines

running on a Personal Computer. This drives an I/O board which provides pulses to the stepper

motors. Pre-processing of the geometries is implemented off-line, preventing drawing time

from being affected by computing time delays. The pre-processor selects the movements of the

laser head by a tracking technique that keeps real trajectories close to the designed ones, within

an assigned error which can be as low as 5 J.Ull, as limited by the hardware setup. This concept is

sufficient to describe generic surface geometries. An important task of the pre-processor is the

optimization of the drawing time. This is accomplished through different mechanisms, one of

which is the generation (when it is strategically convenient) of acceleration and deceleration

ramps for the stepping motors. Of course this would lead to uneven exposure, unless proper laser

power control is accomplished. When needed, the low level code generated by the post-processor

is able to assign a specific exposure energy to each pixel of the 3-D frame. By interfacing the I/O

board to an acousto-optic beam modulator, single pixel programmed exposure can be easily

achieved. Where possible, the pre-processor also minimizes computing time and disk space

requirements. Although only 5 axis are used for the machine described here, the IRC code allows

up to 7 axis to be simultaneously in motion.

4. Coating technique

For two-dimensional surfaces, substrate coating with resist is usually done by spinning

with controlled speed and acceleration. The film dries off during the spinning time, and leaves a

solid film whose thickness depends on the spinning speed and the initial liquid viscosity. This

Workstation PC (fast) PC standard IRC

3-D CAD format

Pre-processing Sistem driving format

Fig.2. Flow diagram of the writing system control.

319

process leads to unavoidable lack of thickness unifonnity near the edges of non circular

substrates and, most important, cannot be used for 3-D surfaces. In order to overcome the above

difficulties, a novel coating technique based on ultrasonic spraying has been developed [2].

Circularly symmetrical substrates have been mounted on a slowly rotating holder while spraying

the resist, with good results in tenns of coverage and unifonnity. Coatings as thin as 0.5 f.UIl have

been obtained with this method. In the specific application discussed below, low resolution (10 ~

m) was needed. Therefore, in order to minimize process sensitivity to particle contamination, 2 ~

m thick resist was used.

5. Test application

An application around which the whole system has been tested is the fabrication of

conductive paths from the flat top of a cylinder down its side surface, across a 45° edge, as

shown in Fig.3. The sample is made of glass, and evenly coated with a thin gold film. A positive

process is used to define the conductors: after coating the sample with resist, the exposed

portions of the sample make the insulation gaps (30 ~m) among the gold lines (100 ~m wire).

Development and etching complete the patterning process. The diameter of the sample is 8 mm,

and its length 50 mm.

Fig.3. Detail of the processed test sample, illuminated from inside.

320

ACKNOWLEDGEMENT. The authors wish to thank G. Busacca, head of Microwave Tubes

Division of Alenia S.p.A. for encouraging this work, and F.Bracciante, P. Cusumano, G. Lullo

and F. Trapani for their helpful technical assistance. This work has been partially supported by

Italian National Research Council, under "Progetto Finalizzato Tecnologie Elettroottiche".

References

1. Arnone, C. (1992): "The laser-plotter: a versatile lithographic tool for integrated optics and microelectronics" - Microelectronic Engineering 17,483 .

2. Giaconia, c.: private communication.

LASER TECHNIQUES IN POWER PLANT COMPONENT MANUFACTURING AND REPAIRING

W. CERRI, L. GARIFO CISE Tecnologie Innovative SpA, p.o. Box 12081 20134 Milano,ltaly

D.D'ANGELO ENELICRTN Via Rubattino, 54 20134 Milano,Italy

ABSTRACT. This chapter deals with higth power laser applications concerning power plant components. Examples of laser cutting, welding and surface treatments (hardening and cladding) for manufacturing and maintenance are reported. Laser materials processing is becoming widespread in manufacturing of components for generation and transmission of electrical power. Recently some field applications have been reported that offer cost effective solutions for maintenance problems. We present examples of laser cutting, welding and surface treatment transferred on the production floor, and describe in-situ repair systems under development.

1. Cutting and welding applications

Some components used in energy installations are machined using laser cutting systems.

For example, an EPRI Report [1] mentions the use of a 1 kW C02 laser for production cutting of

tube plates used in heat exchangers, while at Kawasaki Heavy Industries a multikilowatt laser

beam is used for the blocked-out cutting of steam turbine blades [2].

As far as welding is concerned, the main advantage of the laser process is the low heat

input. For example, it has been observed that the laser welding technique does not produce any

precipitation effect at grain boundaries in austenitic stainless steels, thus eliminating corrosion

problems, as in nuclear fuel containers [3]. Due to this advantage, the laser welding technique

has been submitted to experimental and qualification tests at CISE [3] and at

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S. Martellucci et al. (eds.), Laser Applications/or Mechanical Industry, 321-324. © 1993 Kluwer Academic Publishers.

322

the UKAEA Culham laboratories. A particular by interesting field of application for laser

welding is in manufacturing of diagnostic sensors (noise emission sensor, corrosion probes, etc.).

In general, these applications call for assembly with modest static and dynamic stresses. For

example, laser welding has been applied to noise emission sensors for boiler components, where

sealing of the sensor case has to be carried out with no damage to electronic components.

A further interesting field of application, currently at the R&D stage, is the butt joining of

pipes. This type of operation is presently limited to joint thickness up to 15 mm; qualification of

laser processes on low-alloy ferritic steels , with considerable advantages in time execution and

operational simplicity compared with the TIG process, have been reported [4J.

2. Surface treatment applications

In general, laser hardening is a convenient process when it is necessary to harden

well-defined areas of components which have a complex geometry. A typical example of such a

manufacturing requirement is the hardening of the leading edge of turbine blades (in martensitic

stainless steel) to enhance their resistance to erosion due to droplets of condensed steam [5]. In

Germany a laser hardening process for these components has been transferred to production in

the manufacturing of 200 MW units [6], following favorable erosion tests results.

A second type of surface treatment with lasers of interest in energy plant component

manufacturing is in surface cladding. In this case surface melting is produced while the clad

material is introduced into the molten pool by various means. Generally these alloy elements can

be predeposited, or they can be injected directly into the laser beam's focal point by the flow of

an inert gas. Deposition techniques are in the experimental stage, but in some cases have already

been transferred to production, as in the case of gas turbines.

A particular type of laser cladding known as "reverse machining" consists of the injection

of a material having characteristics similar to the substrate, so as to locally restore the

dimensions of partially worn components. A process of this type, developed by General Electric

[7], has been transferred to production [8] on components such as bearing seals in IN7l8.

Application studies have been recently started up in reconstruction of Co-base superalloy

components, such as gas turbine nozzles [9].

A laser cladding process by means of predeposition of the alloy PW A 694 is in production

on gas turbines at Pratt & Whitney [10]. For several years Coherent General has been

manufacturing a laser cladding system for refurbishment of gas turbine parts, which has been

323

transferred into production at Wstinghouse and General Electric [11). Rolls Royce reported that,

compared with TIG techniques, the laser process provides better control of the dilution of base

material in the deposited layer, a 50% reduction in the required deposition material and a

reduction of the machining cycle by a factor of 10 in gas turbine cladding [12). CISE together

with ENEL/CRTN, Milan [13) is carrying out experiments to optimize the laser deposition

process, and also to extend it to metal carbides.

Finally, special mention should be made of a particular type of laser treatment that is

especially interesting for energy transmission: the so-called "magnetic domain refinement" of Fe­

Si magnetic laminates for transformers (14). Laser surface treatment, by reducing the size of the

magnetic domains, can achieve significant reduction (30%) in magnetic losses in the laminates of

electrical machines. Laser installations dedicated to this specific process are now in production at

ARMCO and Nippon Denso [15).

3. Power laser processing techniques for in-field repairing

Nd-Y AG lasers are compact and easy to install; furthermore the output beam can be

delivered by an optical fibre (due to its near I.R. wavelength). These characteristics make this

kind of laser particularly suitable for field applications in power plants, where difficult access or

hostile environment conditions exist, for repair welds or diagnostic inspections. A report by

EPRI (16) describes a tube repair process in heat exchangers by laser cladding using an optical

fibre system. Westinghouse (17) developed a laser welding and cladding system for repair of

boilers, now installed at Waltz Mill Plant (USA), using a 1 kW Nd-YAG laser. In Japan the most

important application has been developed by Mitsubishi Heavy Industries for tube repair in heat

exchangers by laser welding using an optical fibre system. The system perfomlance was shown

at a feasibility prototype stage. In Europe the first application of laser technology in power plants

concerned tube drawing, an optical system in prchcaters of power plants jointly developed by

ENEL/DPT and CISE [18). The process was performed using a 300 W average power Nd-Y AG

laser and has been applied in several plants. More recently, Framatome has developed a

prototype system for remote welding in nuclear plants.

References

1. EPRI Report EM 3465 (1985): "Assessment of Material Processing Lasers", prepared by Illinois Institute of Technology IITRI, EPRI, Palo Alto CA, USA.

324

2. Alsula, T. (1987): "The application of laser and electron beam to the manufacturing of gas turbines and steam turbine engine components". Proceed. of LAMP Conference, OSAKA, Japan.

3. Mor, G. and Muller, M. (1991): "Qualification of laser welding process of nuclear fuel box presented at 3rd Int. STRAHL 91, Essen, Germany.

4. Balbi, M.,Cerri, W., Mor, G.P. and Zavanella, T. (1989): "Laser welding and subcritical annealing of a 2Cr-IMo steel for pipes operating at high temperatures": Metallurgical Science and Technology, vol. 7, pp.1I3-120.

5. Bedogni, V., Vivoli, P., Mar, G.P. and Cerri, W. (1898): "Laser and BR in surface hardening of turbine blades", Proc. of LAMP Conference, OSAKA, Japan ..

6. Storch, W. (1988): "Laserhartnung von turbinschaufelkanten"; Proc. Konf. IHC Dresden, Germany.

7. Mehta, P., Cooper, E.B. and Otten, M. (1984): "Reverse machining via C02 laser"; Proc. of ICALEO 84 Conf., Boston, USA.

8. Mangaly, A.A. (1990): "Industrial application of laser cladding"; Industrial Laser Handbook. 9. D'Angelo, D.,Regis, V. and Bracchetti, M. (1990): "A C02 laser approach to remanufacturing of

turbines vanes and blades"; Conf. on High Temp. Materials for Power Engin., Liege, Belgium. 10. MacIntyre, R.M. (1985): "Laser hardfacing of turbine blade shroud interlocks"; in Laser in

Materials Processing ASM. 11. Coherent General, USA, Commercial Data Sheet. 12. Dekumbis, K. (1987): "Surface treatment of materials by lasers"; Chern. Engineering Progress,

December Issue. 13. Cerri, W., Be, C.A. and Fiorini, O. (1989): "Laser deposition of composite c1addings"; Int. Conf.

on Evolution of Advanced Materials, AlM-ASM, Milan, Italy. 14. Barisoni, M. (1989): "Recent progress and possible improvements in the magnetic quality of

laminations for electrical machines", La Metallurgia Italiana, Vol. 81, pp. 595-602 (In Italian). 15. Neiheisel, G. (1986): "Full production processing of electrical steel", SPIE, Vol. 668 p. 116-123. 16. EPRI Report RP 1585, (1987): "Laser repair of heat exchanger tubes", Battelle Columbus Div.

EPRI, Palo Alto, CA USA. 17. Elder, G. (Westinghouse), private communication. 18. GiovanneIIi, M. and Motta, P. (1987): "Improvement in reliability and accuracy of heater tube

Eddy Current Testing by integration with an appropriate destructive test"; Proc. of IV Non destructive Testing, London, UK, Pergamon Press.

LASER MEASUREMENT TECHNIQUES

VIDEO SPECKLE INTERFEROMETRY AN OPTICAL MEASURING TOOL FOR INDUSTRY

O. J. L0KBERG Department 0/ Physics The Norwegian Institute o/Technology N-7034 Trondheim. Norway

ABS1RACT. Video speckle interferometry provides full field information about extremely small displacements of the test object. Incorporation of the computer has made fringe interpretation easy and increased the measuring accuracy. Applications within industrial research will be discussed.

1. Introduction

Since hologram interferometry was introduced a quarter of a century ago, its application

within industrial research and inspection has been constantly growing (see for example Refs. [1]

and [2]). The technique shows very small displacements of the test object's surface as contour

maps (fringe patterns) where each contour spacing represents a differential movement of about

half a wavelength. The technique has been extensively used for measuring vibrations and

defonnations of a wide variety of industrial objects.

The conventional holographic process is too slow and cumbersome to be effectively

applied for high volume testing under industrial conditions. However, if we record the holograms

directly on the photosensitive surface of the video camera, and process and display the image by

electronic means, we produce 25/30 (European/American TV-standard) holograms each second.

The resulting system provides fringe patterns of lower quality than conventional holographic

interferometry, but this is amply outweighed by the speed and real time operation of the system.

The use of video systems to record holograms and speckle patterns for interferometric

purposes was proposed independently and almost simultanously by three groups [3,5] as early as

1970 - 71. The technique did not, however, grow into the industrial inspection tool as expected

and for many years research within the field was carried out only in a few academic institutions.

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S. Martellucci et al. (eds.), Laser Applications/or Mechanical Industry, 327-350. © 1993 Kluwer Academic Publishers.

328

A major change occurred around 1985. At that time full field determination of the optical

phase by digital computers and image processing had been making a gradual progress from

classical interferometry [6] to hologram interferometry [7) and finally came to video speckle

interferometry [8, 10]. With the introduction of computerized fringe analysis, interpretation and

numerical analysis of the fringe patterns became simple, which removed one of the biggest

obstacles for the general acceptance of video speckle interferometry (as well as for conventional

hologram interferometry) as a diagnostic measuring tool. In addition, the emergence of computer

analysis has coincided with the introduction of exciting new hardware like sophisticated CCD -

cameras, fiber optics and (semiconductor) lasers, which have given an extra boost to the field of

video speckle interferometry.

In this chapter, we give a general overview of the present status of the field, describing the

most common methods illustrated by a few representative examples of applications. For a more

detailed presentation, the reader should consult papers from the Proceedings of the Society of

Experimental Mechanics 25 year anniversary for Hologram Interferometry [11) or the

Proceedings from the annual SPIE meetings in San Diego, which contain numerous papers on

video speckle interferometry, see e.g Ref. [12).

2. The video speckle system

We first briefly outline the basic principles of video speckle interferometry with reference

to the flow diagram shown in Fig. 1. Later we will discuss in more detail how the various

systems are built up, see also Ref. [13, 14]. In Fig. 1, we have split the system into an optical

registration part and an electronic processing part. The optical registration part is essentially a

two wave interferometer. The light from the laser is split into two branches - the object wave and

the reference wave. The object wave consists of the light reflected from the illuminated object

surface and imaged onto the video camera. The reference wave is a uniform or a speckle wave -

in some cases it might even be a slightly shifted version of the object wave. In most systems, the

phase of the reference wave can be changed either in discrete steps or continously by the phase

shifter - PS. This controlled phase variation is necessary for most computerized methods of

fringe analysis, as described later.

All information about the object's position relative to the recording medium is contained in

the phase of the object wave. We encode this information by combining (interfering) the object

wave with the reference wave as indicated in the flow chart. The result is an image of the object

329

which is covered with a seemingly random intensity pattern (speckles). Due to the interference

mechanism, the pattern changes cyclically as the phase(s) vary in the two waves.

The video camera, today almost exclusively a CCD camera, converts the intensity

distribution of the image interference pattern (or the hologram) into a corresponding video

signal. In this way the phase information is transformed from the spatial to the temporal domain,

which enables us to proceed with conventional electronic processing of the interferograms. The

video signal is then electronically processed. In most cases, the signal is processed by band-pass

filtering followed by a full wave rectifier. This process can be shown to simulate the holographic

reconstruction process. In some systems, such as in - plane and shear systems, electronic

subtraction of frames is necessary to obtain fringes of sufficient qUality. In modem systems, the

video images are digitally processed in a computer based image processing system. Typically,

the system digitizes the video images to an array of 5l2x5l2 pixels, each pixel with 256 grey

scale levels, using a frame grabber board with 16 on-board frame buffers to store 16 video

images (frames). A co-processor board is used for 16 bit arithmetic and logical operations on the

stored images. The final processed signal is fed into a video monitor. The resulting video image

is covered with fringes, which represents the movement of the test objects during the frame

exposure or between separate frames. The computer may also control the excitation of the object

and phase modulation of the reference wave, as indicated by the dotted lines in Fig. 1. In

practice, a time consuming procedure like analyzing the vibrations of a complex systems can be

fully automated.

For practical use of a commercial system for video speckle interferometry, there is

OPTICAL REGISTRATION ELECTRONIC PROCESSING

......... ~..,....:..t"'..,-------------T

Obj.Wave (Image)

Video Camera

Analogue -Digital Proc. Frame grab

Computer

.,. ________ ' I

• • • • I • eXCItation 1_ J

• • .. _------_ ..

Fig. 1. Block diagram of a video speckle interferometer.

Monitor

330

nonnally no need to understand its physics and construction. The optical head containing the

laser, the opto-mechanical set-up and the video camera is contained in a small compact unit.

Pointing the illuminating beam at the test area and choosing an appropriate image magnification

will in most cases ensure that good recordings of the vibrations or defonnations are immediately

seen on the monitor screen. If the optical head is rigidly constructed and the object is reasonably

stable, the short, repetitive exposure time of the video system enables practical holographic

inspection outside the controlled environment of the optical laboratory .

2.1. THE FRINGE FUNCTIONS

As in classical interferometry, video speckle interferometry measures changes in the

optical pathlength, defined as the distance in wavelengths multiplied by the refractive index of

the surrounding medium. Assuming that the refractive index is constant, we can study and

measure defonnations, dynamic or static, of the test object. Changes in the refractive index only,

on the other hand, lead us into the study of temperature and pressure distribution in gases, and

this is of extreme importance. Changes in refractive index and/or changes in wavelength may, as

in hologram interferometry also be used to contour surfaces. For a further discussion of

contouring by video speckle interferometry the reader should consult Ref. [13]. As in hologram

interferometry, video speckle interferometry records the metric infonnation either by averaging

the displacement during the exposure or by combining the position at two (or more) exposures.

We have therefore basically two different modes of operation, depending on whether data

acquisition is taking place during a single frame, or two or more frames are combined.

2.1.1. Single frame exposure. This mode can be further subdivided depending whether the

exposure is continous (time-average) or broken up into subexposures. The time-average methods

are, like their holographic counterparts, mainly used for measurements of objects vibrating at a

single frequency, Assuming the vibration amplitude and phase to be respectively given by

ao(x,y) and <J'o(x,y), we obtain the well known fringe function for the reconstructed image

intensity:

2 [2It ] ITA (x,y) "" IO Tg($)ao(x,y) (1)

where: IO is the Bessel function of zero order and first kind; and, g(<\» is common interferometric

331

sensitivity factor containing the illumination-observation - displacement directions. Note that

Eq. 1 contains only information about the amplitude distribution, whereas vibration phase

information is lost during the averaging process. The time-average vibration method can be

further improved by introducing "phasemodu1ation" at equivalent amplitude and phase ar<x,y)

and <Ilr<x,y). This leads to a slightly modified reconstructed intensity distribution:

(2)

We see that the fringe function is unchanged from Eq. 1, but its argument is now given by the

vector difference between the object and reference vibrations. By either observing the zero order

fringe or working on the linear part of the zero order, we can determine phase distribution and

extend the measurement range, as described in earlier publications [15-17]. Stroboscopic

operation of the interferometer can also be used to study harmonically vibrating objects [18]. In

this case short exposures effectively freeze the displacement in two positions to give a resulting

intensity distribution:

(3)

where: ~(x,y) is the displacement between the exposures; if the exposures are placed at the

extrema of the vibration cycle ad(x,y) represents the peak to peak displacement. Fig. 2 shows the

same vibration recorded by time-average and stroboscopic speckle techniques. Note how the

nodal lines are easily detected on the time average recording because the zero order fringe

a) b)

Fig. 2. a) Time-average and b) stroboscopic recording of a vibrating steel plate.

332

00 2(0) = 0), which is far brighter than the higher order fringes. In the stroboscopic recording all

fringes have equal brightness and fringe order has to be detennined by observing several fringe

patterns at varying excitation levels. If the movement is hannonic but with a strong random

component, for example caused by running machinery, we have to sample the motion with only

two pulses a frame, giving the same fringe function as in equation [3]. As this kind of object

nonnally moves at hight velocities, the only practical alternative to arrest the motion is the use of

a pulsed laser, preferably a repetively pulsed laser which can be run at video-frame rates [19].

2.1.2. Frame combination. In most measuring cases we combine the infonnation from two or

more frames, as for example in the analysis of slow defonnations. We record a reference frame

which is subtracted from frames coming from the videocamera. The subtraction process is

preferred for noise reduction, and essential for some techniques such as shearing and in-plane

video speckle. We get essentially the same fringe function as for addition processes like the

stroboscopic one in Eq. [3], except that the zero order fringe is dark. The dark field background

can be used with advantage in the visual detection of small displacements. The resulting intensity

distribution for a fringe pattern generated by "subtraction" is given by

(4)

For numerical analysis of all the fringe functions mentioned here, several frames are combined

using various algorithms to obtain the displacements, as will be discussed later.

object __ -----illu:;:-:7'?'V

Imag.lens

~

* Fig. 3. The uniform reference system.

Rec. med,

333

2. 2. THE BASIC SYSTEMS

We briefly describe the optical arrangements and the resulting sensitivities of the three

principal systems used for video speckle intenerometry.

2.2.1. Separate (uniform) reference system. The unifonn reference wave system has so far been

the most popular system (in the literature it is called several names such as ESPI, DSPI, TV­

holography, electronic holography etc.). Its basicl construction is described in all the pioneering

papers. We show in Fig. 3 some of the elements of a unifonn reference system, where for

simplicity we have omitted the laser and some of the standard components used in every two­

branch intenerometer. The illuminated test object is imaged through the beam combiner Be onto

the video target. The reference wave is superimposed on the image wave by the beam combiner,

and adjusted to be centered in the exit aperture of the lens. The choice of a good beam combiner

solution is of utmost importance for good optical quality in the system, and the reader should

consult Ref. [12] for a discussion of this problem. A related problem is to reduce the effect of

multiple reflections of the reference wave in the target, which can be solved as indicated in Rg. 3

by cementing a glass wedge to the target. The controlled phase change of the reference wave (the

box labelled PS-PM) may be implemented by displacement of mirrors, elongation of optical

fibers or double refraction in electro-optical modulators. As phase stepping routines have become

integrated into all modern video speckle intenerometers, the phase box will be omitted when

discussing the remaining systems. The directions of the illumination and observation are

important for the sensitivity of the system, as indicated in Fig.4. The sensitivity vector kg is

given by the difference between the illumination vector ki and the observation vector kg. Virtual

...... object surface . Ir ..• ~ ...... ~.~ ...... ~ .. ~7 -~~~~

------- ----------------z -:: :;l' _:

illumination r:/ l' ... ~~

'\. senl.Slt1VltY ~/ 1<0 , ~ , ~ . ~ • • , . .. ~~~

... 1<5 .. Fig. 4. Intenerometric sensitivity.

334

surfaces of equal optical path lengths from the set-up are indicated as dotted lines on the figure.

The distance between neighbouring surfaces corresponds to an optical patblength period of one

laser wavelength. The fringe order resulting from a certain displacement is given by the number

of surfaces crossed by the displacement vector d or found by projecting the displacement vector

on the sensitivity vector. In most systems, the illumination and observation direction almost

coincide, which results in maximum sensitivity to out-of-plane displacements of the object.

Before describing systems where the reference wave is more complex, we should point out that

the uniform reference wave can be replaced by a speckle reference wave [20]. However, the

quality of the fringe patterns is reduced due to less favourable speckle statistics.

2.2.2. Double illumination system [21]. This system has a clear parallel to Laser Doppler

Velocimetry (LDV) both in terms of construction and interferometric sensitivity. Fig. 5 shows

the usual design of such a system. The system has no separate reference wave, but the object is

illuminated by two waves and we might use the object/reference name for either one of the

reflected waves. To simplify the fringe interpretation, the illumination waves should preferably

be plane and symmetrical to the normal of the (plane) test surface. From LDV we know that the

sensitivity vector will be normal to the interference planes as indicated in Fig. 5b, which gives

pure in-plane sensitivity to the displacement. As the in-plane component of the displacement is

of importance to engineers computing stress and strain levels, this system is technically very

important. Note that the assumption about pure in-plane sensitivity is only exactly true for a

plane object. In the case of non-planar objects, we only measure the x (or y) component of the

displacement and information about the object topography is necessary to calculate the in-plane

component. To determine the x and y components, we have to rotate the object or use a "double"

doubleillumination [22]. The fringe quality and measuring range are lower than in a uniform

reference system due to speckle statistics and speckle decorrelation.

sensitivity

..... 1<·2

1 -.

illurn.2

a) b)

Fig. 5. a) in-plane system; b) The corresponding interferometric sensitivity.

object Imag. lens c::::::::::::>

leHHHe

Rec.Med.

a)

TV -holography ••••••••• · ... ... . · ....... . · .. . .. . · .. . ... . · . . ... . · ....... . · ... .. . · ....... . · ...... . · . . . .. ••••••••• ................... ................... . ................. . ' .. :::::: .• ' . '::::::::.

':':':':-:':-:':"'::':':':':':':':', ................... · . . . . . . . . .

Shear

b)

335

Fig. 6. a) A general shear system; b) Wavefront shearing (top) and detection of a defect in the presence of tilt in the case of TV -holography (uniform reference beam) and shear.

2.2.3. Speckle shear system [23-25]. There are numerous ways to construct a speckle shearing

system, in fact there are probably as many variations of the shear system as there are workers in

the field. Some of the systems are very simple, which is one of the attractions of shear

interferometry. In Fig. 6 we use a standard Michelson interferometer to provide shearing of the

wavefront. By tilting one of the mirrors a small angle o~, we move the wavefronts relative to one

another. This gives us linear shear in any desired direction given by the tilt angle of the mirror.

Similarly, if we move the mirror a distance ox, we obtain interference between images of

different magnifications or radial shear. Rotational shear can be obtained by replacing one of the

mirrors with a right angle prism in reflection, which can be rotated to give the desired shear. In

practice, linear shear is most often used, as it provides constant shear across the image and fringe

patterns which are fairly easy to interpret. When the shear is small compared with the

displacement, the resulting fringe pattern can be approximated by the derivative of the

displacement in the direction of the shear as indicated in Fig. 6 b ( for drawing clarity, the shear

~x is exaggerated). As a result, uniform moments are cancelled, giving only a constant fringe

order which the resulting fringe pattern "rides" on. The shear interferogram displays only abrupt

changes in the displacement, which is of key interest in non-destructive testing. The lower right

side of Fig. 6 b shows the advantage of using shear for defect detection in the presence of tilt.

Due to either increased pressure or heat, the defect gives rise to a non-homogenous displacement

of thc surfacc. In the ordinary uniform reference set-up (TV-holography), we observe an

interference pattern between the original position (dotted line) and the deformed state. The tilt

336

produces a straight line pattern where the contribution from the defect results only in small

undulations on the regular pattern, which are hard to detect. In shear, we observe interference

between the shifted versions of the deformed and the tilted state. The effect of the tilt is removed

and the abnormal deformation is easily detected. Note also that the insensitivity to large uniform

displacement allows for higher levels of loading whereby the defects can be more easily detected.

Since this method used a common path interferometer the effect of turbulence is also negligible

for small shear. It is, however, not strictly correct to claim that the interferometer is completely

insensitive to uniform movements and turbulence, as speckle decorrelation might reduce the

fringe contrast to an unacceptable level. We want to keep the shear as small as possible to ensure

stability and a pure derivative fringe pattern which is reasonable easy to interpret. However, as

the shear decreases we enter the regime of speckle photography where the contrast is very poor.

In addition, the interferometric sensitivity is reduced to a level where defect detection is difficult

even with image processing. In practice, most shear interferometers are operated with a shear

which is comparable to or larger than the smallest deformation details we wish to detect. Then

the interferometer acts more like a speckle reference interferometer with a reference that is

strongly spatially phase-modulated. However, even at this larger shear the interferometer is still

superior to the uniform reference interferometer in terms of stability, due to the common path

and due to compensation for tilts in the shear direction.

3. 1. COMPUTER INTEGRATION

The incorporation of the computer into video speckle interferometry has permitted not

only fringe analysis, but also made possible the manipulation of the interferometric speckle

pattern, giving visual display improvement and numeric information.

A single frame video interferogram recording is noisy, and averaging several frames is

very beneficial, both visually and for data acquisition. The averaging is most efficient and easily

performed in vibration analysis, where the same pattern can recorded and displayed over an

extended period of time. Speckle averaging is very useful, where in the speckle carrier is changed

(decorrelated) from frame to frame to further improve the final, averaged image [20,26]. Speckle

averaging is easily performed automatically and the result displayed in real time using the

computer and a modem frame store. The recordings shown in Fig. 2 have both been speckle

averaged.

The visual improvement by frame manipulation may be quite astonishing. Using a fringe

improvement technique specially developed for unstable objects [27], up to 100 time average

fringe orders have been detected across the image, which is quite remarkable considering the low

337

intensity of the J02 - fringes at such high fringe orders. See also reference [28] for the description

of a highly computerized system providing fringes of nearly holographic quality in real time.

3. 2. FRINGE ANALYSIS

As we have stressed several times, the surge of interest in video speckle interferometry is

mainly based on the use of computers in the system. In that context, most important is the

computer s ability to analyse the fringe pattern and assign numerical values of displacement to

each pixel, see e.g. Ref.[29].

It is useful to divide numerical analysis of fringe patterns in video speckle interferometry

into three major groups: static displacement, repetitive displacements and dynamic, single event

displacements. Each group demands a different type of fringe analysis approach. We will discuss

the specific problems for these groups and present solutions, especially stressing solutions which

are not commonly used in classical and hologram interferometry.

In the following discussion we assume the direction of the displacement to be known and

parallel to the sensitivity vector. That is, in Eq. 1, 2, 3, and 4 we set g (cp) = 2. In practice, the

direction of the displacement vector is often unknown and varying across the surface of the test

object. In such cases we only know that we are recording a fringe pattern where the optical path

length change is the projection of the displacement onto the sensitivity vector.

In video speckle interferometry we have at our disposal the in-plane technique, which

provides directions of the sensitivity vector not available in hologram interferometry. If we

combine the information from a uniform reference system with two orthogonal recordings from

an in-plane system, we obtain directly x, y and z displacements from which the total

displacement vector can be found [30].

Another approach is to change the direction of illumination in a uniform reference system

[31] to map the displacement field, where the essential surface contours can also be mapped in

the same set up. Which of these techniques is the best has not been investigated. Intuitively, the

separate determination of the components seems most promising, but it may be that the more

favourable signal to noise ratio of the uniform reference system makes it competitive.

3.2.1. Static displacement. This class consists of measurements where we have a no-load load

situation or the displacement is slow or controlled, as in an experiment using heat or pressure

increments. We want to measure the displacement between separate frames and suppose the

displacement during the recording of one frame to be negligible or zero. The primary result after

such an experiment is an intensity distribution I (x,y)

338

I(x, y) "" 10(x,y) + Ir(x,y)+ 2J10(x,y)lr(x,y) cos{a(x,y)-8(x,y)} (5)

where: 10(x,y) is the intensity of the (speckle) reference wave; Ir<x,y) is the intensity of the object

speckle wave; a(x,y) is the speckle phase tenn, indicating the relative phase between object and

phase wave; and, 8(x,y) is the phase change due to the displacement between the frames. The

fringe analysis problem is to isolate and quantify 8(x,y). The most common solution is to use

phase stepping, where the phase in one of the interferometric branches is stepped in equal and

controlled steps [29]. The corresponding N + I intensity distributions or frames are recorded in a

frame store. By combining the resulting N + I equations we eliminate the trivial unknowns

10(x,y), Ir<x,y) and a(x,y) in the interference equation to solve for the phase-function 8 (x,y). The

great advantage of these phase step techniques is that the effects of uneven intensities in the

illumination and reference waves etc. are to a large extent eliminated. Note that while in nonnal

interferometry the phase function represents the wavefront itself, we are concerned here with

changes in the wavefront due to the displacement. Therefore, we have to use the classical

algorithms to compute the wavefront before and after defonnation. The phase change is then

given by simple subtraction of the computed wavefronts. Most commonly used for wavefront

computation is N = 3, where we acquire from four frames IN (x,y, N-1t/2) . Combining the four

frames gives the solution for the phase angle of the wavefronts

.!l.=!L 8( x, y) = arctan 10-12

(6)

The calculated wavefront before and after displacement is then subtracted to give the

deformation wavefront. There are numerous variations of these discrete phase stepping

techniques which are used for application to video speckle techniques. Some algorithms combine

the frames in such a way that the difference phase is directly computed. For practical use in

industry, we should chose an algorithm which is tolerant of external and instrumental

instabilities. Note that all these algorithms are only useful for rather stable conditions, since we

assume that at least one, usually two frames are recorded in the reference position. If we are

working under realistic industrial conditions, for example looking at running machinery,

comparing the rest position with the vibrating position might be impossible or lead to an

excessive number of fringes. Therefore, even if a pulsed laser can be used to stop the most

violent movements, it does not pennit computerized analysis of the interferograms unless the

infonnation is recorded within a very short timespan. In practice, we have to use a double pulsed

339

laser, where the two pulses are recorded within a single frame. Possible solutions to the analysis

of single frame recordings will be discussed in Section 3.2.3.

As the basic phase-stepping routines are well known and discussed in the optical literature

[29], we will describe the max-min scanning routine which has been developed for defect

detection under unstable conditions [32,33]. The routine is based on the constant production of

interferograms in the video system and is difficult to imagine being applicable to hologram

interferometry. A defect in the test object is supposed to give a non-homogenous movement of

the surface. The image processing procedure is aimed at accentuating these (small)

inhomogeneities which are hidden in the fringe pattern. In addition, the procedure is optimized

for use on test objects which are slowly moving due to e.g thennal loading. The algorithm is

outlined with reference to the flowchart in Fig. 7. The first step is to detennine the phase <I> of

the wavefront reflected from the test object before defonnation (the object phase as defined

relative to the phase of the reference wavefront). This is done by a maximum-minimum scanning

wherein the object phase changes through values from 0 to 21t by drift or as controlled by a

phase-shifting device. During the cycle the extreme values of the phase are recorded. The

absolute sign of the object phase is detennined by adding a known phase-shift a and analyzing

the resulting intensity shift. The object is thereafter defonned, for example by heating, and the

new "defonned" phase II' is detennined by the same procedure as above. The phase values <I> and

II' representing the undefonned and defonned states are thereafter subtracted. The subtracted

phase-distribution is sent through a low-pass filter. However, a straightforward low-pass filtering

introduces excessive noise in the regions where there is a 21t phase shift. This is avoided by

shifting the image by 1t and low pass filtering (convolving) both images. The two images are

subsequently combined into one optimum image by the Lopt operator. This operator combines

the best parts of the two images, a process controlled by the pixel values in the original phase

~-----------, ~ I

I PRE I

DEFORMATION +-m-ax---m~in---I <I> ~ .... __ - ...

scannin

POST DEFORMATION <p

max-min scanning I

I

Isi'fu_d~f. }-----------:

Fig. 7. Flowchart of max - min scanning for defonnation phase measurement.

340

a) b)

Fig. 8. Subsurface structure of a honeycomb antenna revealed by slight heating of the sample. a) Raw fringe pattern: b) Gradient picture

a) b)

Fig. 9. Subsurface damage of laminate due to drop test. a) Fringe pattern; b) Gradient picture

Fig. 10. The J02 - fringe function at low amplitudes.

.~ c £ .s

slope k

amplitude

341

distribution. Finally, the gradient of the resulting image is calculated and the results are

displayed on the video-screen as a two-dimensional contour map, a false colour plot, or a mesh

plot. In Fig. 8, the technique has been used to enhance subsurface structure in a honeycomb

antenna for space applications. Note that, while the individual honeycomb cells are difficult to

detect in the "raw" interferogram, they are clearly seen in the gradient picture. Fig. 9 shows the

detection of subsurface damage of a laminate caused by a standard drop test. Visual inspection of

the surface showed no sign of damage where the steel bill had hit. Slightly heating the sample

produced the "raw" interferogram shown in Fig. 9 a, while the calculated gradient is shown in

Fig. 9 b. The damage of the fiber structure can be clearly observed in the gradient picture.

3.2.2. Vibration analysis. In deformation analysis we are concerned with the absolute value and

direction of the displacement across the deformed surface. In vibration analysis, we find the

projected displacement vector, or its amplitude, readily by time-average recordings. However, we

need to know during what part of the vibration cycle the maximum displacement is reached, that

is we need to know the distribution of vibration phase across the surface. In so-called classical

modes, the phase within each anti-node is constant, while its neighbour mode will vibrate in­

phase or out-of-phase. As the phase values thus shift between 0° and 180°, it is fairly

straightforward to analyze the vibration from the displacement map alone as obtained for

example from double pulsed interferograms. In real life, however, the classical mode is most

often replaced by a superposition of modes which makes the phase vary continuosuly across the

surface. A typical example is vibration in loudspeakers, where uniform movements can be

observed only at the lowest frequencies. Therefore, to obtain adeguate information about a

general vibration, we need to measure both amplitude and phase distributions. Here we will

describe a technique [34] based on time - average recordings using sinusoidal phase modulation

[17] (see Eq. 2). The numerical amplitude and phase values of the test object are found by

subtracting the unknown vibration vector from a reference vibration vector which is given

calibrated changes in amplitude and phase. The technique is interesting because of its ability to

measure vibration amplitudes and phase values at very low excitations. Using low excitation

levels ensures that the object vibrates within the linear region, which makes comparisons with

finite element calculations more reliable. In addition, there is a definite need to measure low

vibration amplitudes as we go to high frequency analysis, where the displacements are in the

sub-micrometer region. There are a few cases where the method cannot be used because the

excitation has to be at a certain level to induce the vibration, one example being pressure induced

vibrations in pipe structures. The fringe function for a time average recording of an object

vibration at low amplitude levels is shown in Fig. 10. The middle part of the function is linear

342

around the working point ar and we adjust the excitation so the amplitude of the object is within

this linear part. In the measuring procedure, we start with the object at rest. The reference mirror

is vibrated at amplitudes corresponding to the extremal parts ar- and ar+ of the linear part of the

fringe function. The corresponding Imax and Imin are recorded as the two first frames. These

frames are used to calibrate the sensitivity of the system in terms of an intensity vs. amplitude

factor k. Alternatively, we may record the intensity values 10 and Idark, where 10 corresponds to

no modulation and Idark corresponds the first minimum of the fringe function. A simple

calibration factor gives the correct k - value. This procedure has the advantage of giving a more

accurate k-value, due to a larger dynamic range in the initial measurements. The reference mirror

is vibrated with an amplitude ar while the object is vibrated at an excitation level which keeps its

maximum amplitude within the linear part of the fringe function: that is, about 30 nm. We step

the acoustical phase of the reference wave by 00 , 900 ,1800 and 2700 and record the corresponding

intensities 10, 190, 1180 and 1270 . These 4 recordings can be combined to solve for the

unknowns of interest, namely the phase 1I'0(x,y) and the amplitude aO (x,y)

[nmJ

57.0

J2.9

26.S

14 . 7

[dO OJ + 617 t '65 -

JI2

160,

Fig. 11. Amplitude and phase distribution for a loudspeaker vibrating at f = 1750 Hz

(und

2 J 6

lO e

1] ::f

l lJ

JlO . " Fig. 12. Amplitude and phase distribution for a crystal vibrating at f = 768,000,000 Hz.

343

() ~I90(X'Y)-1270(X'Y)] <PO x,y = arc 1180(x,y)- IO(x,y)

(7)

[II80(X, y)- Io(X,y)f +[ I90(X,y)- I270(X,y)f ao(x,y)= 2k(x,y) (8)

The mesh plots in Fig. II show the amplitude and phase distribution calculated by the

algorithms given in Eq. [7] and [8]. The object in this case is a "woofer" type loudspeaker

vibrating at 1750 Hz. The speaker has started to break up into a ring pattern, the first being seen

at the edge. The central part is still vibrating reasonably uniformly, although we have variation

both in phase and amplitude values as several vibration modes are excited at the same time. In

Fig. 12 we show the same analysis being used on a small (diameter 2mm) transducer vibrating at

7,680,000 Hz. The vibrations have two central peaks surrounded by a ring pattern. The phase

distribution shows a discontinuity, indicating a nodal point between the two central peaks. Nodal

points are observed whenever two or more modes are excited at the same time resulting in

varying phase values or travelling waves across the surface.

There are other ways to analyze time average vibrations which provide amplitude and

phase distributions. A somewhat similar algorithm is based on phase modulation [35] at higher

excitation levels. As it relies on measurements of 102 fringes of higher orders with

correspondingly lower intensities, the signal to noise ratio needs to be high. Vibration analysis

may also be carried out using stroboscopic techniques, where we measure the displacement in

several parts of the vibration cycle using normal phase shifting techniques. From these

measurements, both the amplitude and phase of the vibration can be calculated. The stroboscopic

technique should also be of interest when we are dealing with non-sinusoidal vibrations, since

sufficient sampling will give information about the higher harmonics as well. In general,

whenever we suspect that non-sinusoidal vibrations are present, stroboscopic or double pulsed

techniques should be used for analysis and measurement of the vibration instead of time-average

methods.

The most challenging problem in industrial vibration testing by spatial interferometry is

analysis of vibrations induced by white noise excitation. Quite extensive work in the author s

group has still failed to come up with a good approach to this problem, although time-average

recordings provide a partial solution by providing information about the areas where excessive

vibrations exist. A technique reported in hologram interferometry - so called holographic

vibration spectroscopy [36) has turned out to be of extremely marginal value. Numerous

344

experiments with digital analysis of time-averaged techniques - with and without phase­

modulation - in addition to pulsed laser recordings have given interesting information about the

time history of white and pink noise vibration. However, we are still looking for a good solution

which provides information about the statistical amplitude and phase. ( The meaning of vibration

phase is somewhat vague in this context - the probleme is more to determine the areas

contributing on the average in-phase to the sound pressure.) A different approach for analyzing

vibration due to general excitation, for example due to a motor running, is to measure the

resonant frequencies and their relative sound levels by acoustic means under real, working

conditions [37]. Keeping the set-up and surroundings unchanged, the resonance vibration

patterns are thereafter excited separately and analyzed by the digital methods just described. In

this way, we have enough information to synthesize the vibration due to random excitation. This

entire procedure is very time consuming.

3.2.3. Single dynamic events. In the problems described so far, we have assumed that either we

control the displacement, or the displacements are stable and periodic. As already mentioned in

section 3.2.1., under realistic industrial conditions the information has to be gathered within the

time of a single video frame. Consequently, the phase shift techniques described so far break

down, as they all rely on gathering information from several frames. In addition, pulsed lasers

with associated stability problems have to be used in these experiments, As of today, we must

combine two consecutive video fields by using a CCD camera [38], and in most cases we have to

analyze single frames, which obviously will reduce the measuring sensitivity and range. For a

general discussion of analysis of single frame interferograms, see e.g. Ref. [39]. There are several

ways of attacking the single frame problem. The simplest solution is to add reference tilt fringes

to the deformation to be analyzed, a trick well known from visual fringe analysis in classical

interferometry. This increases the read-out accuracy and gives the sign of the deformation. The

resulting interferograms can be automatically analyzed, for example, by Fourier transform

methods. Due to the speckle noise in the straight line fringe carrier, we must expect reduced

accuracy as compared with classical interferometry. Using [10-15] reference fringes across the

field we should be able to measure with an accuracy on the order of A/25. A promising technique

is based on using a uniform reference set-up where the reference wave impinges on the target at a

small angle. This angle is chosen to give a 1200 phase shift between consecutive pixels. The

three neighbouring pixels are assumed to be within the same speckle, which can be ensured by

reducing the aperture of the imaging system or by using a camera of high resolution. By treating

the identical phase pixels as individual frames in the readout we obtain three phase shifted

interferograms due to a deformation and can perform a suitable algorithm and unwrap the phase.

345

The technique may work with double pulsed lasers if we are using a camera where single fields

can be read out separately. The drawback is a loss of resolution and light, which can be partly

compensated for by using a high resolution CCD camera.

3. 3. DIGITAL FRINGE ANALYSIS IN VIDEO SPECKLE INTERFEROMETRY

When we use digital techniques in video speckle interferometry to analyze the

displacement we face several problems and sources of errors. Most of these, such as drift,

unwanted vibrations, uneven phase settings, electronic noise, nonlinear recordings and

quantization errors are common to classical and holographic interferometry [29] and will not be

commented on here. However, we should be aware that the production of nice looking three­

dimensional displacement maps is no guarantee of the accuracy of the measurement. Most

algorithms will give seemingly good results even when used under quite unstable conditions, but

the data acquisition should be very fast and controlled to avoid errors.

The speckle pattern has areas where the intensity is low or even zero. These areas do not

contribute to the interferometric signal, as their modulation is too low for reliable detection. In

addition, we have dust particles on the target, diffraction and interference patterns on the

reference wave etc. All this adds up to a substantial number of "bad" pixels in a single video

frame. These pixels leave holes in the information mapping and give rise to noise spikes in the

final results. To correct for these problems, it is often necessary to combine several processing

techniques. We have to use algorithms which search for the pixels having too low modulation for

reliable calculations and replace their values with the mean of the neighboring pixels. If possible,

several frames should be added together. In analysis of repetitive displacements, the contrast of

the speckle pattern can be reduced by speckle averaging as already described. The speckle

averaging technique has proved very efficient in reducing the number of rejected pixels. In

addition, the computed values can be presmoothed by reducing the initial 512 x 512 pixels to e.g

a 100 x 100 matrix, while a spatial N x N filter smooths noise in the final image. However,

excessive spatial filtering reduces the spatial resolution and thus filtering should be used with

care.

Another fundamental problem is how video spectcle interferometry differs from the low

resolution of the video camera. We would like to point out the difference between video speckle

interferometry and video analysis of ordinary hologram interferometry, where the video camera

is used to analyze interferograms recorded on, and reconstructed from, a high resolution medium.

In the latter case the video sensor is analyzing secondary fringes with low speckle noise, while in

video speckle interferometry the video sensor records the primary interferogram from which the

346

secondary fringes are produced. Due to the low resolution, we have to work with small apertures,

which results in speckle decorrelation as tilt of the object s surface brings new speckles into the

aperture. Opening the aperture does not help, as the resolution bandwidth acts an effective

aperture. The speckle decorrelation reduces the contrast of the fringes and thus the number of

fringes across the object which can be allowed. Speckle decorrelation is a greater problem in

techniques based on speckle reference waves as the speckle statistics are more unfavourable.

Therefore, a uniform reference set-up is preferable to avoid decorrelation.

Most experimenters avoid the speckle decorrelation problem by analyzing keeping the

number of fringes small, which makes subfringe detection particular interesting. Whenever we

are faced with a large deformation which can be controlled, the best solution is to subdivide it

into many deformation steps which are integrated into a total deformation.

4. Applications

There have been a steady growth in industrial applications of video speckle interferometry.

Here we will mention only a few examples, mainly from the research of our group. For a broader

assortment of applications, the reader should consult for example Refs.[Il] and [12]. It is

obvious that video speckle interferometry is gradually taking over much of the routine testing

formerly done by hologram interferometry. The speed, ease of handling and ability to work

outside the laboratory are most beneficial for industrial research, and we should expect this

market to grow strongly during the coming year.

Most widely accepted within the industry is shear interferometry [25], which is recognized

as an official testing method for defects and delamination in the airplane industry. It is obvious

that the shear technique will be the inspection method to be used whenever we only want to

pinpoint abnormal displacements indicating failure, weakness or other sources of trouble, and we

are not too concerned about the exact analysis of the movement. As the shear system can be very

simple and inexpensive in its construction, it is tempting to suggest that such systems can be

used in great numbers for continous surveillance of critical structures, for example, in atomic

power plants. The interferograms are available as video signals and can be relayed by cable to a

central control station.

The standard uniform reference system has also proven to be capable of applications inside

and outside of the laboratory. The biggest application area has been and still is vibration testing,

where fringe patterns of high quality can be produced in real time under very unstable

347

conditions. The lower resolution compared to hologram interferometry is not detrimental, as we

can vary the magnification to resolve any details needed. If the test surfaces can be covered by

retro-reflective coatings, very large structures can be tested even with a low power He-Ne laser

[40], as shown in Fig. 13. The Rover group has used a system based on fiber optical connection

to a powerful CW laser for analysis of vibrations and deformations in automobile bodies and

motor parts [41]. The system has worked with great success on the factory floor with no stability

precautions.

It has also been shown that digital analysis can be used to analyze vibrations under

extremely hostile conditions even, when the system uses continous wave lasers. A typical

example is the analysis of vibrating objects heated to +1000 Co, where recordings at high

temperature were less noisy than at room temperature [42] . This somewhat unexpected result is

probably due to more effective speckle averaging taking place at high temperatures. We would

also like to mention measurements of very high frequency vibrations, where the amplitude and

phase distribution have been mapped at frequencies in the 10 MHz range [43].

The same digital procedure has also been used to measure vibration of passive sonar

transducers in water tanks [44] .

Measurements under very unstable, rough conditions can be done by using a pulsed laser

system, a development pursued especially at Loughborough University [19] . This work has

concentrated on developing a pulsed system for in-plane measurements in both directions, and

interesting results have been achieved by using the polarizing properties of metal surfaces [45].

Pulsed lasers have also been used for in-plane studies of rapid events such as rotating

structures (46) .

Fig. 13. Vibrations of a car body recorded by a 5 mw He-Ne laser.

348

Fig. 14. Crack in concrete.

In pure defonnation testing using low power lasers the stability of video speckle

interferometry are not comparable to vibration testing. In this case, we are comparing different

frames which might be several seconds apart, while information about the vibration is acquired

within single frames . Nevertheless" the system has been used in practice in this mode, also due

to its real time presentation and ease of handling. A good example is the study of crack

development in a church wall due to heating and cooling during the day. The opening and

closing of the cracks have been followed and measured during periods of weeks [47] . The

speckle instruments were in this case firmly attached to the wall, whereby gross movement of the

building was effectively cancelled out. The behavior of micro-cracks in concrete and larger

cracks in sandstone has also been studied by other groups [43,49]; for an example of crack

detection in concrete, see Fig. 14. In this case, the crack appears as a discontinuity in the straight

line fringe pattern. The concrete sample was in this case submerged in water. Interesting work in

the same direction is being pursued in Japan, where speckle interferometry monitors irregular

displacements in the wall of a mine [50]. The information will hopefully be able to detect

earthquakes at an early stage. Similar work has been reported in progress in California [51].

ACKNOWLEDGEMENT. The author wants to thank S. Ellingsrud, J. T. Malmo and E.

Vikhagen for their kind pennission to use some of their results and documentary pictures.

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in situ monitoring on buildings", Opt. Eng. 29.816 - 820. 48. Lokberg, O.J., (1989): "Electronic Speckle Pattern Interferometry and Its Applications in Rock

Mechanics", Laser Holography in Geophysics (ed. S. Takemoto), pp. 168 -196. Euis Horwood Ltd-John Wiley & Sons.

49. Aasved-Hansen, E. (1990): " Real time TV - holography observation of the fracture process zone". "Micromechanics of Failure of Quasi-Brittle Materials" S. P. Shah, S. E. Swartz and M. L. Wang, Eds.(Elsevier, London) pp. 504 -513

50. Takemoto, S. and Tsubo, T., (1989): " Application of holography and electronic speckle pattern interferometry techniques to earthquake prediction" SPIE, 952, 230 - 234.

51. Rosenthal, D.M., Trolinger, J.D. and Weber, D.C., (1991): " The use of double pulsed ESPI for earthquake mitigation oflarge structures" SPIE, 1553.175 -184.

ADAPTIVE PROFILOMETRY FOR INDUSTRIAL APPLICATIONS

G. SANSONI, F. DOCCHIO, U. MINONI, L. BIANCARDI Dipartimento di Elettronica per I'Automazione Facolta di Ingegneria Universitii deg/i Studi di Brescia Via Branze. 25133 Brescia. Italy

ABSTRACT. In this chapter a review of optical profi!ometry techniques is presented. Particular emphasis is given to industrial applications of these methods, especially their system adaptiveness. The main features of an optical adaptive profilometer, which makes use of a Liquid Crystal projector, are presented. In particular, the characterization of the Liquid Crystal Display unit is discussed.

1. Introduction

F1exible manufacturing requires a number of measuring techniques, which are very useful

as far as quality control, on-line inspection, computer vision, robotics and CAD/CAM

applications.

Among the available measuring techniques, those based on optical methods are

particularly appealing, due to: i) the non contact nature of the measurement; and, ii) the trend of

technology, which favors sensors and systems characterized by decreasing costs, increasing

flexibility, good performance, and relatively small dimensions [1].

Among the available optical measurement methods, optical whole-field profilometry has a

large spectrum of possible applications. For example, computer vision needs optical profilometry

techniques to obtain the 3-D surface measurement of the workpiece. The observation of transient

deformations in thermal processing of the workpiece can be done by a whole-field profile

measurement of the object surface. The possibility of measuring the surface profile of a prototype

in order to transfer its shape to a CAD system represents another interesting application [2, 3].

351

S. Martellucci et al. (eds.). Laser Applications for Mechanical Industry. 351-364. © 1993 Kluwer Academic Publishers.

352

Measurement requirements generally depend on the particular industry, as shown in Table

1, but also vary with the production system, especially when high-speed on-line dimensional

control is required.

Table l. Summary of the applications areas for non-contact, optical profilometry.

INDUSTRY DIMENSION OF THE FIELD (m) ACCURACY (mm)

Automotive lxl 0.5

Sfiipyards 20x20 10.0

Aeronautics lOx 10 2.0

Mechanical lxl 0.3

Microelectronics O.lxO.l 0.001

When a whole-field profilometer is intended to operate in an industrial environment, one

must take into account that: i) surface measurements have to be done in relatively small times,

especially when an on-line dimensional control of a surface is required; ii) good precision is

required; and, iii) the profilometer should be flexible enough to adapt itself to measure surfaces

differing in shape. Moreover, the complexity of the target to be measured, the presence of

shadow regions along preferred directions, changes of illumination and nonuniformities of the

reflectivity of the surface have to be managed. These characteristics, correlated with the

particular environment, suggest that the system must offer a large degree of adaptiveness.

In this chapter, a review of optical profilometry techniques is presented, with particular

emphasis on those based on grating projection methods. We also discuss how adaptiveness has

been added to an optical whole-field profilometer designed for industrial applications.

2. Whole field optical profilometry techniques and systems

Using optical, non-contact methods the tree-dimensional profile of the surface of an object

can be measured in two ways: by scanning techniques and by imaging techniques [4).

2.l. SCANNING SYSTEMS

Surface shape is determined at each point by a device that measures the distance between

the light source (usually collimated) and a point in space, along the direction of the beam. A

353

scarming system is needed in order, to cover the two-dimensional field. Among developed

techniques, those based on triangulation methods [5], on heterodyne interferometry [6] and on

laser lidar methods [7] have produced a number of systems which are currently used in industrial

applications [8].

2.2. IMAGING TECHNIQUES

In imaging techniques, one or two whole-field intensity-based images of the object, are

acquired. The images are then suitably analyzed in order to obtain the object profile. Many

techniques have been developed, providing different illumination schemes and implementing

different software analysis procedures. Among these, those based on stereo vision and grating

projection methods are described in the following.

2.2.1. Stereo-vision. The three-dimensional profile of an object can be determined by stereo

vision methods. In principle, the concept is that of observing the field at two different angles and

locating corresponding points in both images [9,10). Generally, full-field maps are difficult to

obtain by using these methods, which are rather slow.

2.2.2. Projection of grating. Grating projection represents a promising method to measure the

profile of the object surface. As shown in Fig. 1, the projection of a fixed, periodic pattern (a

grating) on the surface to be measured, and the observation of the resulting pattern from a

different perspective, produce a deformed pattern, according to the topography of the object. In

order to extract 3-D topographic information, the 2-D deformed image is stored by a video-

Macro Profilometry Area

Fig. 1. Schematic of the approach to profilometry.

354

camera and then analyzed. Analysis can be performed by means of a matched reference grating,

leading to the analysis of the so-called Moire fringe patterns, or by a direct analysis of the

deformed grating itself, which leads to the so-called direct methods:

Moire topography - In the Moire fringe pattern, contour lines of equal separation are produced on

the object [11,12]. Several techniques have been developed to automatically measure these

contour lines, mainly in order: i) to distinguish a depression from an elevation on the object; ii)

to make fringe order assignement, and iii) to increase both measurement accuracy and resolution

[13,18].

Direct methods - The use of direct methods led to the development of a number of analysis

algorithms. Among these, particularly interesting are those based: on Fourier transforms [19]; on

real-domain demodulation and processing of the grating, [20]; and, on phase-shift methodologies

[21,22]. All these techniques basically present the same optical geometry, and allow one to

obtain the surface profile by evaluating the phase perturbation introduced by the object. In the

following, both the schematic layout of illumination and detection geometry and the phase

evaluation procedures are described.

3_ Whole field optical profilometry with direct demodulation: system approach

3.1. OPTICAL GEOMEIRY

A typical illumination and detection geometry is shown in Fig. 2. The projection and

imaging optics are placed at points P and C respectively; d is the distance between them and L

the distance to plane R, having the coordinates x in the plane of the figure, and y in the direction

normal to the figure. This x-y plane is the so called reference surface, with respect to which the

height of the object is evaluated. In principle, the method of measurement is a triangulation

method: a ray striking plane R at point B is seen by the video camera as if it originated from

point B, when no object is placed on plane R. In the presence of the object, point B moves to

point A on plane R. Correspondently, a phase-modulation term, strictly related to the shape of

the object, is present in the object-deformed pattern. The height of the object at any point H(x,y)

can be evaluated by the formula: (ZH + L) / d = ZH / BA, which takes into account that triangles

PHC and BHA are similar. In this formula, ZH represents the height of the object at point H, L

and d are the geometrical parameters of the system, and BA represents the shift of the ray,

which is proportional to the phase shift term in the deformed pattern.

355

grating.f image Pli~ne /\\ d 11

p"- '.rc \ \ ! i i \. ! i i \. I . i \. i i \. i i \ i i \.1 i \. ! \ '/.. \ i \ \ i \ \ . \ i ! \ i / \. . . __ ~. -+· ______ ~\~~----~------x

;r-BL/A 0

~l~ OBJECT

L

Plane R

Fig. 2. Optical geometry of the projection and recording system.

3.2. PHASE DEMODULATION TECHNIQUES

The basic technique using phase demodulation is based on the demodulation of a

defonned, square grating, by means of the Fourier Transfonn [19]. Fig. 3 shows the schematic

flow of the procedure. Each row of the 2-dimensional acquired signal is Fourier-transformed, the

proper spectrum component is selected, and then the inverse Fourier-transform is applied. In

order to separate the phase term from unwanted surface reflectivities, the complex logarithm of

the signal is computed. The phase information is now contained in the imaginary part of the

signal. In order to eliminate phase discontinuities of 21t, are eliminated by a suitable

unwrapping algorithm.

This procedure automatically distinguishes between a depression and an elevation in the

object surface. Moreover, fringe order assignement and data interpolation in the regions between

contour fringes are not required. On the other hand, measurements of stecp object slopes are

difficult, and particularly sophisticated hardware is required. An alternative approach is to

process the fringe pattern in the real domain instead of the frequency domain, following the

procedure depicted in Fig. 4 [20]. Again, the image is analyzed by rows. Two signals are

obtained from a single row, by multiplying it by the sine and cosine functions, whose frequency

is equal to (or close to) the grating carrier frequency. Both signals are low-pass filtered. The

phase information is obtained by evaluating the arc tangent of the ratio between the two signals.

356

PHASE

Fig. 3. Schematic diagram of the Fourier demodulation procedure.

PC2

VGA-PAL converter

PC 1

Video-Camera

PHASE

Fig. 4. Schematic diagram of the real-domain demodulation procedure.

Monitor

Fig. 5. Schematic diagram of the profilometer layout.

357

This technique can be easyly implemented in computer software, and this improves the speed of

analysis, making the system particularly suitable to industrial applications, in which processing

time is often critical. The phase-shift method [21, 22] is based on the projection of sinusoidal

gratings. By using several phase-shifted frames of deformed grating image data, a high degree of

precision in the phase measurement is achieved. Fringe order assignment is required.

All the techniques mentioned are characterized by: 1) good accuracies even with coarse

gratings (period 10 - 15 mm); 2) limitations in the maximum resolution range achievable,

depending on the maximum slope of the object, which in tum depend on the grating period; and,

3) both fixed pattern projection and non adaptive demodulation procedures. As mentioned above,

an adaptive profilometer could be very useful for industrial applications. Specifically,

adaptiveness of the projection could provide: 1) the possibility of managing environmental light

changes and correcting for surface nonuniformities of reflectivity before image acquisition. A

pre-corrected image would be acquired, and this would greatly shorten the overall analysis time;

and, 2) the possibility of modifying the period of the grating, which could decrease the limit of

resolution determined by the slope of the object. In fact, coarse fringes could be used to measure

large surface variations, while fine fringes could be used to measure fine surface details.

Adaptiveness of projection requires adaptiveness of the demodulation. The system should

present an algorithm able to vary the demodulation parameters according to the variations

imposed by the adaptivess of the projection. A possible solution is represented by: i) designing a

suitable, adaptive, projection unit; ii) choosing the fastest and easiest to implement algorithm

provided that both resolution and accuracy match the measurement requirements; and, iii)

modifying the algorithm itself in order to add adaptiveness. With these goals in mind, an

adaptive whole-field profilometer has been developed at our laboratory. The profilometer

provides an adaptive projection unit based on a Liquid Crystal Device (LCD) and implements an

adaptive demodulation technique based on the real domain demodulation method described

above. The main features of the system are presented in the following section, with particular

emphasis on the characterization of the LCDs and the realization of the projector.

4. The profilometer

The complete structure of the system is presented in Fig. 5. The system follows the

crossed-optical-axes geometry described above. The projector illuminates the object placed on

358

I /

I /

I / /

/ / / /

/ /

/ /

/ /

/ /

/

/ /

/

VIDEO-CAMERA

/~~-=-=--~ \ \ -~-

\ -'-\ \

\ \

\

L

Fig. 6. Example of object-defonned grating.

Fig. 7. Acquired image.

359

the reference plane. The CCD-matrix camera detects the 2-D perturbed fringe pattern, which is

frame-grabbed by a Personal Computer (PC1) through an image processing board. The monitor

allows direct image visualization. The user can directly select the most suitable pattern to be

projected by the software of PCl and then send the chosen configuration to the projector,

through the dedicated Personal Computer PC2 and the VGA-P AL converter.

4.1. THE SOFfW ARE INTERFACE

A suitable software interface has been developed. It is based on a series of nested menus.

Among the menus provided, the most important ones perform the following actions: 1) image

acquisition and pre-processing, which basically consist of background subtraction and image

normalization; 2) object height measurement, implementing the demodulation algoritlun; 3)

graphic 3-D visualization, giving the representation of the measured object; and, 4) pattern

generation and driving of the projection unit. Fig. 6 shows, as an example, the grating

deformation due to the shape of an object under measurement. Fig. 7 depicts the corresponding

acquired image. In Fig. 8 the reconstructed image is presented. In this example, the illuminated

area is equal to O.75m by O.75m, the acquisition and projection units are placed 3 m from the

object, and a square grating of period equal to 11.36 mm is projected, corresponding to 5.88

pixel per period on the video-camera (a 512 by 512 CCD camera is used).

Fig. 8. Object profile reconstruction.

o III

750 o Y·AXIS

~~250 500

360

.. 00 20 S 8~ .! i 300 I t 15 , 6 c: ;; ~ I ~ ~ 10 a) ~ 200 , E ~ I 4 ~ 8 ~ 100

J ~ 5 , 2

~ ~

-----~ 00 DO 2 .. 6 0 2 4

b)

6 Brightness Voltag. (Volt) Brlghtn.ss Voltage (Volt)

2 40 -;::.

~ 8~ := 1.5 30 .. /I § .. 6- ..

Iii I ~ ~ 20 C) i .. E c:

~ 8 ~ 0.5

, ...... _ .. 2! 1O

~ \. 0

D° ~,.2 ·1.2 48 ·0." -0.11 -0."

d)

o Brightness Voltage (Volt) Brightness Voltage (Volt)

Fig. 9. Transmission and contrast curves for the passive matrix monitor (plots a, b) and for the active TV display (plots c, d) at 21°C. Light transmission is monitored, in both cases, for white (solid circles) and black (solid squares) pixels as a function of the brightness voltage.

30 6 50 0 .. 0:- .. 0:-Il ... III ...

~ 24 ., , ~ ~ 40 ~ c: \, '- c: '-

8 ,

4 : 0 -0.2 " , (.) 30 " 18

~, • ... ~

oS ...

i b) 0 0

a) ! i III ~ 20 ~ ... 12 , l: ~ -0.4 Qj ~

.. 2111 , , ,

~ "' ... ~ ole \ ole

~ III 10 \ III

I 6 ~ "'-", ........ -- ..... ~ .... It---.

00 10 20 30 40 500 00 10 20 30 40 511.6

Temperature (OC) Temperature (OC)

Fig. 10. Plots of the peak value of contrast (solid line, solid circles, left scale) and of the brightness value giving the peak contrast (dashed line, solid squares, right scale) for the displays under test, as a function of temperature. a) Passive matrix display; b) active matrix display.

361

4.2. LCD CHARACTERIZATION

In order to develop the projection unit, we made a comprehensive characterization of

Liquid Crystal Displays. Super twisted-nematic displays were chosen for this application. We

focused our attention on commercially available, low-cost and high pixel number LCDs. They

are used as video-camera viewers, pocket TV's and computer monitors [23]. We compared

passive-matrix and active matrix LCDs.

The passive LCD was a computer monitor display, while the active LCD was a display

mounted on a commercial liquid crystal TV. Both LCDs were driven by a high-level personal

computer interface. The former was driven directly, through its in-built VGA interface, the latter

was driven through a VGA-to-PAL converter.

Active matrix displays have proven to be superior to passive matrix displays in terms of:

(i) transmittance and contrast; (ii) dependence of the transmission curves upon temperature; (iii)

spatial characteristics of transmission; and, (iv) transient behavior of the transmission.

In the following section, the main results of this characterization are summarized; they are

discussed in more detail in Ref. [24, 25].

4.2.1. Transmittance and contrast measurements

Due to the influence of fringe contrast on measurement accuracy, both transmission and

contrast of the two devices have been studied. Typical results of static transmission

measurements are shown in Fig. 9. Plots a) and c) show the transmittance of a white (solid

circles) and of a black (solid squares) pixel as a function of the brightness signal for the passive

and active matrix devices respectively, at 21°C. Plots b), and d) show the dependence of the

dynamic contrast upon the voltage brightness for the two LCDs.

4.2.2. Temperature dependence of transmission

To assess the performance of an LC display when inserted in a light pattern projector for

an industrial environment, it is important to evaluate how the display parameters change with

temperature. We did measurements for various temperature values ranging from 6 to 40.5 °C.

The plots in Fig. 10 express the peak contrast dependence on temperature of both displays. They

show, for both displays, (i) the value and (ii) the position of the peak of the contrast curve as a

function of temperature.

362

4.2.3. Spatial transmittance distribution measurements

Fig. 11 shows typical examples of near-field patterns of sequences of black and white

pixels of the two LCDs, taken at the two extreme values of the brightness voltage for each LCD.

Figs. Ita) and b) show near-field patterns for the passive matrix display. Note the increase of the

centred black pixel transmittance at high brightness voltage. The pixel transmission of the

passive LCD shows a dip at the center of the pixel. This is the so-called "pixel segmentation"

effect, and is determined by the presence of the signal-carrying conductors for rows and columns

[26]. Fig. Ilc) and d) show near-field patterns for the active display. Here, the absence of the

pixel segmentation effect is evident (the connectors are deposited in the inter-pixel space). Fig.

lId) points out that the contrast at the high brightness level is still high, unlike the previous case.

The results obtained from the single pixel transmission measurements are confirmed.

4.2.4. Temporal dependence a/transmission

When an LCD has to be used for adaptive projection, the response to a change in the

>5OO"S

a) b)

'\.) t W '~ ~J ¥ \~ : ________ 1

H

c) d)

Fig. 11. Near-field pattern of the transmission of a sequence of two white pixels and of a black pixel of the two LeOs considered, for two different values of the brightness voltage. a) passive matrix display, 1.75 V; b) passive matrix display, 5.4 V; c) active matrix display, -0.7 V; d) active matrix display, -0.1 V.

363

b)

a)

Fig. 12. Transient characteristics of the two LCDs:a) passive matrix display; b) active matrix display.

projected pattern should be compatible with the frame time of the acquisition video camera. We

compared the settling times of the two LCDs, by blanking a set of pixels and by monitoring the

transmission as a function of time. Fig. 12 shows the transient characteristics of the two displays

under test. The passive matrix display (Fig. 12.a) exhibits an off-to-on rise time of about 350 ms,

and an on-to-off decay of about 310 ms. The active matrix display, conversely, exhibits a more

prompt response (on-to-off = off-to-on) of less than 40 ms, equivalent to 1.2 frame periods.

Superiority of the active-matrix display over the passive matrix display is quite evident from

these results . Specifically, the absence of the phenomenon of pixel segmentation in the active

matrix device results in a uniform light distribution across the pixel width. This aspect is relevant

in applications involving fringe demodulation.

5. Conclusions

Optical profilometry is now reaching its full maturity, mostly due to irmovations in

electrooptical components and systems and to the high speed of modem personal computers and

mainframes. It is envisaged for the near future, that true real time operation of profilometers will

become possible and this will greatly favour the use of such intruments for in-process

dimentional control.

Most beneficial for the success of optical profilometers with adaptive projectors will be a

greater attention of LCD manufacturers to this field. The generation of static patterns whith a

high contrast, by means of low-cost display with good thermal characteristics, will be strictly

required for high accuracy systems which can effectively substitute non-adaptive profilometers.

364

References

1. Langenbeck, P., (1982): 'In-Process Optical Metrology for Precision Machining', Proc. SPIE, 802.

2. Fagan, W.F., (1987): 'Industrial Optoelectronic Measurement Systems Using Coherent Light', Proc. SPIE,863.

3. Boehnlein, AJ., et al. (1987): 'Closing the loop between metrology and CAM', Proc. Sensor Exp. 345.

4. Bieman, L.H., (May 1988): Three-Dimensional Machine Vision', Photonics Spectra, 81-92 5. Blais, F. and Rioux, M.,(1986): 'BIRIS: A simple 3-D Sensor', Proc. SPIE 723-735. 6. Sommargren, G.E., (1981): 'Optical heterodyne profilometry', Appl. Opt., 20, 610-617. 7. Svetkoff, D.J., Leonard, P.F. and Sampson, R.E. (1984): ' Techniques for real-time, 3-D feature

extraction using range information'. Proc. SPIE, 521, 302. 8. Wesolowicz, K.G. and Sampson, R.E., (1987): '3-D Imaging Sensor for Robotic Applications',

Proc. of the 5th ICALEO, IFS Publications 9. Hersman, M. et al. (1987): 'Coherent laser radar application to 3-D vision', VISION '87, SME

MS87,385-391. 10. Hobrough, G. and Hobrough, T., 'Stereopsis for robots by iterative stereo image matching',Proc.

SPIE, 449-94. 11. Takasaki, H., (1970): 'Moire Topography', Appl. Opt., 9,1467-1472. 12. Meadows, D.M., Johnson, W.O. and Allen, 1.B., (1970): 'Generation of surface contours sy moire

patterns', Appl. Opt., 9, 942-947. 13. Idesawa, M., Yatagui, T. and Soma, T., (1977): 'Scanning Moire method and automatic

measurement of 3-D shapes', Appl. Opt., 16,2152-2162. 14. Moore, D.T. and Truax, B.E., (1979): 'Phase-locked Moire fringe analysis for automated

contouring of diffuse surfaces', Appl. Opt., 18,91-96. 15. Cline, H.E., Holik, AS., and Lorensen, W.E., (1982): 'Computed-Aided surface reconstruction of

interference contours', Appl. Opt., 21, 4481-4488. 16. Cline, H.E., Lorensen, W.E. and Holik, AS., (1984): 'Automatic moire contouring', Appl. Opt.,

23,1454-1459. 17. Indebetouw, G., (1978): 'Profile measurement using projection of running fringes', Appl. Opt.,

17,2930-2933. 18. Dirckx, J.J.J., Decraemer, W.F., and Dielis, G., (1988): 'Phase shift method based on object

translation for full field automatic 3-D surface reconstruction from Moire topograms', Appl. Opt., 27, 1164-1169.

19. Takeda, M. and Mutoh, K., (1983): 'Fourier transform profilometry for the automatic measurement of 3-D object shapes', Appl. Opt., 22, 3977-3982.

20. Tang, S.and Hung, Y.Y., (1990), 'Fast profilometer for the automatic measurement of 3-D object shapes', Appl. Opt., 29, 3012-3018.

21. Srinivasan, V., Liu, H.C., and Halioua, M., (1984):'Automated phase-measuring profilometry of 3-D diffuse objects", Appl. Opt., 23, 3105-3108.

22. Srinivasan, V., Liu, H.C., and HaIioua, M., (1985): 'Automated phase-measuring profilometry: a phase mapping approach', Appl. Opt., 24, 185-188.

23. Liu, H.K. and Chao, T.H., (1989): 'Liquid crystal television spatial light modulators', Appl. Opt., 28,4772-4780.

24. Sansoni, G., Docchio, F., Minoni, U. and Viviani, N., (1992):'Characterization of commercial liquid crystal displays for adaptive pattern projection in industrial profilometry', International Journal of Optoelectronics.

25. Docchio, F., Sansoni, G., Minoni, U., and Viviani, N., (1992): 'Light-induced transmission changes in liquid crystal displays for adaptive pattern projection', IEEE Transactions on Instrumentation and Measurement.

26. Liu, H.K. and Chao, T.H., (1989): 'Liquid crystal television spatial light modulators', Appl. Opt., 28,4772-4780.

ELECTROOPTICAL SYSTEMS AND TECHNIQUES FOR DIMENSIONAL MEASUREMENTS FOR INDUSTRY

F. DOCCHIO, U. MINONI, G. SANSONI, E. GELMINI Dipartimento di Elettronica per /'Automazione Facolta di Ingegneria Universitii deg/i Studi di Brescia Via Branze, 25133 Brescia, Italy.

ABS1RACT. In this chapter a review of the current electrooptical techniques and systems for measurement and control are presented, with particular reference to the industrial applications of these techniques. Triangulation, shadow casting, interferometry and profilometry are considered. The major advancements in each field are reviewed, in relation to: i) the activity carried out in our laboratory; and, ii) the applications domain. Finally, the prospects for electrooptical techniques in view of technological breakthroughs in optical components are discussed.

1. Introduction

Modem manufacturing systems are characterized by a large amount of flexibility, as well

as a continuously increasing quality of the end product, thanks to in-process and off-line quality

control and production diagnosis [1]. The achievement of both these features has been made

possible by the availability of sensors, systems and techniques for the dimensional control of the

working tools and of the workpieces, during and after the manufacturing process [2].

The trend is particularly marked in the case of laser manufacturing. The non-contact nature

of the laser process calls for accurate sensing, to profit from optimum laser-material interactions.

Optical measuring techniques are particularly important, in the overall context of laser- and non­

laser flexible manufacturing and quality control. In fact, their impact on the market is constantly

increasing. This is also due to the increasing number of devices, techniques and components

presently available.

365

S. Martellucci et al. (eds.), Laser Applications for Mechanical Industry, 365-379. © 1993 Kluwer Academic Publishers.

366

2. Overview of optical metrology

Optical dimensional measurements take advantage of all the characteristic properties of

either coherent or incoherent light. The principal optical measurements for industrial applications

in terms of the above mentioned quantities are: i) Intensity; ii) Phase/amplitude; iii) Polarization;

iv) Spectrum of radiation emitted, diffused or reflected; and, v) Time of flight.

Position measurements rely mainly on the straight-line propagation of light in unperturbed

media. Any deviation of a light beam from its original direction is due to misalignment and/or

reflective index modifications, as a result of any perturbations induced by the environment.

Intensity measurements mainly rely on the fact that a collimated beam can propagate virtually

unchanged if non-absorbing or dispersing media are present. The measurement of the beam

intensity is therefore useful to measure the amount of light that has undergone such changes.

Measurements based on phase/amplitude determination probably have the most accuracy and the

highest resolution. They profit from the wave nature of light and make it possible to measure

distances or changes in the position of an object on a micron, sub-micron and almost Angstrom

scale. Polarization measurements, on the other hand, are based on the fact that a polarized beam

of light remains polarized until it encounters a medium which is able to alter the polarization

state of the wave. Therefore, polarization measurements are useful to increase the resolution or to

add extra information to either intensity or phase/amplitude measurements. Spectral

measurements rely on the ability of the media under test to disperse light according to its

wavelength or to convert monochromatic light into polychromatic light by virtue of their internal

atomic/molecular structure. Therefore, spectral measurements are particularly useful for

combustion, luminescence, and chromatic aberration measurements. In tum, spectral

measurements can be further distinguished into measurements of emitted light, diffused light,

and reflected light. Finally, time-of-flight measurements are based on the fact that, in a

controlled environment, light travels at a known speed. Thus distance, or distance variations, can

be measured by measuring the time interval required for a light pulse to traverse the distance

under test.

Almost all the above mentioned characteristics of electrooptical metrology are made

possible by to the unique characteristics of laser radiation. Lasers, in fact, are able to emit light

both temporally and spatially coherent, that is, monochromatic and collimated. They are able,

under certain circumstances, to produce very short light pulses (and are therefore suitable for

time-of-flight measurements). Finally they are able to emit tunable radiation, which makes

possible spectral measurements.

367

The most relevant devices for optical measurements are: i) Optical components (miniature

optics, holographic optics, diffractive optics); ii) Optical fibers (single-mode fibers, polarization

maintaining fibers, couplers, graded-index lenses); iii) Photo-detectors (photo-diodes, photo­

multiplier, intensifiers, CCD camera); and, iv) Modulators (integrated modulators).

Enhancements in the technology such as miniature optics, holographic optics and diffractive

optics provide significant improvments in light processing and handling. Fiber optics are a

particular type of optical component, which greatly increases the flexibility of light delivery onto

a target, and which may constitute the basis for optical sensors as well as for laser material

processing. Single mode fibers polarization-maintaining fibers, together with suitable light

couplers and graded refractive index lenses, are among the most outstanding recent advances in

optical fibers. Photo-multipliers and semiconductor photo-diodes are the most widespread types

of photo-detectors. Semiconductor photo-detectors can now be integrated to form detector

matrices (as in TV cameras or intensified cameras). Finally, light modulators are of particular

interest in order to produce phase- or frequency-shifts of the carrier light beam in those cases

where the laser source cannot be directly modulated.

Coming to the specific topic of this chapter (Le. laser-based dimensional measurement

systems), Table I summarizes the techniques which form the basis of modem optical

measurements for use in the industrial context.

Table 1 Summary of optical metrology techniques for industrial measurements.

Techniaue Sources Ran~e Resolution Repeatibility Avvlications Shadow casting Gas,Diodes mm tocm IJm tomm tens oflJm Wires, shafts Triangulation Gas,Diodes mm tocm /lm to mm IJm to mm Diameters,

surfaces Incremental Gas, Diodes, IJm to Km nm to /lm nm to mm Calibration, Interferometry YAGs displacements,

robots Absolute Diodes, Gas cm tom /lm to mm /lm to mm Distance Interferometry YAGs measurements,

robots Profilometry Gas Depends on !Jm to mm !Jm to mm Surface

technique monitoring, dimentional vrofilometrv

Dimensional measurements are usually based on: 1) shadow casting techniques, that is, the

measurement of the shadow projected by an object; 2) triangulation techniques, which make use

of the straight-line propagation of light; 3) interferometry, which makes use of the wave-like

nature of light and can be further subdivided into (i) incremental, or relative, interferometry and

368

(ii) absolute interferometry; and, 4) profilometry, which can be based either on direct fringe

projection and demodulation by the detector system, or on optical demodulation of the projected

fringes (Moire techniques). As one can see by comparing the techniques in Table 1, the range of

measurement, the resolution and the repeatability vary consistently according to the technique,

and as a consequence their application domains also vary. In the following sections, we present

an overview of the different techniques investigated and developed in our laboratory within the

context of multi-partner projects.

3. Triangulation techniques

A schematic diagram of the basic layout of dimensional probing using optical

triangulation is shown in Fig. 1 [3.4]. If a beam of light is projected onto a target and the light

spot produced is observed by an imaging system at an angle e, the lateral position Ax of the

resulting spot in the image plane depends on the height of the target. The relation between

vertical and lateral displacements is given by the simple trigonometric relation Ax = Az tan (0).

A number of devices based on optical triangulation are available today, and constitute a powerful

set of distance and displacement sensors whose measurement range varies from a few millimeters

up to tens of centimeters, and whose resolution is generally a few times 10-4 of the measurement

range.

3.1. TRIANGULATION TECHNIQUES COMBINED WITH TIME-OF-FLIGHT TECHNIQUES

An innovative approach to point-based laser triangulation techniques is given by the

Lambda probe [5], which is substantially based on a time-of-flight triangulation technique. The

principle of operation of this measurement head is shown in Fig. 2. A laser beam is scanned

across an angle. Two photo-detectors are used in the head: one photo-detector (the start

photodetector) is illuminated when the laser beam is reflected by the mirror placed at point P and

virtually generates a light spot on a so-called virtual (or reference) plane at the intersection point

of this plane with the axis of the imaging system. The measurement photo-detector is placed at

the image side of a lens and is illuminated when the laser beam generates a light spot onto the

target surface at the intersection point with the axis of the imaging system. Therefore, the

distance of the target with respect to the reference plane is given by the time interval between the

start signal and the stop signal, each of which is provided by the respective photo-diode. The

IIIuminatin beam

IMAGE STOR ED BY THE VIDEO CAMERA

369

Fig .. 1. Basic layout of dimensional probing using optical triangulation. A I-D, or 2-D light pattern impinges onto the surface to be probed; a CCD camera, placed at an angle with respect to the illumination direction, acquires the light pattern deformed by the presence of the object to be measured.

system just described can operate in a range from one hundred millimeters up to few hundred

millimeters with an accuracy of about 10-5 of the range. The main limitations to the

performances of the technique based on time-of-flight are, to date, the fixed focal length of the

imaging system, and the fixed focal position of the imaging system itself. This fixed focal length

results in a fixed range of measurement. On the other hand, the fixed focal position results in

errors of measurement due to the dependence of the leading edge of the stop signal.

z

Fig. 2. Optical diagram of a laser triangulator based on the technique of time-of-flight. bO, b, b I: start, current and stop angle, respectively. z: position of the object plane with respect to the reference plane. DI, D2: stop and start photo-detectors respectively.

370

Studies currently perfonned in our laboratory, in collaboration with Microcontrol S.p.A.

[6], are aimed at overcoming the limitations in perfonnance of the sensing head by introducing i)

an auto-zooming system to increase the measuring range of the device and ii) an auto-focusing

system to keep the spot size at the image plane as small as possible, and relatively independent

of the position of the target. An example of the results obtained by the use of an auto-focusing

system is given in Fig. 3. In Fig. 3.a, in particular, the shape of the stop signal is given for three

different displacements of the target: it is evident in the figure that huge variations in the

25

20 ........ ::; ~ 15 .:;:;. .~ 10

Q) --.E 5

o -0.4

16 14

.-. 12 ::; ~ 10 .:;:;. 8 ·in a; 6 c

4 2

--39

I~ .. - - 36

- - - 33

- -- 24 .. , . . - - - 3

_~1 •• f- .~~ -~

-0.2 o 0.2 0.4

Distance from center (cm)

-Z=1

- - Z=22

---Z=40

.... o ~ ..... --..... ~ ..... ~~ ..... ~ ..... --..... ~ -0.06 -0.04 -0.02 o 0.02 0.04 0.06

Distance from center (cm)

Fig. 3. Plots of the intensity distribution of the light spot measured onto the image plane as a function of the distance of the object plane from the reference plane. a) no auto-focus; b) auto-focus.

371

detennination of the stop pulse occur by defocusing. In contrast, Fig. 3.b shows the same

situation with the use of the auto-focusing system. The dramatic increase in precision of

detennination of stop signal edge can be observed.

3.2. PROFILOMETRY USING TRIANGULAR TECHNIQUE

Triangulation techniques can be successfully applied to 2-D or 3-D topographical analysis

of solid objects. 2-D triangulation (as again in Fig. 1) consists in projecting a light pattern onto a

target, and in monitoring at an angle 6 the deformed image resulting from the intersection of the

light pattern with the target surface. 3-D measurements are based on an extension of the 2-D

concept. In this case, a two-dimensional pattern is projected onto the target surface and the

defonned light pattern is acquired and analyzed at an angle 6. Both analog and digital procedures

for the detennination of the 3-D topography are available. It is beyond the scope of this paper to

deal with these techniques, as this will be the subject of another chapter in this volume [7].

3.3. APPLICATIONS OF TRIANGULATION SCHEMES TO ROBOTICS

Simple triangulation schemes can be successfully applied for the measurement of

misalignment errors in industrial robots. In collaboration with the Department of Mechanics, a

detailed study of the misalignment errors of a SCARA robot is underway using triangulation as

the sensing technique, and using analytic models to interpret the results [8]. Misalignment errors

of the robot axes may cause inaccuracy in the position of the end-effector. Sampling the angular

position of the robot by means of triangulation as shown in Fig. 4 allows us to quantify these

angular errors. Fig. 5 shows a typical result of these measurements and the matching of the

measurement with the values predicted by the model.

4. Interferometry

4.1. INCREMENTAL INTERFEROMETRY

In interferometry, the distance to be measured is compared to a known distance by means

of phase measurements of a laser beam which is sent to the target with respect to phase of an

auxiliary beam sent to a reference mirror. The beams returning from reference and target are

372

SCARA ROBOT

Zo

350

_300 "t:J as ..

Yo

I-

Irl 250 Q .... >< -200 c o ~ 150 l­

=; ~ 100 I-(I) Ul as ...J 50 I-

Zz

BS

Fig. 4. Example of triangulation techniques applied to robot tracking and misalignment measurement. The deflections of the laser beam reflected from a plane mirror placed on the end effector of a SCARA robot are monitored at the screen plane by a camera.

u CAMERA

o theoretical

* experimental

Measurement points

Fig. 5. Plot of the laser beam deflections of the plane mirror mounted on the Robot end effector, for various measurement points. Theoretical values obtained by matrix inversion are compared to experimental values.

373

made to interfere and the required distance information can be derived from the interference

signal. Laser interferometry is a relative measurement: the interference signal Isig is a sinusoidal

function of the distance d and is periodic with period given by the laser wavelength (lsig oc cos<i>,

where <i>=2nd/A.). Therefore, in most interferometers the measurement needs constant tracking of

both the measurement and of the reference beam. A number of different techniques have been

developed in order to apply laser interferometry to various measurement situations (surface,

topography, point measurements, different ranges) [9-11]. One of the major advantages of laser

interferometry compared with triangulation and other types of distance measurement techniques

is that the accuracy of the measurement is independent of the measurement range. The range, in

fact, is limited only by the coherence length of the laser source used in the interferometer. He-Ne

lasers have generally been used to build interferometers, whereas C02 lasers used for infrared

interferometry. In recent years diode lasers have been increasingly used as interferometric

sources. the reason for this success being the increase in the coherence length of available laser

diodes (single longitudinal mode laser diodes) [12].

Most of the limitations of normal interferometry arise from the sensitivity of the

interferometric signal to fluctuations of the laser source and to the quality of the target.

Overcoming these limitations is possible by using: (i) phase modulation techniques; and, (ii)

heterodyne techniques [13]. A typical example of phase modulation techniques is given by the

optical layout of Fig. 6. In a standard Michelson interferometer one of the beams is modulated by

means of an electrooptical modulator operating at 10 MHz, which introduces a phase shift in the

beam. When the two beams are recombined, a modulated signal is produced; demodulation of

PBS MR

Fig. 6. Block diagram of the sinusoidal phase-modulation interferometer. EOM: electro-optical modulator; MR, RR: measurement retroreflector and reference retroreflector; PBS: polarizing beam splitter; fm: modulating frequency; P: polarizer; D: detector.

374

200r------------------------, 14

12

10

Frel 8

(% 6

4

2

0 0.2 0.4 0.6 0.8 ·200 .100 0 100 200

Time(s) Error (nm)

Fig. 7. Plot of the difference of the displacement as measured with the SPMI interferometer and as measured with a reference interferometer for an arbitrary displacement. a) temporal plot; b) statistical distribution of the data.

lnterferom.

Laser

Rlvelatore

Camera

Fig. 8. Block diagram of the optical layout combining a laser interferometer in the folded configuration to a laser triangulator, for the measurement of both misalignment and positioning errors in a SCARA Robot.

0.05 n-------------------------, 0.04

~ ~U~~~~~~AM~~~~~ ...s 0.03 ~. 0.02

fii 0.01

j OH-----------------------; .~ -0.01 c:>

-0.02 -0.03 L-______________ ...J

o 5 10

Time (s)

·0.07 ~ ...s -0.09

1l -0.11 E " g -0.13 g-o -0.15

-0.17 L-__________ ---'

1.8 6.8

Time{s)

11.8

Fig. 9. Experimental results obtained with the layout of Fig. 8. a) transient displacement in the proximity of the quiescent position. b) micro-displacements at the quiescent position due to the control loop.

375

this signal gives two signals in quadrature to each other. This allows one to extract both

displacement magnitude and direction. FurtheImore, the signal is less sensitive to fluctuations in

the laser intensity than is the standard interferometry signal. Modulating at 10 MHz makes it

possible to operate the interferometer at in range of the m/s. A typical example of the

perfoImance of this interferometer is given in Fig. 7: here, the interferometer has been cross­

calibrated using a commercial interferometer based on the Doppler shift principle [14]. The

accuracy of the system has been found to be better than 80 nm, corresponding to ').J8. The major

advantage of the phase modulation scheme with respect to heterodyne or other kinds of laser

interferometers is the fact that this system is capable of being integrated in a very small package,

and when this happens integated electrooptical modulators will become available at low cost.

4.2. APPLICATIONS OF INCREMENTAL INTERFEROMETRY TO ROBOT TRACKING

Optical interferometry has a number of actual and potential applications in industry

including robotics and manufacturing control. An example of robotic control by interferometric

techniques combined with triangulation techniques is shown in Fig. 8. Here the interferometer is

designed in a folded configuration, in order to tolerate lateral displacements of the end effector.

The interferometer monitors the static and dynamic micro-displacements of the end effector

under the action of the controller. The angular displacements of the end effector are monitored

using the auxiliary camera, which is illuminated by a fraction of the laser beam in the

measurement arm. The position of the spot is analyzed using triangulation methods as in the

previous section. A typical result of the interferometric measurement of the robot aIm is depicted

in Fig. 9 [15]. In particular, Fig. 9.a shows the transient displacement of the robot when

controlled by a step command. Fig. 9.b shows the steady state vibrations of the robot aIm and the

influence of the controller. It is believed that this and other measurements will allow us to obtain

successful calibration of the robots in relation to possible axis misalignment and positioning

errors. These measurements will be used to perform software compensation of such errors.

4.3. ABSOLUTE INTERFEROMETRY

The limitations of interferometry of being a relative measurement technique may be

overcome by the use of multiple-wavelength interferometric schemes. An interferometric

measurement can be made absolute when the measuring wavelength is made longer than the

measurement range. Since it is not possible to use real laser wavelengths on the order of

centimeters, the measuring wavelength is obtained in a synthetic way by combining two or even

376

0,

Ii,· 1i2

Fig. 10. Optical layout of a 2-wavelength absolute interferometer with heterodyne detection. AOI +A03: acousto-optical modulators.

more laser wavelengths. In general, when two waves at different frequencies are forced to beat,

after low-pass filtering only the difference between the two frequencies of the waves remains,

and this is equivalent to a synthetic wave whose wavelength is 1.11.210"1-1.2). It is obviously

clear that the closer the two wavelengths are, the longer the synthetic wavelength is. In absolute

interferometry, distance or displacement measurement is then accomplished by means of phase

measurements within the synthetic wavelength [16,17].

Fig. 11. Example of an absolute distance measurement using the 2-wavelength interferometer. ,.,

:[ Q) 4.5

~ .s::; 0..

3L-______ ~ ____ ~~ ____ ~ ______ ~------~

o ~ ~ ~ ~ ~

Displacement (mm)

377

A typical laser interferometer using the two-wavelength approach [18] is shown in Fig. 10.

Each wavelength from a He-Ne or semiconductor laser is heterodyned in order to set up a double

heterodyne scheme. Fig. 11 shows results of absolute measurement within a range of about one

meter using the above scheme. In an absolute interferometer, the two wavelengths can be

generated either using two different lasers which are frequency-locked or using a single laser

source followed by an acousto-optic modulator. Work is in progress in many centers to obtain

multiple-wavelength interferometers with a sufficient degree of compactness and reliability in

order to be used in the industrial environment.

5. Outlook and prospects for electrooptical dimensional measurements

There is little doubt that systems and techniques for optical metrology will dramatically

expand in the near future. Optical measurements in the past decade have been mainly confined

to the laboratory, and their role in the manufacturing environment has been limited. The reasons

for this limited impact have been: (i) the bulky nature of the optical components required; and,

(ii) the lack of ruggedness of most of the laser sources available on the market.

The development of semiconductor lasers opens new possibilities for optical dimensional

measurements. Semiconductor diodes, used at first only for triangulation techniques, are being

increasingly used for interferometry (due to their increased temporal coherence), and for time-of­

flight measurements (due to their direct modulation capability). It is therefore likely that, at least

in the field of dimensional measurement, the semiconductor laser will be the source of the future.

Detectors have kept pace with the light sources. Small-sized photomultiplier tubes, simple array

and matrix photodiodes, and avalanche photodiodes are now commercially available at costs

compatible with the overall system costs. When needed, image intensifiers with gating

capabilities are also available. In addition, computer interfacing of the matrix detectors is

facilitated by advanced array- and frame-grabbers.

However, the most dramatic breakthrough, which really makes it possible to expand the

role of electrooptic dimensional measurements, is the evolution of optical components. Fig. 12

summarizes progress in this area. From large-sized optical components, miniaturized

components are now available. Consumer optics have tremendously favoured both mass

production of miniature optics and the decrease in their costs (a typical example is given by the

sensing head of CD players).

378

Miniature ./I 0 ~ Optics LJ

Diffractlve Optics .... .AAA ........

-we&tn-

Integ~ted~ OPtlc"/2-Y

Fig. 12. Evolution of optical components for dimensional analysis in industry.

New kinds of optical components are available today. Holographic optics (presently

mostly computer-generated) are particularly useful whenever complex functions are required of

the optical components. Diffractive optics, whose development is now possible due to the

availability of micro-machining techniques, are likely to be the building blocks for very compact

optical systems, thus eliminating many bulky optical components. Finally. integrated optical

systems (the first devices are already available on the market) will contribute to the increase in

compactness, robustness, reliability and benefit to cost ratio of optical systems.

In conclusion, optical metrology will have a leading role in the context of automated

manufacturing, dimensional gauging and control, as well as calibration, and will certainly be

beneficial for increasing the potential of laser material processing systems.

ACKNOWLEDGMENTS. The authors gratefully acknowledge the contributions of Prof. C.

Bussolati and A. Taroni in the preparation of the manuscript. The authors also acknowledge the

contribution of Ing. U. Perini and G. Re Garbagnati of elSE S.p.A., of Dr. P. Biazzi and Dr.

Brusaferri of Microcontrol S.p.A., and of Prof. PL Magnani, Dr. R. Faglia. and Dr. G. Legnani

of the Department of Mechanics of the University. for their partecipation in joint activities.

379

References

1. Langenbeck, P. (1982): 'In-process optical metrology for precision machining', Proc. SPIE 802. 2. Fagan,W. F. (1987): 'Industrial optoelectronic measurement systems using coherent light', Proc.

SPIE863. 3. Brenci, M., Mencaglia, A., Mignani, A.G. and Scheggi, A.M. (1990): 'Sensori a fibra ottica

basati su tecniche di triangolazione', Proc. 1st Natl. Congr. 'Strumentazione e metodi di misura elettroottici', 67 -74.

4. Blais, F. and Rioux, M. (1986): 'BIRIS: A simple 3-D sensor', Proc. SPIE 723,235. 5. The Lambda System, Data Sheet, Applied Laser Tecnology. 6. Docchio, F. and Minoni, U. (1991): unpublished results. 7. Sansoni, G., Docchio, F., Minoni, U. and Biancardi, L. (1992): 'Adaptive profilometry for

industrial applications', published in this volume. 8. Faglia, R., Legnani, G., Docchio, F. and Minoni, U. (1992): 'Experimental evaluation of the joint

axes misalignment in a SCARA Robot by optical measurements', Mechatronic Systems Engineering.

9. Minoni, U., Sansoni, G., Docchio, F., Paganini, E., Perini, U. and Re Garbagnati, G. (1989): 'Interferometria laser per misure dimensionali e sue applicazioni industriali', Fisica e Tecnologia 12,113-131

10. Chen, J., Ishii, Y. and Murata, K. (1988): 'Heterodyne interferometry with a frequency­modulated laser diode', Appl. Opt., 27,124-128.

11. Minoni, U. and Docchio, F. (1991): 'A new SPM interferometer for displacement measurement at high velocity', Proc. 37th ISA '91, 1,997-1008.

12. Belov, M. 1., Gur'yanov, A. N. Gusovskii, D. D., Dianov, E. M., Kuznetsov, A. V., Pencheva, V. K. and Prokhorov, A. M. (1987): 'Investigation of single-frequency semiconductor lasers with a fiber Michelson interferometer', Sov. J. Quantum Electron., 17,549-551.

13. Docchio, F. and Minoni, U. (1992): 'A high-frequency sinusoidal phase modulation interferometer for displacement measurements', International Journal of Optoelectronics.

14. Minoni, U., Sardini, E., Gelmini, E., Docchio, F. and Marioli, D. (1991): 'A high-frequency sinusoidal phase-modulation interferometer using an electrooptic modulator: development and evaluation', Rev. Sci. Instrum. 12,2579-2583.

15. Avanzi, P. A. and Bertanza, G. (1990): 'Calibrazione di robot SCARA mediante sistemi ottici e tecniche numeriche di compensazione', Tesi di laurea, Universita degli Studi di Brescia, Italy.

16. Massie, N. A. (1987): 'Absolute distance interferometry', SPIE 816, 149-157 17. Dandliker, R., Thalmann, R. and Prongue, D. (1988): Two-wavelength laser interferometry using

superheterodyne detection', Opt. Lett. 13,339-341.

LASER VELOCIMETRY FOR COMBUSTION

D.F.G. DURAo, M.V. HEITOR Instituto Superior Tecnico Technical University of Lisbon Department of Mechanical Engineering Av. Rovisco Pais 1096 Lisboa Codex, Portugal

ABSlRACT. The most critical aspects of the application of laser velocimetry to the analysis of combusting flows are reviewed in this chapter and include: i) the presence of gradients in refractive index across the scattering medium; ii) particle seeding; iii) interferences with visible emissions; iv) window transmissivity in confined flows; v) uncertainties in the average performed; and vi) velocity bias due to non-uniform particle seeding in the reactants (and entrained) streams. Attention is focused on the analysis of homogeneous combusting systems in small-scale environments, and sample results obtained in premixed and non-premixed recirculating flames are briefly discussed to demonstrate the use of the technique. Its extension to the analysis of turbulent heat fluxes is briefly summarized and, in addition, recent developments which are expected to improve the analysis of combusting flows with practical interest in the near future are described.

1. Introduction

Laser Doppler velocimetry, LDV, has been extensively used for spatially- and temporally­

resolved measurements of velocity characteristics in a number of different laboratory-flow

configurations, ranging from simple boundary-layer flows to complex engine configurations, and

the most important implications of combustion for the successful application of the technique are

briefly reviewed in this paper. Other recent reviews with emphasis on the discussion of relevant

results obtained in a range of com busting flows, including those typical of liquid fuels, have

been published [1,2].

The continuing interest in this area arises from the increasingly stringent requirements of

improved efficiency, reduced emissions and the use of advanced combustion systems and

381

S. Martellucci et af. (eds.), Laser Applications for Mechanical Industry, 381-401. © 1993 Kluwer Academic Publishers.

382

alternative fuels [3], which call for improved understanding and the related need for diagnosis

[4]. It has been emphasized in the past [5,6] that the information to be obtained in practical

combustion systems includes the spatial and temporal distribution of the characteristics of the

combustion mixture before and after ignition, which requires knowledge of the mean flow,

turbulence levels and fuel burnout and dispersion. While intrusive probing for velocity

characteristics is inadequate for most com busting flows of practical interest, which include those

with wnes of recirculation, laser velocimetry is the most reliable technique, and one which has

unique capabilities when used either at fundamental or applied levels. The principles of the

technique and its more general practice have been extensively described in the literature and the

reader is referred to Refs. [7-9] for further information.

The LDV technique exploits the combined effects of scattering particles and the incident

light in the intersection region of two laser beams, where a fringe pattern is formed. It uses the

properties of monochromaticity, coherence and directionality of laser light (together with linear

polarization in some occasions), and continuous wave lasers are commonly used in the red, green

or blue lines with an output power which is dependent on the flow configuration under analysis

[10]. A typical configuration is shown in Fig. 1, for which the light scattered from the particles

in the flow is collected and focused onto a photodetector, which provides a Doppler signal with a

frequency proportional to the instantaneous velocity of the particles. Other optical configurations

can be used (e.g. light collection in the backscatter direction), but it is important to note that it is

the velocity of the particles which is measured and, for this to correspond to the flow velocity,

the particles must be small enough to follow the flow. Typical time scales for the flows

considered in the scope of this paper range from 100 !lsec in laboratory flames at atmospheric

pressure to about 0.1 !lsec in jet engines and combustors at design pressures of 10 atm [11],

which are associated with characteristic lengths varying between 0.5 mm and 10 !lm,

respectively.

While the spatial resolution of the technique is derived from the geometry of the collecting

optics and from the dimension of the velocimeter measuring volume (say, 1 x 0.1 mm), the

necessary temporal resolution can only be achieved if submicron particles are used once the

available detectors are fast enough to follow the type of temporal fluctuations mentioned above.

On the other hand, the particles must be uniform in size and shape, chemically inert and

characterized by a scattering efficiency independent of the flame characteristics in the range of

interest, and these aspects are discussed below on the basis of comparative studies making use of

different seeding materials. In addition, it is essential that the intersection of the two light beams

at their focal point be accurately known and that their interference pattern is maintained along the

scattering medium and, in practice, this may cause problems for the application of the technique

383

to combusting flows [121. Other problems, such as directional ambiguities, can be easily

resolved with the use of frequency-shifting techniques (e.g., as provided by a rotating diffraction

grating or acoustico-optic modulator) and are not discussed here.

It is evident that when laser velocimetry is used in a variable density flow, such as a

combusting flow, the measuring volume contains more particles when it corresponds to high

density gas (Le., cold flow), than when it corresponds to low density gas (Le., hot flow), so that

near-density weighted averages may be measured, and the extent to which this has been

analysed in the literature is also considered in this chapter.

The various aspects mentioned above are discussed in detail in the next section. Section 3

presents two cases studied, namely a swirling non-premixed flame and a bluff-body stabilized

premixed flame, which are used to demonstrate the successful application of laser velocimetry to

study the homogeneous combusting system associated with complex recirculating flows. The

TRANSMITING OPTICS (BEAM SPLITIER. FREQUENCY SHIFTER MODULE AND TRANSMITING LENSES)

LASER

FRINGE PATIERN AT THE MEASURING VOLUME

FLAME SEEDED WITH SCATIERING CENTERS

Fig. 1. Schematic diagram of the typical arrangement of a laser Doppler velocimeter capable of measuring time-resolved local velocity characteristics. A typical Doppler signal before and after filtering is shown.

384

analysis includes the extension of the technique to the analysis of turbulent heat fluxes through

its combination with other techniques for scalar measurements. Novel methods in laser

velocimetry which are expected to be helpful in com busting flows are analysed in Section 4, and

the last section presents the most important conclusions and recommendations derived from this

review.

2. Critical aspects of the application of laser velocimetry to combustion

The purpose of this section is to discuss the main implications of combustion for the

successful application of laser velocimetry. The analysis is based upon results presented in the

literature and includes the following topics: i) laser beam refraction bias; ii) the suitability and

efficiency of current particles as scatterer centers; iii) the effects of visible emissions in practical

combusting systems; iv) window transmissivity in confined flows; v) uncertainties in the

averages performed; and vi) possible bias effects due to non-uniform particle seeding in the

reactant streams of non-premixed flames.

a)

LASER

b)

LASER

UNBURNT REACTANTS

UNBURNT REACTANT

BURNT GASES

BURNT GASES

Fig. 2. Schematic diagram of the effects of the presence of a flame boundary between burnt and unburnt gases across the two transmission beams of a laser velocimeter, after Ref. [12]: a) The phase effect; and, b) The Schlieren effect.

385

2.1. GRADIENTS OF REFRACTIVE INDEX ACROSS THE SCATTERING MEDIUM

This problem was first addressed [12) on the basis that a flame boundary acts as a moving

optical interface which can modify both the phase and direction of a traversing laser beam. Fig. 2

illustrates these two aspects, considering the interaction of a dual-beam LDY system with a

turbulent boundary between hot and cold gases due, for example, to a jet of combustion products

emerging into the atmosphere. The first effect (Le., the "phase effect") is due to the changing

phase difference between the two beams, which occurs at a rate

dx ~ dy -z--dt D..y dt

(I)

where dy/dt is the flame boundary's upward velocity. The other effect is caused by the changing

deflection due to variations in the optical path gradient, which affects the closely adjacent beams

together. This is usually referred as the "schlieren effect", and it is evident that the related

deflection of the beams increases with the optical path length within the hot gases which,

ultimately, may preclude the intersection of the two laser beams.

However, considering a variation of the refractive index across the turbulent "phase

boundary" of the order of 3 x 10-4, it has been shown [12) that, for an average path length in the

hot gases of 5 cm, the phase and schlieren effects may give rise to spurious velocities 10% larger

than the convection velocity of the interface only for short periods during near-tangential

incidence of the beams to eddies in the interface. Otherwise, refraction-induced bias errors can be

assumed to have a negligible impact on the accuracy of velocity measurements, and this agrees

with the experimental analysis, making use of a combustion bomb [13) and a spark ignition [14),

respectively. The authors of Ref. [14] have however found a large decrease in the LOY data rate

when measurements were made over the longest path length available in the engine, probably

caused by a reduction in signal quality related to the thermal boundary layer at the window

interface. Measurements in larger flames [15), appeared to lead to similar conclusions, although

the path through the furnace led to a reduction in the intensity of light in the intersection region

of the two laser beams. Other current applications of laser velocimetry to combustion flows, at

least for unconfined small flames [2], have confirmed that combustion does not impose special

requirements on instrumentation other than that it should make best use of available signals and

avoid the usual ambiguities associated with noise and gradients. These conclusions suggest that

the theoretical analysis [l6) about the effects of fluctuations in refractive index on the accuracy

of a laser velocimeter is overestimated, probably because in practice a large proportion of

"spurious" Doppler signals are rejected by the validation criteria normally imposed by the signal

processing systems.

386

2.2. PARTICLE SEEDING FOR COMBUSTING FLOWS

The suitability of seeding particles to be used as light scattering centers when LDV is

applied to combusting flows represents an unique experimental difficulty. In addition to the

usual sub-micron size requirement, the particles introduced into a combusting environment must

survive high temperatures (at least 2000 K) and, if the flow is confined, must be both non­

abrasive to the walls of the combustion chamber and non-adhering to window surfaces. Naturally

occuring particles will usually bum and soot particles are generally too small to provide reliable

Doppler signals. Also, the droplets of the mineral oils commonly used to seed non-reacting flows

will evaporate and bum. Among the powders listed in Table 1, aluminium oxide particles [17]

are used in most of the gas flame applications to act as light-scattering centres, have Sauter mean

diameters of around 1 mm, and will follow turbulent frequencies of at least 5 kHz. Titanium

dioxide particles undergo a severe loss in scattering efficiency at the high temperatures

encountered during combustion but are particularly suitable to provide conditional information of

the burned and unburned velocities in premixed flames [18].

Table 1. Typical properties of particles used as light scattering centres in laser velocimetry.

MAlERIAL SIZE SPEC. REFR. MELT. BOIL MOHS

(/lm) GRAV INDEX TEMP. TEMP. HARD.

Aluminum Al203 0.05 3.97 1.76 2300K 3250K 9 Oxide 0.5 Magnesium MgO < 1 3.58 1.74 3125K 3900K 6 Oxide Talc 3Hg04 .Si02.H20 ? 2.8 1.54 ? ? 1 Titanium Ti02 0.22 4.17 2.6 2100K 2800- 5.5-Dioxide 3300K 6.0 Zirconium Zr02 1.5 5.89 2.20 3000K 5300K 6.5 Oxide

Note: The particle diameters are taken from supplier's specifications, while material properties are from the CRC Handbook of Chemistry and Physics, CRC Press. Inc., 1982.

These powders are commercially available in nominally unagglomerated form with diameters

below IIJ.ffi and can readily be dispersed in the dry state with a steady particle flux by a reverse

cyclone device [19]. In complex flows such as those encountered in the combustion chambers of

internal combustion engines, it has been found [20] that zirconium oxide particles are those more

likely to give high laser - Doppler velocimeter data rates and unbiased density measurements,

although this was not found in a comparatively simple open flame [21]. In oil flames, the droplet

387

diameter is measured in regions where it is larger than added solid particles and in solid-fuelled

flames the coal particles are similarly measured. In both oil and coal flames, optical

measurements are impossible where the particle or droplet concentration is high as in the

immediate vicinity of the burner [221.

2.3. EFFICIENCY OF LIGHT TRANSMISSION AND COLLECTION

The application of LDV to confined and luminous com busting flows at high pressure, such

as those typical of combustion chambers in internal combustion engines, may suffer from

additional problems related with window transmissivity and interference with visible emissions.

Ref. [14] provides clear evidence of these problems and is briefly summarized here.

The first of these problems is related with water vapour condensation on the windows due

to the large water vapour content of the combustion products and the high pressure of the engine

flows. This particularly occurs when sapphire windows are used due to their large thermal

conductivity and, as a result, fused silica (quartz) windows should be used although they are

weaker than those of sapphire.

The second problem considered here arises from the increased emission noise observed

when LDV is applied to luminous flames, such as those typical of stoichiometric propane/air

mixtures. Analysis has shown that although the use of current laser line interference filters in the

collecting optics reduces the intensity of the scattered signal by some 50%, the data rate of valid

Doppler signals is significantly improved because of rejections of emission noise.

2.4. THE TYPE OF AVERAGES PERFORMED

Apart from the limitations of the above paragraphs, the application of laser velocimetry to

combusting flows differs little from that in isothermal flows. Possible problems due to particle

thermophoresis appear to be small, but uncertainties in the averages performed should be

considered. This is because it is expected that the measuring volume contains less particles when

it corresponds to high density gas than when it corresponds to low density gas, so that density­

weighted averages of velocity may be measured (Le.,U=pU/p . where U is the measured velocity

component).

Heitor et al. (1985) The kind of averaging process which is performed by a laser

velocimeter by comparing the probability density function of temperature, T, taken from a

continuous temperature record with the distribution of T conditioned by the arrival of valid

velocity signals at the same point in a turbulent premixed flame front making use of a combined

388

4.---------------, b) c)

3

~ 2 0..

o ~~-~-~-~~ 0.0 0.2 0.4 0.6 0.8 1.0 0.0 0.2 0.4 0.6 0.8 1.0 0.0 0.2 0.4 0.6 0.8 1.0

Fig. 3. Density-weighted and unweighted probability density functions of temperature making use of a combined LOy-thermocouple system. a) Unweighted average of continuous temperature record as taken by the thermocouple; b) Unweighted average of conditional temperature measurements by the arrival of a valid. and simultaneous. Doppler signal; and. c) Density-weighted average of the continuous temperature measurements of a).

LDV-thermocouple system. has been investigated; the results, Fig. 3, show that the probabilities

of the higher values of T are reduced in the conditioned measurements, which is due to the

concentration of seeding particles being lower in the hot products than in the cold, dense

reactants. This suggests that the probability density function of the conditioned measurement of

T (Le., Fig. 3b), and thus that of velocity, is close to the density-weighted averaged one (Le., Fig.

3c) and the results of the experiment support this suggestion. It should be noted that these

density-weighted probability functions are biased towards high velocities [23], but for the

turbulence intensities considered here, the errors are less than +4% and -4% for the mean and

variance values respectively [24].

2.5. VELOCITY BIAS DUE TO NON-UNIFORM PARTICLE SEEDING

As a final remark it is important to note that when the fuel and oxidizer are not premixed,

the addition of particles to the air stream alone implies that any measurement of velocity will not

contain the history of the fuel stream. Similarly, if an unconfined flame in still air is to be

analysed. the absence of seeding particles in the entrained air will preclude the quantification of

the history of the entrained flow which is important if the correct flow field is to be measured.

This can lead to uncertainties but can also be put to positive use when conditional sampling is

required [21,25].

For example. the limits of bias effects in flames gases where mixing with ambient air

occurs were established by comparing velocity measurements obtained by seeding with

389

powdered Al203 either a burner flow or the ambient air, with measurements obtained by seeding

only the fonner. The ambient air was seeded by dispersing a large quantity of particles against

the backward face of a plate located upstream of the burner head and the results have shown

differences up to 3xlO-2Ui and 4xlO-3Ui2 (where Ui is the characteristic velocity) for the mean

and variance, respectively. These values are small and may be considered unimportant for the

case discussed here; however, in general, the dispersion of seeding particles in the ambient air

surrounding flames allows one to quantify the details of the mixing which occurs between fuel

and oxidizer, as well as the mass flow of entrained air, and is recommended for most of the

combustion studies.

3. Cases Studied: The Application of LDV to Turbulent Recirculating Flames

The above sections consider only the main limitations of the application of LDV to

combusting flows and it can be concluded that the present state of understanding of the technique

and available instrumentation are considered to be adequate for the measurement of the velocity

characteristics of many complex reacting flows. An important advantage of laser velocimetry is

that it does not disturb a flow in any way other than by requiring an adequate concentration of

particles and optical access. Both requirements can conveniently be met in small-scale flows and,

at the same time, can often penn it the use of forward-scattered light and simple, and therefore

cheap and easy-to-use, instrumentation. It is emphasized that greater benefits are derived at

smaller cost if the flows in small-scale and carefully designed simulations of real flows are

investigated, although spatial resolution and Reynolds number can impose limitations. This is

best explained in tenns of examples, and two have been shown and are described below so as to

make clear the requirements and achievements in the application of LDV to laboratory flows

including those with zones of recirculation, which are of particular interest in engineering [26].

3.1. SWIRLING RECIRCULATING NON-PREMIXED FLAMES

Swirling flames with non-premixed reactants are commonly used in a wide range of

engineering applications, including the burners in industrial furnaces and the combustors of gas­

turbine engines [27]. The combustion process, and the consequent generation of pollutants, in

these flames is partly dependent on the mixing of the fuel and air streams, which in tum is

detennined by the state of the flow turbulence. Improved knowledge of the turbulent structure of

390

NON-REACTING FLOW 2.5

2.0

1.5

1.0

0.5

-0.7 -0.5 _-

o~.~ ___ ~.0~1.5_-_ ~.0-O:2.5 ---- -0.1 0 ___ ___ .5

----------------------0.25 ____ ----------

0.0

0.5

1.0

1.5

( 2.0

2.5

REACTING FLOW

Fig. 4. Measured streamline distributions for the non-reacting and reacting (APR = 27) flows. Re = 49500, after Ref. [21].

4.0 4.0 a)

3.5 REACTING FLOW b) 3.5

REACTING FLOW '" 2 2 ,-.....- 2 2

3.0 contours of k/Uox10

3.0 contours of u'v' / Uox 10

0.1 2.5 2.5 0.5

0 2.0 -0.1 0

2.0 2 0 0 ";:::: ";:::: 4

1.5 1.5 6

1.0 1.0

0.5 0.5 10

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0

0.1 x/Do x/Do

Fig. 5. Iso-contours of density-weighted turbulent kinetic energy. k, and Reynolds shear stresses u V' for the reacting flow of Fig. 4.

391

swirling flames, including the analysis of quantities with direct relevance to turbulence

modelling such as budgets and correlation coefficients of Reynolds stresses, is therefore

desirable. A dual-beam laser velocimeter has been used [28] to achieve these flames. The burner

comprised two coaxial swirling flows of propane gas and air, with swirl numbers of Si = 0.85

and So = 0.77, respectively. The bulk velocities, defined as the ratio between the flow rate and

the cross sectional area, were equal to Uo = 30 mls (Reo = 49500) in the air stream and Ugas = l.8 mls (Rei = 3000) in the gas stream, corresponding to a flame with an air-to-fuel volumetric

ratio of 27.6 and a heat load of about 350 kW. The velocimeter was based on an argon-ion laser

light source at 5l4.5nm (lW nominal output) with sensitivity to the flow direction provided by

light-frequency shifting from acoustico-optic modulation (double Bragg cells) with a resulting

shift of the Doppler signal in the range O-lOMHz. The half-angle between the beams was 4.92°

and the calculated dimensions of the measuring volume at the e-2 intensity were l.528 and

0.132mm. Forward-scattered light was collected and focused into the pinhole aperture (0.300

mm) of a photomultiplier tube with a magnification of 0.74. The band-pass filtered Doppler

signals were processed by a commercial frequency counter (TSI 1980B) interfaced with a l6-bit

microcomputer.

Fig. 4 shows the measured streamline distribution in the vicinity of the burner head for the

flows with and without chemical reaction, which are typical of those observed in highly rotating

flows. They include a central swirl driven recirculation zone surrounded by an annular forward

flow region where the maximum tangential velocities occur with absolute values about 80% of

the annular bulk velocity. For reacting conditions the reverse flow zone is characterized by

uniform and high mean temperatures, as in other recirculating flames. Despite the qUalitative

similarities between the two flows, combustion induces significant quantitative differences: the

mean axial and tangential velocities increase because the density is lowered and the axial and

angular momentum must be conserved; the recirculated mass flow rate decreases from 67% in

the non-reacting flow to 14% in the reacting flow as a result of the reduction in density; the

length of the recirculation zone decreases by 32.5% and its maximum width decreases by about

12%.

Fig. 5 quantifies the main turbulence characteristics of the reacting flow and shows that

the shear layer surrounding the recirculating bubble is a zone of strong generation of turbulence,

which is affected by the curvature of the mean streamlines. Analysis (not shown here for lack of

space) has shown that the turbulent flow is strongly anisotropic with probability-density

distributions suggesting the presence of some flow periodicity in the vicinity of the stagnation

zone. The sign of the shear stress is related to the sign of the shear strain au I ar in accordance

with a turbulent viscosity hypothesis [5], except for a narrow zone in the upstream part of the

392

shear layer adjacent to the reverse flow zone in the reacting flow where the shear strain is close

to zero.

The measurements of the mean and turbulent velocity characteristics reported in Figs. 4

and 5 allow us to estimate the convection and production terms in the transport equations for

turbulent kinetic energy and for Reynolds shear stresses. and help to quantify the mechanisms

involved in the generation of turbulence. The estimates are approximate because of the error in

evaluating the spatial gradients. but the values are sufficiently accurate for the purpose of

establishing the relative importance of the separate terms in the conservation equations. The

results show that the interaction between normal stresses and normal strains influences the

turbulent flow in the vicinity of the rear stagnation point and suggest that extra source terms.

such as those due to the effects of mean pressure field. may be important in the present

unconfined flame. The results also indicate that turbulent diffusion and dissipation are likely to

be important in the balance of turbulent kinetic energy. particularly near the stagnation point and

along the annular swirling jct.

3.2. DISC-STABILIZED. RECIRCULATING PREMIXED FLAME

The use of a laser-Doppler velocimeter has been extended [29] to the analysis of turbulent

heat transfer in a strongly sheared disc-stabilized methane-air flame through its combination with

either laser Rayleigh scattering or digitally-compensated fine-wire thermocouples (see. Ref. [30]

for the analysis of these choices). The laser velocimeter was based on a conventional forward

scattering system using the green light of a 5W Argon-ion gas laser. while the Rayleigh signals

used the blue line of the same laser as shown in Fig. 6.

The procedure for the numeric compensation of the thermocouple signals included the

analysis of the effects of velocity and temperature on the time constant of the thermocouple and

was optimized to allow combined velocity-temperature samples acquired by a custom-built

digital interface with a frequency up to 2000 Hz without deterioration of the thermocouple by

particle accretion. The results were used to discuss the extent to which the "thin flame" model of

burning represents the aerothermochemistry of premixed flames with practical interest [5]. and to

quantify the processes of non-gradient diffusion in a strongly recirculating premixed flame.

The largest random errors incurred in the values of velocity-temperature correlations when

the thermocouples were used are due to the spatial separation of the measurement locations of

temperature and velocity. because the thermocouple junction must lie outside the measuring

volume of the anemometer and it is difficult to reliably place the two measurement locations

closer than about 1 mm. Analysis has shown that the dependence of the velocity/temperature

393

correlation on the spatial displacement between the two measuring zones is weak along

directions characterized by shallow temperature gradients, such as along the streamwise

coordinates. In contrast, the influence of errors in the radial positioning of the thermocouple may

be large and reach 15% of the maximum value of the local time average heat flux, with absolute

magnitudes proportional to the thickness of the reaction zone.

When the LDV was combined with the laser Rayleigh scattering system, post-processing

of the Rayleigh signals which bracketed each valid Doppler realization was required to determine

whether they were contaminated by Mie scattering, and this precluded real time data analysis.

The Rayleigh signal was continuously sampled and stored in a circular buffer, which was

triggered by each valid Doppler signal, and then analysed to select the last Rayleigh signal free

of Mie contamination. In general, the time interval between the velocity and the Rayleigh signal

was below 200 ms, although the maximum combined data rate which could be used was around

1 KHz. This is because the method involved the use of low seeding rates, which were

conveniently controled making use of reverse cyclone seeders and forward scatter LDV optics.

A = 514,5 nm

I' I' I'

~ , I

\ \ INTERFERENCE \1 FILTER • 4BB I'm

POLARIZER

Fig. 6. Schematic diagram of the combined LDV/Rayleigh scattering/thermocouple system used in turbulent recirculating premixed flames, after Ref. [29).

394

12

aJ ~

~ / "-~ /

10

/ \ II \ /' ~ '-. '" ~ y \ p ~~ .-~

~J' ~-n-~ //../ ~

~ ~

~ 0/ oV -u-

02 o 8 1

~ ~ // / I

~ ~~ ----------,

2

6

4

2

o

- 2

- 4

- 6

- 8 rid

------ U I UO'10 ---0--- V I Uo'10 -.-~ U02'100 ----<>-- v-7 U02'100 ---- ~'/UO '200

0,5

............

~ b)

~ r-~ ~

o 1 02 o 3 o 4 o 5 "\" of 0

I

o

-0,5

\ II J -, V

-, ,5 rid

Fig. 7. Radial profile of velocity, a) and heat flux, b) characteristics across the recirculation zone of the disc-stabilized premixed flame, after Ref. [29).

395

Fig. 7 shows sample mean and turbulent velocity characteristics which are typical of baffle

stabilized recirculating flames and exhibit regions of reverse, accelerating and near uniform

velocity. The isotherms, which are not shown here due to lack of space, are highly curved and

reveal a non-planar flame oblique to the oncoming reactants. The reaction region occurs outside

the locus of the mean separation streamline and is curved along its length. This curvature

imposes mean velocity effects on the turbulence field and the result is a strongly sheared flame.

The results have allowed us to estimate the axial and radial turbulent heat fluxes across the

reacting shear layer, as also shown in Fig. 7. These quantities represent the turbulent heat

transfer rate (or, equivalently, the exchange rate of reactants), which is responsible for the

phenomenon of flame stabilization around the recirculation zone and for flame propagation

downstream of this zone. It should be noted that it is the spatial gradients of the heat flux

components, and not the components themselves, that give rise to heat transport. The axial

gradients of u"c" are substantially smaller than the radial gradients of v"c" and hence the

turbulent transport of heat is principally in the radial and not in the axial direction. In general, the

results show that the net turbulent scalar flux occurs along directions in which the mean scalar

gradient is not large and, therefore, calculations using a conventional gradient transport

hypothesis would be inaccurate for any system of co-ordinates. If the present results are

transformed into streamline co-ordinates, the net turbulent scalar transport occurs mainly along

the streamlines as a consequence of the highly oblique flame.

4. Recent developments in laser velocimetry

Many novel arrangements of laser Doppler velocimeters have been proposed since the first

descriptions [31]. So far as optical systems are concerned, two new topics have been considered

in the last years. First, optical fibers have been developed to link the light source to the

transmission optical components and to link the collection optical components to the

photodetector, and this is of particular interest in the analysis of complex configurations such as

those typical of engines [7, 32-34]. Both graded-index [35) and monomode fibers have been used

successfully for transmitting optics, but monomode fibers are preferable due to lower model

dispersion although their efficiency is no higher than 60%. Second, semiconductor materials

have been developed and optimized so that laser diodes can replace gas lasers in the transmitting

optics of current velocimeters and avalanche diodes can replace photomultipliers in conventional

collecting optical systems. As a result, small and flexible laser velocimeters can now be built,

396

LASER DIODE

NETWORK

CONSTANT CURRENT SOURCE

TEMPERATURE CONTROL

PULSE GENERATOR

BEAM SPLITTER

LENS 1

INTEGRATOR

Fig. 8. Schematic diagram of miniature, semiconductor-based laser-Doppler velocimeter, after Ref. [36]

which allow extending the application of the technique to industrially relevant flows [28].

Fig. 8 shows a typical configuration of a miniaturized laser velocimeter based on laser

diodes and avalanche photodiodes [36]. Unlike conventional units, the size of these systems is

not determined by the space occupied by the light source and the detector, but rather by the size

of other optical components, such as the lenses and the beam splitter. As an example, a

semiconductor-based velocimeter is described in Ref.[37], which houses both the focusing and

the receiving optics in an assembly of dimensions 40mmx40mmx230mm. The construction is

modular and allows for easy replacement of parts. As a result, it can easily be forecast that

semiconductor systems will replace current velocimeters for at least some applications, and will

open the possibilities for many new applications including combustion and other industrial

processes, where they may be used as an accurate flow monitoring device. Advantages of diode

lasers over gas lasers are not only their small size, low energy consumption and high reliability,

but also, from the physical point of view. the fact that the spectral sensitivity of Si-photodiodes

shows a maximum at a wavelength of about 830 nm, which is the same wavelength at which

GaAIAs-laser diodes emit maximum power. In this region, the quantum efficiency of Si­

photodiodes can attain 90% and maintain a very low background noise. Therefore, much higher

signal-to-noise ratios can be achieved as compared with the combination of gas lasers and

photomultipliers.

397

The main drawbacks in the use of laser diodes and avalanche photodiodes include the

utilization of optical systems in the invisible part of the spectrum, limited lack of wavelength

stabilization and reduced power. At present, the output power of single element laser diodes is

limited to about 100 mW which, together with current high-sensitivity photodiodes, give rise to

signal-to-noise ratios similar to those obtained with the combination of 200 m W gas lasers and

current photomultipliers; such systems are therefore suitable only for small-scale flames. An

interesting possibility to increase the optical output power of single element laser diodes is to

operate them in the pulsed mode [38,39].

Of the different methods of processing the Doppler signals in electronic form, frequency

counting has found increased use in combusting flows, although its applicability depends on

signal quality, range of frequencies and availability. Where reasonable signal quality can be

achieved, for example in most small scale flows, the use of counters is to be preferred. This near­

standard form of instrumentation can lead to errors due to bias effects [40,41], but these are

comparatively small and certainly smaller than those which can occur due to poor signal-to-noise

ratios. With care, mean values can be measured within 1 % and rms quantities within 3%,

although current improvements in validation arrangements used in burst counters may help to

improve these values.

Where available light is of very low power, or where the particles or droplets are very

small (e.g. <0.2 urn), photon correlation [31] offers a possible alternative, although limited to

signal frequencies (including any shift frequency) of less than around 5 MHz. The combination

of the two-spot optical arrangement [42] and photon correlation can operate in situations where

the light level is in single photons rather than continuous, but this is unlikely to occur in many

practical situations.

Novel signal processing systems include hard-wired frequency-domain processors [43-45],

and the use of Digital Signal Processors [46]. These techniques allow the processing of signals

with extremely low signal-to-noise ratios and, therefore facilitate the extension and common

applicability of laser velocimeters to the study of complex combusting systems.

s. Conclusions

The implementation of laser velocimetry for spatially and temporally-resolved

measurements of velocity characteristics in com busting flows is reviewed in this chapter.

Available knowledge and instrumentation has permitted the use of the technique for the solution

398

of a wide range of combustion problems, although it is argued that greater benefits can be

obtained from small-scale carefully designed simulations of real flows. As an example of this

latter use, results obtained in premixed and non-premixed turbulent recirculating gaseous flames

are briefly discussed.

It is shown that flame-induced beam refraction due to the existence of gradients of

refractive index in the flow is unimportant in small laboratory-scale experiments, but may

decrease the rates of measurement as the size of the combusting flow increases and, ultimately,

may cause the failure of intersection of the transmitting beams. The efficiency of current seeding

particles is discussed and those of alumina appear to be the most appropriate for laboratory

environments, but in high pressure engine applications zirconia may be preferred. Optical access

in confined flows should be provided by fused silica windows, and the use of interference filters

significantly improves the measurement rate in luminous flame zones.

Of course, the measuring volume contains more particles when it corresponds to high­

density cold gas than when it corresponds to low-density hot gas, so that density weighted

averages are measured. In addition, when the fuel and oxidizer are not premixed, the addition of

particles to the air stream alone implies that the results do not contain the history of the fuel

stream; this can lead to uncertainties, but can also bring advantages if conditional averaging is

sought.

In the context of the application of laser velocimetry to com busting flows the paper also

addresses the extension of the technique to the analysis of turbulent heat fluxes, through the

combination of LDV with other techniques for scalar quantities. Again, this has only been

reported in well-controlled laboratory environments, where the combination of laser velocimetry

with Rayleigh scattering may be preferred, at least for premixed flames, although the use of

digitally compensated thennocouples has allowed considerable insight into the analysis of

recirculating flames. For non-premixed flames fonns of Mie scattering have been successfully

combined with laser velocimetry.

Recent developments in laser velocimetry which are expected to improve the analysis of

com busting flows in the near future include the use of fibre optics and the miniaturization of

optical systems making use of semiconductor materials , as well as the implementation of

improved signal processing techniques such as those based on discrete Fourier transfonns.

ACKNOWLEDGEMENTS. The assistance and contribution of our students A.L.N. Moreira and

P. Ferrao for the successful use of laser velocimetry in our laboratory is gratefully acknowledged.

Thanks are also due to Prof. J.H. Whitelaw and Dr. A.M.K.P. Taylor of the Imperial College,

399

London, for discussions held over many years. The manuscript has been typed by Marta Pereira

and Carlos Carvalho and the figures drawn with the help of Mr. Jorge Coelho.

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14. Witze, P.O. and Baritaud, T.A. (1989): "Influence of Combustion on Laser Doppler Velocimetry Signal Quality in a Spark Ignition Engine. In:Instrumentation for Combustion and Flow in Engines", eds. D.F.G. Durao et ai., Kluwer Academic Press, pp. 255-266.

15. Baker, R.I., Hutchinson, P. and Whitelaw, I.H. (1974): "Velocity Measurements in the Recirculation Region of an Industrial Burner Flame by Laser Anemometry with Light Frequency Shifting", Combustion and Flame, 23" pp. 57-71.

16. Buchave, P., George, W.K. and Lumley, I.L. (1979): "The Measurement of Turbulence with the Laser-Doppler Anemometer", Ann. Rev. Fluid Mech., 11, pp. 443-503.

17. Kennedy, I.M. (1982): "Some Aspects of Seeding Flames with Refractory Oxide Particles", Comb. Sci. and Tech., 27, pp. 247-252.

18. Moss, J.B. (1980): "Simultaneous Measurements of Concentration and Velocity in an Open Premixed Turbulent Flame", Combust. Sci. and Tech., 22, pp. 119.

19. Glass, M. and Bilger, M.W. (1978): "The Turbulent Jet Diffusion Flame in a Co-Flowing Stream - Some Velocity Measurements", Comb. Sci. and Tech., 18, pp. 165-177.

20. Witze, P.O. and Baritaud, T.A. (1988): "Particle Seeding for MIE Scattering Measurements in Com busting Flows", In: "Laser Anemometry in Fluid Mechanics - III", eds. Adrian et aI., LAOOAN; pp. 489-502.

400

21. Durao, D.F.G., Heitor, M.V. and Moreira, A.L.N. (1988): "On the Effect of Combustion in Multijet Swirl Stabilized Flames", Proc. 4th IntI. Symp. on Appl. of Laser Anemometry to Fluid Mechanics, Lisbon, 11-14 July; and, Durao, D.F.G., Heitor, M.V. and Moreira, A.L.N. (1991): "Turbulent Transport Processes in Swirling Recirculating Non-Premixed Flames", Proc. 8th Symp. on Turbulent Shear Flows, Munich, September 9-11. Paper 31.5

22. Self, S.A. and Whitelaw, 1.H. (1976): "Laser Velocimeters for Combustion Research. Combust", Sci. and Tech., 13, pp. 171-197.

23. MacLaughlin, D.K. and Tiederman, W.G. (1973): "Biasing Correction for Individual Realization of Laser Anemometer Measurements in Turbulent Flows", Physics of Fluids, 16, pp. 2087-2088.

24. Glass, M. and Kennedy, I.M. (1977): "An Improved Seeding Method for High Temperature Laser Doppler Velocimetry". Comb. and Flame, 29, pp. 333-335.

25. Dibble, R.W., Hartman, V., Schefer, R.W. and Kollman, W. (1987): "Conditional Sampling of Velocity Scalars in Turbulent Flames Using Simultaneous LDV-Raman Scattering", Experiments in Fluids,.5.. pp. 103-113.

26. Heitor, M.V. (1992): "On the Analysis of Turbulent Heat Transfer in Recirculating Flames", Submitted for publication in IntI. J. Heat and Fluid Flow.

27. eitor, M.V. (1989): "Velocity and scalar measurements in model and real gas turbine combustors", In: "Instrumentation for Combustion and flow in Engines". Eds., D.F.G. Durao et and, Kluwer Acad. Publ.. pp. 1-44.

28. Heitor, M.V. (1991): "Advanced sensors for the application of CIME technologies in the process industries", A review. Industrial Methodology, 2, pp. 1-31.

29. Ferrao, P. and Heitor, M.V. (1992 b): "Simultaneous Measurements of Velocity and Scalar Characteristics for the Analysis of Turbulent Heat Transfer in Recirculating Flames", Proc. 6th IntI. Symp. on Appl. of Laser Techniques to Fluid Mechanics, Lisbon 20-23 July 1992.

30. Ferrao, P. and Heitor. M.V. (1992 a): Probe and Optical Techniques for Simultaneous Scalar­Velocity Measurements. In: "Combusting Flow Diagnostics", eds. D.F.G. Durao et al.. Kluwer Academic Publ.

31. Durst, F .• Melling, A. and Whitelaw. J.H. (1981): "Principles and Practice of Laser Doppler Anemometry", Academic Press

32. Obokata. T. Bopp. S. and Tropea. C. (1988): "LDA fiber optic probe with adapter for two point. spectral velocity correlations". JSME. 55 (513-B). pp. 1490 - 1493.

33. Bopp., S. Durst, F. Tropea, C. (1990): "In cylinder velocity measurements with a mobile fiber optic LDA system". SAE paper 900055, Soc. of Automative Eng ..

34. Ikeda, Y., Nakajima. T., Hosokawa, S. and Matsumoto, R. (1990): "A Compact Fiber LDV with a Perforated Beam Expander", Meas. Sci. & Tech., 1 (3) pp. 260.

35. Bopp., S. Tropea, C. and Zhan, L. (1989): "The use of graded-index fibers and fiber optic LDA probes", Rev. Sci. Instr., 60, p. 315.

36. Dopheide, D., Faber, N. Mein, G. and Taraux, L. (1988): "Laser and Avalanche Diodes for Velocity Measurements by Laser Doppler Anemometry", Experiments in Fluids, Q, pp. 289-297.

37. Bopp, S.S., Durst, F., Muller, R. Naqwi, A, Tropea, C. and Welb, H. (1989): "Small Laser­Doppler anemometers using semiconductor lasers and avalanche photodiodes", In. Laser Anemometry in Fluid Mechanics - IV ed, Adrian et ai, Springer Verlag, pp. 315 - 337.

38. Dopheide, D., Pfeifer, H., Faber. N. and Taraux, G. (1989): "The Use of High Frequency Pulsed Laser Diodes in Fringe Type Laser Doppler Anemometry", J. Laser ApI., 1 (4), pp. 40-44.

39. Naqwi, A., Durst, F., Muller, F. and Weller, P. (1991): "LDA and PDA systems based on diode­pumped Neodym: Y AG laser", Proc. Sensor 91 ,N uremberg, May 13 - 16, 1, pp. 15 - 30

40. Durao, D.F.G., Pita, G. Velho, A. Founti, M.A., Laker, 1. and Whitelaw, 1.M. (1984): Some Consequences of bias Effects. In: Laser Anemometry in Fluid Mechanics, eds. Adrian et ai, LADOAN, pp. 381-390.

41. Edwards, R.V. (1987): "Report of the special panel on statistical particle bias problems in laser anemometry", 1. Fluid Engineering, 109, pp. 89-93.

401

42. Selbach, H. and Lewin, A. (1987): "Fibre Optic Flow Sensors Based on the 2 Focus Principle", In: Laser Anemometry in Fluid Mechanics - II", eds. Asanuma et aI., LADOAN, Lisbon, pp. 195-206.

43. Lading, L. and Andersen, K. (1989): "A covariance processor for velocity and size measurements", In "Applications of Laser Anemometry to Fluid Mechanics", Eds. R. Adrian et at Springer-Verlag, Berlin, pp. 454-472.

44. keda, Y. Nakajima, T, Matsumoto, R. and Shimazer, N. (1990): "Performance of prototype hardware of burst digital correlator", Proc. 5th IntI. Symp. on Appl. of Laser Tech. to Fluid Mech., Lisbon, July, 9 - 12. Paper 25.1.

45. Ibrahim, K.R., Werthimer, G.D. and Bachalo, W.O. (1990): "Signal processing considerations for laser Doppler and phase applications", Proc. 5th IntI. Symp. on Appl. of Laser Tech. to Fluid Mech., Lisbon, 9 - 12 July.

46. Kobashi, K., Hishida, K. and Maeda, M. (1990): "Measurement of fuel injection spray flow of I. C. Engine by FFT based phase Doppler anemometer", Proc. 5th IntI. Symp. on Appl. of Laser Tech. to Fluid Mech., Lisbon, 9-12 July - Paper 21.2.

INDUSTRIAL APPLICATIONS OF HOLOGRAPHIC INTERFEROMETRY

N.H. ABRAMSON Royal Institute of Technology Industrial Metrology 10044 Stockholm Sweden

ABSlRACT. This chapter will describe how holographic interferometry can be used for the measurament of dimensions, deformations and vibrations. In particular, the application of holographic interferometry to non-destructive testing will be described. To explain complicated fringe patterns the holo-diagram will be used. Finally the method of light-in-flight recording by holography will be discussed as applied to dimensional measurements.

1. Holographic interferometry for non-destructive testing

Hologram interferometry has during recent years become more and more accepted as a

precision tool for the measurement of mechanical deformations, displacements, dimensions, and

vibrations. By studying the deformation caused by loading, it can be used for holographic

nondestructive testing (HNDT), where it has to compete with ultrasonics and x- rays. It is also

used for the study of stress and strains which have to be calculated from the observed surface

displacements. In the latter case hologram interferometry has to compete with, e.g., photoelastic

analysis using polarizing effects, in comparison to which it has the great advantage that it can be

applied directly to the actual part. In holographic interferometry there is no need to make a model

of a material with specific optical properties that does not behave as the material of the original

part. Photoelastic stress analysis, on the other hand, has the great advantage of producing

information about the entire stress field inside the object, while hologram interferometry only

reveals the resulting surface displacement.

When machine tools are designed, however, the main interest is the stability of the

machine. The stresses and the strains are usually so low that they are only of secondary

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404

importance. The important factor is the static and dynamic deformation of the machine caused by

the production forces. For these studies hologram interferometry is almost without competition.

It produces a 3-D image of the object covered by interference fringes representing lines of

constant displacement. Thus we get directly a map of the motions of the total machine in one

single view. The only other way to produce a similar result would be to measure the machine

point by point using some type of noncontacting sensor connected to a computer with a graphic

display. Even then, however, we would run into trouble because all the measurements would not

be made simultaneously. With these thoughts in mind, we started studying milling machines

using hologram interferometry.

The vertical knee-type milling machine used in the experiment was deformed by a

simulated cutting force produced by a pneumatic membrane which was placed between tool and

workpiece, and the first exposure was made with the machine at rest. Air pressure was applied to

Fig. 1. The cutting force of the milling machine was simulated by static loading. Every fringe represents a displacement of about 0.3 ~m normal to the plane of the photo. Straight fringes represent a tilt around an axis parallel to the fringes. Curved fringes represent deformation. A fixed reference surface is seen on the lower right.

405

the membrane to simulate a low cutting force, and after a few minutes a second exposure was

made. The result is seen in Fig. 1. The top of the machine is the head, while the big part

protruding to the right of the main body is the knee which is supported from the base by a heavy

screw. On top of the knee is the table, which has totally disappeared from the holographic image

because it reached far outside the ellipsoid representing the limited coherence length. On the

floor to the right of the machine is the fixed reference surface against which the motions of the

machine could be defmed. Fig. I displays the motions of the machine in reference to the floor,

each fringe representing a displacement of 0.32 lUll along the line of sight. Thus the motion of

any part of the machine can be calculated by counting the number of fringes that have to be

crossed if we trace a path from the foot of the machine to the studied point. Areas that have not

moved in the direction of sight are fringe free (the foot of the machine, the reference surface).

Areas covered by straight parallel fringes have tilted rigidly without deformation. The tilt is

larger when the fringes are closer. It can be seen that the knee has mainly tilted. The tilt angle

can be calculated by counting the number of fringes across its known width from left to right.

Areas of the machine that are covered by curved fringes have been deformed. Thus it can be seen

that the main body has been deformed by the moment caused by the force acting between tool

and table.

2. Evaluation by the use of the holo-diagram

The holo-diagram was first introduced as a practical device for the making and the

evaluation of holograms. When used for this purpose it consists of the ellipsoids of constant path

length from the focal point A to the focal point B (Fig.2) Each ellipsoid represents a path length

that differs from that of the adjacent ellipsoids by one wavelength. Thus the separation of the

ellipsoids along the x-axis is 0.5 A., while the separation is larger at all other places in the space

around AB. The separation expressed in half-wavelengths is represented by the k-value, which is

constant along arcs of circles (toroids). These k-values are printed alongside the y-axis. To

evaluate a displaced point on an object, Fig. 2 can be used in the following way. Let A represent

the point from which the laser light diverges to illuminate the object (strictly speaking it

represents the pinhole of the spatial filter). B is the point of observation, the center of the lens of

the eye or the camera used for the observation. The hologram plate works like a window with a

memory. Its position does not influence the fringe pattern as long as no optical component is

changed between the two recordings of a double-exposure hologram or between recording and

406

observation of a real-time hologram. To simplify the evaluation, however, in the following let B

be a point at the hologram plate, e.g. its center.

Let the holo-diagram represent a vertical view of the horizontal table surface on which the

holographic set-up is situated. Place the object into the diagram so that its size and its position in

relation to A and B are equivalent to those of the real set-up. If there are n fringes between the

point studied and a fixed point of the object, the displacement (d) is calculated from the

following formula: d = nkA./2, where k is found by following the arc of the circle passing

through the point studied and reading the k-value at the y-axis.

The value of d calculated in this way represents the smallest possible displacement that

could produce n fringes, namely the displacement perpendicular to the ellipses. The true

displacement could be in any random direction: what has been calculated here is only its

projection onto the normal to the ellipses (which is equal to the bisector of the directions of

illumination and observation). To retrieve the complete information concerning all three

components of the displacement in space it is necessary to make three readings using different

observation points.

Each fringe on the reconstructed image of the object is formed when the object point under

study intersects one ellipsoid. Thus, the sensitivity depends both on the separation of the

ellipsoids, the k-value, and the direction of the normal to the ellipsoids. The k-value is given by

k = I along the x-axis outside AB, indicating that the sensitivity is at its highest possible value

here; one fringes is formed for a displacement of 0.5 A. along the x-axis. Between A and B the k­

value is infinite, indicating that there is no sensitivity at all to displacements in any direction.

This statement is true only exactly at the x-axis: as soon as we move away from it the k-value

decreases. At large distances the asymptotic k-value decreases to a value k = 1.

The curvature of the ellipsoids will influence the fringe patterns. A large object that is

close to A and B will have a sensitivity that varies to a great extent over its surface, while a

smaller object far away from A and B will have a sensitivity which is approximately constant. In

the latter case the fringe evaluation is simple: each fringe represents a displacement of 0.5A. in a

direction parallel to the illumination and observation, and there is no sensitivity to displacements

in any other direction. In that case the imaginary interference fringe surfaces are flat.

At closer distances, say less than ten times the separation between A and B, one must

make an allowance in the calculation for the fact that the imaginary interference surfaces are no

longer flat but can be approximated as spheres, their separation still being approximately

constant at 0.5 A. • At still closer distance, say within a sphere centered half-way between A and

B and with radius AB, there exists a near field within which one must take into account that the

k-value varies and that the imaginary interference surfaces consist of ellipsoids.

407

y

k=I.1 k<I.1

---- ----- --

Fig. 2. The holo-diagram used for the evaluation of holograms. A is the point from which the divergent laser beam originates; B is the point of observation behind the hologram plate. Light from A to B via the object at C will not change its path length if C is displaced along an ellipse, while the difference in path lengths to adjacent ellipses is a constant number of wavelengths. The displacement perpendicular to the ellipses that is needed to cause one fringe is kA /2, where k is constant along arcs of circles representing different spacings of the ellipses.

408

c

HL----------------' ~ ~

81

Fig. 3. A diverged picosecond pulse from the laser (A) illuminates the propeller (C). Some of the light is reflected by mirrors Ml and MZ to the hologram plate H. Different intersections Sl Sz and S3 are photographed from B( BZ and B3.

409

In this near field the holo-diagram is of its greatest value. Instead of having to study the

direction of displacement and calculate its projections on to the direction of illumination and the

direction of observation at every point on the object, the k-value is given directly in the holo­

diagram and it is known that the maximal sensitivity is normal to the ellipses. In this way it is

relatively easy to make at least a qualitative evaluation of the fringe pattern formed by a

particular rigid body displacement of the object.

3. Light-in-flight recording for dimensional measurement

Since 1977, we have used the light-in-flight method (LIF) in two ways[2,3J . The shape of

a light sheet (wavefront) has been evaluated when the shape of the illuminated surface was

preknown, or the shape of an object has been measured when the wavefront was preknown. In

the following we will discus the possible industrial uses of the latter method.

In Fig. 3 is shown a typical set up for the study of the shape of a propeller. One of the

resulting sections is seen in Fig. 4. The method is based on the fact that a hologram is recorded

Fig. 4. The reconstructed image of intersection S2 as photographed through the processed hologram plate with the camera in position 82.

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only if the hologram plate is illuminated by object pulse and reference pulse at the same place at

the same time. Thus the reference pulse functions like a light-shutter that with the speed of light

moves over the plate. As the point of observation is moved along the plate during reconstruction

a frameless continuous moving picture is seen of the light intersection moving over the object.

The above described method has been automated (Fig. 5). A computer, combined with an

image processing system, moves a CCD-camera along the hologram by a stepmotor-driven

positioning system in two axis, linear and rotation movement, and stores and processes the

camera images. The images are put together from to form a three-dimensional depth map of the

investigated object [4].

Some of the benefits of the method compared to others are: object motions is of almost no

importance because of the short exposure time, the information becomes immediately accessible

in the hologram but can be evaluated at any instant later, and it is a contactless method which in

the future may be fairly inexpensive. Possible applications are: shortening of the path between

artistic design and CAD/CAM systems; quality control; medical applications in e.g. plastic

surgery, measurement of fast moving objects like rotating turbine blades; and, inspection of long

term object deformations.

There is one unusual twist in our method. The hologram can be used the other way round:

master object can be recorded and the holographic real image reconstructed by an ultrashort pulse

projected back onto a test object [5]. Short pulses are sent out from the position of the hologram

plate in such a sequence that, after being scattered by the object, they all combine into one short

pulse at the point where the laser was during recording. The duration of this scattered light pulse

is a measure of the similarity between test object and master.

For our work we have concentrated on the use of picosecond laser pulses, but experiments

have also been carried out using instead light of shorter coherence. As recording media we have

used ordinary silver halide hologram plates instead of the spectral hole-burning materials

proposed in 1991 [6]. The main reasons are that our material is orders of magnitude more

sensitive to light and does not need cooling by liquid nitrogen.

ACKNOWLEDGEMENT. The author expresses his acknowledgement to Torgny E. Carlsson for

contributing to the third section of this chapter.

References

1. Abramson, N.H., (1981): "The Making and Evaluation of Holograms", Academic Press.

412

2. Abramson, N.H., (1983): "Light-in-flight recording: high-speed holographic motion pictures of ultrafast phenomena", Appl. Opt. 22. 215-232.

3. Valdmanis, J.A. and Abramson, N.H., (1991): "Holographic imaging captures light-in-fIight", Laser Focus 27,11-117.

4. Carlsson, T.E. (1990): "Three dimensional shape measurement by Iight-in-flight single line contouring", Proceedings Holographics 1990 in Nuremberg, 38-43.

5. Abramson, N.H. (1991): "A holographic method and device for obtaining a quantitative likeness measure", SPIE, International Symposium on Optical Science and Engineering, Proceedings Vol. 1539

6. Rebane, A. and Feinberg, J. (1991): "Time-resolved holography", Nature, 351, 378-380.

LASER SAFETY

LASER SAFETY DEVICES FOR HIGH POWER C02 LASER SYSTEMS

P.GAY FIAT Research Center Strada Torino, 50 10043 Orbassano (To), Italy

ABSTRACT. The development of laser technology has raised from the beginning the question of laser safety. The problem is much more important than for normal illumination techniques due to the stronger focussing capability of laser radiation due to its coherence. In Italy at present the following two regulations are in effect: - CEI 76-2 based on EC (1984) proposal with sections one and two (General, and Manufacturing Requirements); - Pubblication CEI 1284 as a User's Guide. The main requirement for laser safety is the presence of a protective enclosure containing all dangerous radiation. This chapter will describe the protective enclosure of a robolaser with a high power laser beam (5 kW). The use of two different concepts to simplify the construction of the enclosure are emphasized: the first considers the lens a part of the enclosure and the second uses an active protective housing. A device to detect the presence of the lens is also described. For the second concept a sensor to detect the impingement of laser beam on the enclosure will be described. This causes the interruption of the laser beam during its penetration of the enclosure. Some data for enclosure penetration time for plastic transparent material of different thicknesses (4 - 10 mm) and power densities (3 - 300 W/cm2) are also presented.

1. Introduction

Since the first development of laser technology in the 1960's, laser safety requirements

have investigated deeply, and several national or international standard regulations have been

produced. On the contrary, very few safety rules and standards deal with conventional light

sources. This is due not only to an increased social attention to safety problems, but also to the

fact that laser exposures form an image point on the retina whereas conventional light sources

form large retinal images at hazardous viewing distances. Another important difference is related

to the high power density levels reached with laser radiation. For the C02 industrial laser, since

the radiation is not transmitted to the retina, only the high power density for cornea and skin bum

damage must be considered.

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s. Martellucci et al. (eds.), Laser Applicationsfor Mechanical Industry, 415-426. © 1993 Kluwer Academic Publishers.

416

The expansion of industrial laser applications particularly for material processing using the

robolaser, has created the need for better diffusion of laser safety knowledge and devices. The

Italian standards on laser safety have been recently issued from CEI in December 1989.

Specifically the "Norma italiana CEI 76-2" is related to Harmonization Document CENELEC

HD 482 SI and Sections 1 and 2 of EC 825 (1984), concernig generalities and manufacturing

requirements. The CEI publication 1282 G is a guide for the utilization of laser systems. The

general safety criteria used in these standards are based on the values of Maximum Permissible

Esposure (MPE) [Esposizione Massima Permessa (EMP) 1 for different conditions of the source

(wavelength, pulse duration),of the nature of the light reaching the target (direct, diffuse or

collected with an optical instrument), and the type of the target (eye or skin). Each laser or laser

system according to its construction has a maximum accessible emission level. The safety

standards define specific emission limits (AEL)[Limite di esposizione accessibile (LEA)] and

divide lasers or laser systems, according to these limits, into Gasses 1,2,3A,3B and 4. For Gass

1 AEL is less than the MPE and there should be no risk of personal injury. Any laser with

continous power output higher than 0.5 W is within Class 4, which is dangerous for direct and

diffuse light vision and is a fire hazard. Therefore all lasers used for materials processing are

Class 4 devices. In industrial environments only Class 1 systems are usually allowed for normal

operation; therefore each industrial system has engineering devices to prevent the exit of

dangerous radiation. It is common practice as a first step to enclose the laser with a protective

housing. The description of the protective housing of a robotic laser system is given in this

chapter. Data concerning penetration times for the barriers used in the protective housing are also

described.

2. Protective housing for a robolaser workstation

The first step when studying the safety of a laser system is the identification of the

Nominal Hazard Zone (NHZ) within which the level of the direct, reflected or scattered radiation

during normal operation exceeds the applicable MPE. Engineering measures must be taken to

limit this zone and to enclose it. An example based on a C02 laser is given below.

The HAT Research Center robolaser system (Fig. 1) basically consists of a 5 kW laser

source, a transfer beam section, a robot with internal beam path, and focalization optics. The

laser, as is the case for any Gass 4 laser, has a unique aperture for the beam output and an

automatic shutter which, when closed, prevents the exit of any dangerous radiation. The NHZ

417

starts from the laser aperture and is directed along the laser axis. The robot and the beam delivery

system are contained within an aluminum duct with a thickness greater than two millimeters. If

the alignment of the laser beam along the duct is guaranteed, then the duct is part of the

enclosure and only the beam output from the robot exit determines the NHZ. To control the

integrity of the duct two approaches are followed: one is based on the monitoring of the external

temperature of the duct, and the second is based on the monitoring of the temperature of a ring

positioned around the beam at the robot exit to control the alignment.

The safe power density MPE for a wavelength of 10.6 !lm and an exposure time higher

than 10 s is 1000 W/m2. It can be shown that this radiation level is reached only at more than

about 100 m from the robot output. This NORD (Nominal Optical Hazard Distance) is

substantially reduced considering the output optics as a part of the system. This approach is

clearly accepted by the eEl laser guide. The NOHD distance can be calculated from the

following formula:

Fig.! Robolaser system of 5kW at the Centro Ricerche Fiat with its protective housing.

418

NOHD=F/D·SQR{4·P/(1t·MPE)} ,

where: F = focal distance; and. P = laser power at robot exit. As described in Fig. 2 the NOHD is

15 m for a laser power of 5 kW. a beam diameter of 32 mm and a focallenth of 7.5" (typical for

laser welding applications). The lens, because it constitutes a movable access panel, requires an

interlock. A device for lens interlok, developed at Fiat Research Center, is shown in Fig. 3. A

LED illuminates the lens and a photodiode monitors the reflected beam indicating the presence

of the lens. Other technical solutions rely on a microswitch on the lens mounting and on the

monitoring of the temperature of a ring positioned around the beam near the focus. With such a

sensor, any loss of focussing due to absence or breakage of the lens can be detected. Due to the

movements of the robot, even with the lens, the hazard zone (NHZ) is too large for typical

industrial applications. Moreover, as the standard requires for a Class I laser an AEL of 0.8 W at

the C02 wavelength, an enclosure is needed. To decrease its dimensions a construction material

able to stop the impinging radiation has to be used. An example of an enclosure is shown in Fig.

4 for a COMAU robolaser. The distance L from the focal point to the wall and the lens focal

length are the two main parameters which determine the choice of material and thickness of the

barriers.

CEI standards define two types of barriers: the passive one maintains its functionality with

nearly no inspection, while the active barrier uses automatic devices or sensors to switch off the

LASER P

"""-"'''-''''-''''iROBOT

P Laser power 5kW

D

F NOHD

D Beam lens input diameter 32 mm F Lens focal length 7.5"

MPE

MPE Maximum Permissible Exposure 1000 W/m2 NOHD Nominal Ocular Hazard Distance 15 m

Fig.2. Schematic of the beam path at the robolaser output optics.

419

Fig.3. View of a focussing system with the lens presence detection device.

Fig.4. Comparison of the dimensions of the protective housing with robolaser beam focus position.

420

lASER BEAM

LED

Transparent Barrier

Ught guide Transmission

Photodiode

Fig.5. Functional design of a sensor for an active barrier.

Electronic circuit

laser when a risk of penetration is detected. This standard requires that the barrier penetration

time be at least of a factor five higher than the time to switch the laser off. after detecting the

Fig.6. Four sensors on the protective housing panel with the electronic detection circuit.

421

impinging beam. The enclosure shown in Fig. I has a transparent active barrier of

polymethylmetacrylate (PMMA) of 5 mm thickness. Each panel has a sensor whose functional

principle is shown in Fig. 5. When the laser beam impinges on the barrier a flame is produced.

Its light propagates by light guiding to a panel position where a photodiode can detect it. A

failsafe function is also provided, with a LED and electronic circuit to monitor the functionality

of the photodiode. The switch off time depends on the input power density and varies from 0.2 to

I s when the power density changes from 800 to 50 W/cm2. In Fig. 6 a photograph of the

sensors and the electronic circuits is shown.

Another technique used to create transparent active barriers is to deposit a linear continuity

conductor along the length of the panel with distances between two parallel conductors of several

millimeters. The destruction of the barrier will cause an interruption in the conductor, which can

switch off the laser.

For safe use of the active barrier it is very important to know penetration times for the

material at the expected power densities. In the next paragraph we report data for transparent

plastics and for metals.

3. Barrier time penetration data

Very little data on penetration times for various barriers are avaliable in the literature, and

usually the experimental conditions are not described with sufficient precision to repeat the

experiments, or to use the results with confidence for the design of a barrier. For these reasons

FIAT Research Center carried out a series of penetration tests on transparent plastics and on

metals. Unfortunately the collected data are very few compared with the large amount of

available barrier materials and experimental parameters.

For each material and thickness the following main experimental parameters have been

chosen: the power density, q ; the beam diameter, D ; the sample total surface area; the

absorption and reflection coefficients of the sample; and, the absorption coefficient of the vapor

produced during laser material interaction.

3.1. TRANSPARENT PLASTIC BARRIERS

Test materials were polymethylmetacrylate (Plexiglass from Bayer) and polycarbonate

(Makrolon from Bayer, and Lexan from GEC), with thicknesses from 2 to 8 mm. The samples

422

were irradiated with a 1.5 kW C02 laser (Model 820 Rofin Sinar) through an appropiate optical

system. The penetration of the barrier was detected with a calorimetric sensor with a sensitivity

range from I m W to 300 W (Photon Control 25 S ) and its signal was used to switch off the laser

and automatically measure the penetration time.

The principal parameters for the penetration time were the material thickness and power

density. Less important was the presence of the flame, which increased the penetration time due

to an absorption of the impinging radiation. The beam dimension and the beam axis inclination

had very little influence. Experiments have shown a linear dependence of penetration time on the

material thickness. In Fig. 7 the specific penetration times sp (s/mm) are presented for three

power densities. The higher one (300 W/cm2) is typical for a raw C02 laser beam, the second

(30 W/cm2) is typical for a distance L of about one meter from the focal point and corresponding

to minimum distance of a barrier for a robolaser housing, and the third value is for a distance of

about 3 m. For these different power densities specific penetration time varies from 1.2 to 112

s/mm for polymethylmetacylate and from 2.4 to 500 s/mm for polycarbonate.

1000

-- polycarbonate

..... :::Ii :::Ii 'U;-

-~-. polfJD.ethyJmetacrylate

~

.... :::Ii 1=

100

z 0

~ ~ 15 Q.

0 ii:

f3 10 Q. UI

10 100 1000

POWER DENSIlY [W/Ct.A2]

Fig. 7 Specific perforation time versus power densities for polymethylmetacrylate and polycarbonate.

423

As indicated in Table 1 the presence of a flame increases the penetration time of about

15% for plexiglass. Polycarbonate penetration times were higher than those of

polymethylmetacrylate by a factor varying from 1.4 to 4.5, depending on the power density

level. The increase of penetration times with power density is well explained by a one

dimensional energy balance. In fact, the input energy density (product of power density and

penetration time) is equal to the sum of material evaporation energy and the energy G needed to

reach the evaporation temperature. For plexiglass the specific evaporation energy is 3000 J/cm3

and Gis 6 J/cm2.

Table 1.

Material

Plexigiass Plexigiass Plexigiass Plexigiass Plexigiass Makrolon Makrolon Lexan Lexan Lexan

Perforation specific times for plastics transparent barriers.

Test conditions

with flame without flame with flame without flame without flame

3.2. METAL BARRIERS

Penetration specific speed s/mm

1.41 1.21

10.27 8.52

112.00 2.41

11.83 2.43

18.15 50000

Some preliminary conclusions can be drawn from experiments carried out so far. We

expect the presentation of these results may draw attention to better design of metallic barriers,

and stimulate other laboratories to produce new data on this subject.

Preliminary tests with a power of 3.5 kW have shown the following basic points. Power

density, in the case of metallic barriers, is no longer a useful parameter as the physical

phenomenon is not one-dimensional. Total power and beam diameter strongly influence

penetration times. Therefore, it is important to be aware that all considerations are limited to this

power level. Absorption and reflection coefficients are also very important parameters. For these

tests we painted the surface of aluminum black to increase the absorption. In fact, with a raw

surface on the aluminum the absorption process, characterized by vapor formation, started after a

variable build-Up time. Test materials were low carbon steel and aluminum (99%) 4 mm thick.

424

400

LASER SP975

200

ABSORBING METAL

PLATE

SAMPLE ABSORBING SAMPLE POSmON 1 CONE POSmON 2

~'""":::::~::~::::::;J.l] PLEXIGlAS PLATE

1000

Fig. 8. Schematic of the experiment to measure the penetration time of metallic samples.

SIN 10

8 f-

6 f-

4 I-

2

[8] 600

500

400

300

2

AVIBAGB PERFORATION TIlOI DIVIDED BY DISPBRTION

80 Fe AI

-- beam diameter ~- material

AVERAGE PERFORATION TIllE

200 L-----~2~----~80~----~F~e------~AI~----~

Fig. 9. Analysis of average penetration time and average penetration time / dispersion ratio for the parameters material (aluminum,iron) and beam diameter (2 and 80 mm).

425

As indicated in the experimetal schematic of Fig. 8, a C02 laser beam of 3500 W was focussed

by a parabolic mirror with a focal length of 400 mm on a metallic sample with transverse

dimensions of 200 mm. Changing the sample position the input beam diameter was changed

from 2 to 80 mm. The penetration time was evaluated' by monitoring the burning of the

plexiglass plate positioned behind the sample. A complete set of experiments with two

parameters (material, diameter) and two levels for each parameter (iron,aluminum; diameters of 2

and 80 mm) were repeated four times. The results for average penetration time, and the ratio of

average penetration time to the dispersion are shown in Fig. 9. All penetration times observed

were between 50 and 1800 s. The analysis of the average penetration time indicates the best

parameters for a metallic barrier are 80 mm for the diameter, and iron for the material. The ratio

between the penetration time and the standard deviation of this experiment show that the best

conditions are 2 mm for the diameter and iron for the material. This meams that even if 80 mm

of diameter has a higher average penetration time, the results are more dispersed than with 2 mm

of diameter.

A final consideration is that with a metallic barrier it is difficult to predict penetration

times without tests using experimental conditions which are close to the barrier design

conditions. Then for robolaser protective housings the divergence of the beam emerging from the

focussing optics also must of course be considered.

4. Conclusions

The expansion of industrial laser applications has created the need for a diffusion of laser

safety standard regulations. In that context the design criteria for the protective housing of a 5

kW robolaser workstation have been described. The European Standards CEI 76-2, and the

Italian Guide CEI 1282 allow us to take into account the beam path after the focussing optics to

reduce the NHZ and to design adeguate barriers.

Active sensors on the barriers and detectors to monitor the presence and integrity of the

lens are useful safety devices. With this approach, the knowledge of penetration times for the

barrier is a critical point. Preliminary experimental results for transparent housings and for

metals (iron and aluminum) have been obtained, and used to establish basic aspects of the design

of the protective housing of laser workstations.

426

References

1. "American national standard for the safe use of lasers" (1980), in ANSI Publication Z 136. I, 1986, Rev. ANSI Z 136.1

2. World Health Organization (WHO) (1982): Environmental Health Criteria nO. 23, "Laser and Optical Radiation", joint publication of United Nations Environmental Program, the International Radiation Protection Association and the World Health Organization, Geneva.

3. British Standards Organisation (1984): "Radiation Safety of Laser Products and Systems", Standard BS4803, London.

4. Deutsche Institut fiir Normung (1984): "Radiation Safety of Laser Products", Standard VDE 0837, Berlin, DINNOE.

5. International Electrotechnical Commission (1984): "Radiation Safety of Laser Products, Equipment Classification, and User's Guide", Document WS 825, IEC, Geneva.

6. "Radiation Safety of Laser Products, Equipment Classification, Requirements and Users Guide" (1984), in International Electrotechnical Commission Publication mC-825, First Edition, CEI Geneva.

7. IRPA, International Non-Ionizing Radiation Committee (1985): "Guidelines for Limits of Human Exposure to Laser Radiation", Health Physics, 49 (5): 341-359.

8. ANSI (1986): "Safe Use of Lasers", Standard Z-136.1 - 1986, American National Standards Institute, Laser Institute of America, Orlando, Florida.

9. ANSI (1988): "Safe Use of Lasers in Health Care Facilities", Standard Z-136.3 - 1988, American National Standards Institute, Laser Institute of America, Orlando, Florida.

10. Norma ltaliana CEI (1989): "Laser equipment and installations Electrical safety" CEI 76-1 "Radiation safety of laser products, equipment classification and requirements" CEI 76-2 (in Italian).

11. ACGIH (1990): "TVL's, Threshold Limit Values and Biological Exposure Indices for 1990-1991", American Conference of Governmental Industrial Hygienists, Cincinnati, Ohio.

INDEX

INDEX

Aberration-free image, 81 Ablation model, 167, 168 Abrasive effect, 155, 157 Absolute interferometry, 369, 378, 379 Absorptance, 92, 93 Absorption coefficient, 117,243,245,423,426 Absorptivity, 35,69,131,132,133,135,136,

137,138, 139, 140, 141, 142, 143, 144, 146, 149, 150, 155, 156, 159, 170, 184, 198

Acoustic emission, 156, 158 mirror, 35, 36,43 mirror beam monitoring, 38 nozzle, 35,43 signals, 37, 38, 40, 158

Acrylates, 383 Activation temperature ofreduction, 145 Adaptive control, 31,42

profilometry, 351, 357, 359 projector, 366, 367

AE,158 AEL,416,417,419 Aerosol phenomena, 243,245 Aerospace applications, 116, 263, 282 ~, 129,130, 131, 133 Airy disc radius, 78 Alarm, 39 Alignment devices, 284 Alloying, 47, 48, 54, 55,69, 71, 72, 73, 84,166,

169,179,193,204,205 Aluminium oxide, 388 Analysis of laser beam, 77, 277, 280 Analytical model, 77,99, 100, 108,374 Annular rods, 23 Aramid fibre reinforced plastics, 118, 123 Articulated robolaser, 236, 265, 266 Assembly costs, 264 Auto-focusing system, 372, 373 Auto-zooming system, 372 Automation, 31, 32,46, 115, 152,235,294 Automative components, 9 Avalance diodes, 380,400 Averages, 384, 385, 390,404 Axisymmetric model, 47, 62

429

Background subtraction, 361 Ballistic particle manufacturing, 297, 298 Base material, 308, 323 BDS, 236, 241, 243, 244, 245, 246, 247, 250,

252,253,256,259,260,263,265 Bead inspection, 228 Beam characteristics, 36, 250

delivery systems, 32, 235, 236, 252, 274, 275, 277,279,280,281,282,305,418

diameter, 19,35,38,99,81,110,119,237, 238,272,278,419,423,425,427

divergence, 78, 80, 272, 290 interference solidification, 303, 307 mode, 96, 272 parameter product, 13, 16, 19,21,24,31 power, 35, 133,262 propagation, 237 quality, 13, 14, 15, 17, 18, 19,21,24,26,28,

31,32,77,78,96,99,208,209,217, 235,236,237,241,243,250,251, 252,255,262,263,305

quality factor, 14 Bending fatigue test, 225 Blur-spot image, 81 Boiling temperature, 109, III Building technique, 296, 298 Buoyancy electromagnetic forces, 57,66 Butt-joint welding, 214, 216, 322

CAD system, 299, 293, 352,411 CAD/CAM integration, 235

CAD/CAM systems, 411 Calorimeter, 36 CAM-programming, 297

technologies, 293 Capacitor, inductance, 35 Carbon fibre reinforced plastics, 115, 121 Cartesian/gantry robot, 236, 256 Cavity optics, 7 CCD camera, 41, 231, 329, 344,368,411

-matrix camera, 361 CFFUP, 115, 118, 121, 122, 131, 134 Chaotic generation pulse, 194, 196,200,208 Chemical laser, 7

430

Chemical laser (cont'd) reaction heat, 50, 51, 52,140

Chemically reactive zone, 50, 51, 52, 53 Chip-formation mechanism, 160 Chopper devices, 37 Chromium-nickel steels, 136, 137 Circular birefringence, 18 Cladding, 16,30,31,41,44,45,48,68,70,71,

166,169,321,323 dilution, 35

Closed loop controller, 34, 43 CNC systems, 280, 288, 290 C02 lasers for cutting, 153, 157,277,288 Coating technique, 315, 321 Co-axial lasers, 108 Coefficient of expansion, 85, 86, 88, 90, 94, 129 Combustion, 367,381,382,385,387,388,389,

392,393,395,401,403 chamber, 7,385,388

Components manufacturing, 322 Composite materials, 115, 133 Computer aid design (CAD), 274

aid manufacturing (CAM), 274 Computer vision, 351 Computerized fringe analysis, 328, 329 Continuity defects, 163, 164 Continuous wave laser, 87, 100,382 Contouring precision, 289, 290 Control systems, 34,41,45, 116,279,290 Control-volume finite-difference method, 205 Convection,48,53,54,57,58,60,61,68,69,

79,169,171,205,387,395 Cooling rate, 57, 60, 63, 64, 65, 70, 80,182,

184, 185, 188, 189 -solidification, 165, 169, 175

Copper mirrors, 94, 247 Corrosion behavior, 264

probe, 322 resistance, 70, 278

Crank-Nicolson's technique, 172 Crash resistance, 264 Crystallization, 163 Cutting, 17,42,66,77, 80, 96, 129, 130, 133,

134, 152, 153, 156, 159, 160, 162, 163, 164, 165, 166, 167,207,208, 209,210,211,212,214,235,239, 249,275,277,278,279,280,281, 283, 321,405

depth, 154, 164,303 parameters, 131 quality, 31 speed, 17, 130, 131,132,133,164,165,152,

153,154,155,156,160,287

CW field laser, 99,109,239,240 CW Nd:YAG lasers, 167 Cylindrical coordinate robot, 236, 256, 257

1) ferrite, 307, 310 Decision making software, 31, 42 Decollimation angle, 79, 92 Deformation limit strain, 224 Demodulation algorithm, 361 Dendritic zone, 308 Density-weighted averages, 384. 390, 391,404 Depth of focus, 7,34,209,272,379

of penetration, 119,214,251,252,272 Design Flexibility, 271, 273 DF (tests), 7,209,215,216 Diagnostic sensors, 322 Dielectrics, 131 Diffractive Optics, 368, 381 Digital Signal Processors, 397 Dimensional measurement, 366, 368, 369, 380,

403,410 Diode-pumped slab, 27 Dissolution of large carbides, 305 Dittus-Boelter Equation, 93 Doppler, 377 Double illumination system, 334 Dual wavelength, III

Edge accuracy, 278 quality, 278

Effect of flow on surface deformation, 62 Electric space charge, 42 Electron beam welding, 3, 131,221 Energy density, 121, 133, 140, 145, 147, 166,

168, 169, 170, 172, 173, 175, 176, 181, 185, 190, 192, 193, 194, 197, 198,204,206,207,279,300,303, 424

Enthalpy method, 168,205 Environmental disturbance, 31, 33 ETCA, 31, 42, 43, 45, 49 European/American TV-standard, 327 Eutectic melting temperature, 163 Evaporation front velocity, 165, 168, 172, 175,

176,177,179,196,197,208 rate, 186, 203

Extraction efficiency, 3

Factor F, 245,247 Far field analysis, 239, 240, 248, 249 Fast Fourier transform, 126 Fatigue behavior, 264 Feeler devices, 7

FFr, 126 Fiber damage, 16

optic beam delivery, 274. 277, 281 optic delivery, 277, 281, 282 optic transmission, 8, 16, 17, 30, 31

Fibre reinforced plastics, 122 Filler wire feeding, 228 Flame, 383, 384,385, 386,387, 388, 389,390,

391,392,394,396,399,402,403, 404

Flexibility, 31,207,208,209,235,257,266, 269,280,281,294,296,351,365, 368

Flexible manufacturing, 17,253,257,351,366 Flow behavior with vaporization, 56,57,61,62 Fluid dynamics, 47 Flux density, 78, 81, 83, 87,89,90,93,94,95,

96,99,110,139,143,168 F/number, 3, 208 Focal depth, 15 Focal position, 35,40,41, 45,49,215,371

spot, 122, 128, 154, 156, 192, 244, 304 Focusing optics, 3,6,77,80,208,209,401

head, 288,289 Forming processes, 295 Fourier demodulation procedure, 398

transform methods, 48, 355, 356 Free surface deformation, 83 Frequency counting, 402

-domain processors, 356, 402 Friction coefficient, 165, 166

conditions, 151 conditions' variation, 166 factor, 89

Frictional coefficient, 90 Fringe analysis, 328, 329, 339,340,341,348

pattern, 327,328, 333, 336, 337,338,339, 340,342,351,353,361,382,407, 408,410

quality, 329, 336 FRP, 116,117, 125 Fume exhaust facilities, 279 Fused deposition manufacturing, 297, 298

silica, 389, 403 Fusion zone, 41. 264 Fuzzy logic, 44

Gantry type robot, 236, 241, 249, 252, 283, 290, 291

Gas turbine cladding, 321 Gaussian beam, 13,31,79,80,81,122,123,

235,237,238,240,241,243 GFRP, lIS, 120, 128, 130

Glass fiber reinforced plastics, 115, 123 transition temperature, 129 glazing, 166, 169, 184, 190

Graded index, 367, 395 Gradient index fiber, 29 Grain refinement, 305, Grating projection, 353, 354

Half far field divergence, 13 ~ening,35,40,42,47,48, 156, 158, 162,

166,287,298,300,314,321,322 ~essHv,225

results, 309

431

Heat affected zone (HAZ), 117, 118, 126, 132, 152,166,182,189,205,264,279, 281,308,311

capacity, 117, 133 conduction, 48, 49, 50, 83, 130, 143, 190,

200,205,215 distortion, 273 momentum and mass transfer, 9, 47 processes, 144, 165,166,169,175,176,181,

184,194,196,197,198,200,205, 206,208

properties of metals, 131 Heating gas, 37

mirrors, 37 wire,37

Heavy duty bearing, 307 He-Cd laser, 315, 316,383 He-Ne laser, 154,283,317,347,356 Hermite modes, 237 Herz-Knudsen law, 169, 171 High power laser, 3, 4,11,13,77,78,82,96,

132,235,236,415 weaponry laser, 78

-speed pyrometry, 166,208 -speed recording, 166

HNDT,403 Holo-diagram, 403, 406, 410 Hologram interferometry, 327, 328, 329, 332,

342,344,351,354,355,356,403 plate, 403

Holographic interference solidification, 298 interferometry, 328, 353, 403, 406 optics, 367

Holography, 284, 297, 335,403

Image normalization, 361 processing, 43,328,330,340,344,361

system, 129, 411 techniques, 353, 354

Inconel 718 alloy, 159, 165

432

Incremental interferometry, 368 Inductive sensors, 41 Industrial vibration testing, 351 In-field repairing, 324 Infrared radiation, 42 In-process control, 43, 44,367

diagnostics, 33 monitoring, 36 sensors, 35, 39,45

In-production laser cutting systems, 277, 283 laser welded assemblies, 263, 268, 271, 275 laser welded blanks, 263, 265, 267

In-situ processing, 316 Inspection station, 228 Integrated articulated robot, 260

optical density, 129 optics, 380 sheet blanks, 231

"Intelligent" processing, 43,44, 45 Interference filters, 389, 403 Interferometry,45, 166,327,328,331,335,338,

343,352,355,353,365,368,373, 374,376,377,379,403

sensitivity, 337, 340 International Organisation for Standardization

(IOS),13 IOD,129 Isothermal oxidation, 139

Ioint tolerance, 272

KCllenses,154 Kerf surfaces, 117, 118, 126, 129, 130 Keyhole effects, 5, 8

monitoring, 42 welding, 5,8,264,271,272

Knudsen layer, 69, 169 "Kvant-16" Nd:YAG laser, 184

Laguerre modes, 237 LAM, 151, 152, 153, 155, 156, 157, 158, 159,

161, 162, 163, 165, 166 Lambertian radiator, 97 Laminated object manufacturing, 302 Laser annealing, 161

assisted machining, 151, 152 Laser beam analyser, 35, 36, 37

guidance, 32 litography, 315

Laser cutting, 42, 45, 115, 120, 129, 132,277, 279,281,283,288,321

processes, 132, 277, 278, 279, 280 deposition process, 323

Laser diodes, 27, 255, 374,400,401,402 Doppler velocimetry, 334, 381, 400 efficiency, 14 hardening, 161,322 heating, 50, 132, 133, 135, 136, 138, 140,

142, 149, 153, 155, 156, 160, 161 dynamics, 138

Laser-induced melting and vaporization, 79 reduction, 144

Laser interferometry, 35, 166,376,378 machining, 115 material interaction, 115,279,422 processing, 36,47,54,207,217,271,283,

321,367,380 plasma, 151, 165 power, 5,36,39,44,45,78,79,83, 117,123,

130,162,165,208,210,211,237, 267,272,287,318,418

processing, 3, 31, 34,47,48,54,72,78,83, 272,276,288,315,323,365

radiation, 31,38,138,139,161, 162,367, 415,416

absorption, 156 remelting, 161 robotics, 33, 235 safety, 415, 416 scanning, 51, 52, 300 surface melting (LSM), 47,48,55,58,69,

305,313 thermochemistry, 144 treatment parameters, 305, 308 triangulator, 369 velocimetry, 381, 382, 383, 384, 385, 387,

392,394,396,399,400,401,403 welded blanks, 219, 220, 231, 233, 263, 264,

265,266,267,268 welding, 3, 5, 6, 7,40,214,219,221,222,

271,272,273,275,276,321,322, 323,418

line, 233 systems, 41, 267, 271, 274

workstations, 33,417,426 writing, 315, 316

Latent heat of melting, 171, 180 LBA, 35, 36, 37 LCD characterization, 357,359,364,365,366,

367 LCVD, 52, 53, 74, 75 LDV, 336, 381, 382, 386, 387, 389, 391, 392,

393,397,404 Ledeburite, 10 LlF,41O Light-in-flight recording, 403, 410

Liquid bulge test, 223 chrystal device, 359

display, 351,364 projection, 351

metal + liquid oxide system, 147 polymerization, 297

Location of the beam, 41 LOM,302 LUCAS, 41 Luminous flames, 389,403

Machining of materials, 99, 115, 116, 130, 132 MAD,130 Magnetic domain refinement, 323 Magnification, 4, 5, 264, 330, 337, 351, 394 Marangoni number, 62, 63 Marechal criterion, 82, 249 Mass transport, 48, 71, 72, 205 Material cost, 219, 220, 265 Material processing, 3,11,15,16,17,28,31,47,

54,79,167,208,211,217,271,273, 283,287,321,367,380,416

reflectivity, 272 removal processes, 79, 117, 163, 179,293,

295,306 savings, 265

Maxim urn permissible exposure, 416 Mean absolute deviation, 130 Mechanical treatment, 151, 152, 153, 162 Medical applications, 411 Melt thickness, 165, 168, 169, 172, 173, 175,

179, 181, 182, 186, 191, 192, 196, 201. 202, 203, 206

Melting boundary velocity, 167, 172, 196, 198 -solidification problem, 165, 204, 206, 208 velocity, 168, 172, 175,177,179,180,181,

197,202,203,208 Metal barriers, 424, 425 Metals, 8,12,117,131,132,133,138,139,144,

151, 156, 158, 165, 168, 173,215, 219,223,282,302,304,421

Michelson interferometer, 337 Microcracks, 151, 157, 159,160,162,166,353 Microlitography, 316, 318 Microphone, 41,43 Microprocessing, 17 Microstructure, 48, 55, 58, 67, 84, 264, 272, 305,

306,307,308,309,310 Miniature optics, 367, 379 Mirror bowing, 83, 85, 94

ripple, 83, 85, 94 Modulus of elasticity, 86 Moire topogmphy, 354

Moire (cont'd) fringe pathlength, 355

Molybdenum rod, 37 Monomode fibers, 400 Motion systems, 279, 280 Mott formula, 139 11PE,416,417,418 Multi-gauge blank, 269,270 Multikilowatt lasers, 3,7,208,276,321 Multirod system, 24

433

Nd-YAG laser, 11,12, 17,21,31,166,184,204, 255,271,274,276,277,281,282, 283,284,323

Near field analysis, 250, 252 NHZ,417,419,425 Nickel superalloy, 4 NOHD,417,418

distance, 417 Noise emission sensor, 322 Nominal hazard zone, 417 Non-adaptive profilometer, 363 Non-contact process, 273, 365 Non-destructive testing, 338, 403 Non-thermal distortion, 245 Nuclear fuel containers, 322 Numerical aperture, 16,78,81

simulation, 165, 167, 168, 194, 198

Object pulse, 411 One-dimensional transient model

for cylindrical bodies. 50 for flat semi-infinite body, 49 for laser cladding, 73

On-line control, 17 inspection, 351

Open loop controller, 34 Optical components, 78, 79, 80, 81, 83, 84, 85,

87,90,92,95,96,97,238,239,240, 246,261,365,367,379,380,400, 401,406

distortion, 77, 90 fibers, 255, 301, 317, 323, 335, 367,400 fibre system, 323 measuring tool, 327, 352 metrology, 366, 378, 380 pathlength, 245, 331, 335, 340, 386 profilometry, 351, 353, 355, 367 scattering, 37 sensors, 40, 367

Overlap welding, 213, 214 Oxidation, 74, 131, 139, 140, 142, 143, 145, 146,

156,170

434

Oxide-metal film system, 145 Oxide + metal system, 140, 142 Oxide permittivity, 141

Parabolic approximation, 142 Parallel process, 316 Particle seeding, 381, 385, 387 Particles, 47, 73,156,235,293,306,308,309,

311,349,382,385,387,388,389, 402,403,404

Parts consolidation, 264, 265,266 Penetration ability, 131 Penetration time, 415 Perforated mirror, 35,37 Perturbation model, 61, 62, 64, 65, 68, 84 Phase changes, 151, 160, 161, 166, 168, 171

demodulation technique, 357 effect, 386 -modulation interferometer, 374, 376 -shift method, 355, 358

Photo elastic analysis, 403 stress analysis, 404

Photon correlation, 402 drag in Ge, 37

Piezoelectric detector, 39,43 sensor, 35

Pixel segmentation, 365, 367 Plasma, 6, 36, 67, 131, 152, 165, 166, 170, 173

breakdown, 5, 6 charge sensor, 42, 43 formation, 4,7,8, 9,48 generator, 156

Plastic surgery, 411 Plasticization, 151, 157, 159 Poisson's ratio, 86, 91, 94 Polarization orientation, 207, 211, 214 Polarizing effects, 248, 403 Post assembly cutting, 281 Powder feed rate, 35,45 Precision drilling, 99 Premixed flame, 396,404 Pressure bending, 88, 246 Process analysis, 251

efficiency, 208, 209, 216, 217 flexibility, 208, 213, 217

Processing velocity, 6, 17, 213 Product design, 263,265,270, 272, 273. 276 Production applications, 3, 9, 10,268,271 Productivity, 31,130,132,151,166,221 Profilometer, 352, 358,359,367 Profilometry, 353, 365, 368 Propagation parameter M, 241 Protective enclosures, 280,415

Pulse interruption, 165, 175, 184, 188, 190,194 laser action, 165, 175, 196,200,202,208 laser alloying, 204, 205

Pupil function, 78 Pyroelectric (detectors), 37

Quality control, 351, 365, 366,411 Quasi-steady state models, 58, 60

Rapid prototyping, 293, 294, 295 solidification, 47,69

Rayleigh length, 15,208 quarter wave limit, 81, 82 range, 238 scattering, 393, 398

Real-domain demodulation procedure, 355 Recirculating flames, 392, 394, 399, 403, 404 Reference pulse, 411 Refraction bias, 385, 387 Refractive index, 16, 18,30,84,146,331,367,

385,386,387,403 Refractory metals, 144 Remote welding, 323 Repetitive pulse laser, 88,99, 103 Resist exposure, 316 Resistance spotwelding, 261, 271,272,273 Retained austenite, 305, 306, 308, 309, 310, 313 Reverse machining, 322

thermal wave transform, 99,100, 105, 109 Reynold's number, 60, 61, 90, 93, 392 RMS surface roughness, 85, 92, 249

wavefront distortion, 82, 245, 249 wavefront error, 82, 83

Robolaser, 235, 236, 262, 265, 266, 415, 416, 417,419,422,425

Robot-mounted lasers, 255 Robot tracking, 374 Robotic beam delivery system, 32

cutting systems, 281 laser cutting systems, 277, 283

welding, 274, 275,276 welding systems, 273

Robotics, 211, 235, 280, 281, 317,318 Robotic systems, 235, 351, 376 Robotization, 230, 252 Robotized laser system, 235 Rolling wear, 305 Roughness, 85,92, 151, 157, 160, 165 RP,93, 94, 294, 295, 296, 297, 298,299, 301, Safety standards, 416, 425 Sag radius, 85 Sapphire windows, 389 Scanning systems, 353

SCARA,373 Scatterer centers, 385, 387 Scattering coefficient, 243 Schlieren effect, 386, 387 Seam following, 41

location, 32 Secondary hardening, 306, 313 "See through" rnirror, 42 Selective laser sintering, 296, 298, 301, 302 Self regulating systern, 33

teaching, 45 Serniconductor rnaterials, 400, 404

photo-detectors, 367 Serniconductors, 131 Sensing options sensor, 36 Serial process, 316 Shadow casting, 365, 368 Shape rnelting, 297, 298 Shears, 151, 156, 158,162 Shield gas, 272 Shockwaves, 43 Shroud gas, 43 Skids,40 Slab-lasers, 24 Srnart laser cutting systern, 45 Solid state lasers, 3, 11, 12, 13, 18,204,282,283 Solidification velocity, 180, 181, 188,201,202 Solute redistribution, 54, 60, 63, 64, 80 Space application, 344 Spark discharge rnonitoring, 43 Specific heat of evaporation, 171 Speckle decorrelation, 336, 337, 350

patterns, 339,349,350 shear systern, 337

Speed effects, 6 Spherical coordinate robots, 256 Spot size, 6, 31,32, 192,240,242,248,249,

278,279,301,371 Stability of laser irradiation, 202 Stamping cost, 219 Stefan type boundary condition, 169 Step index fiber, 29, 31 Stereolytography, 299,301,302 Stereovision, 354 Strain analysis, 263, 267 Strehlcriteria,81,82,95,96

ratio,81,82,94,96,242 Strength of joint, 272

ratio, 222 Stroboscopic recording, 333

speckle techniques, 332 Structural analysis, 263, 264

strength, 273

Supersonic nozzles, 210 Surface absorption, 35, 110, 144

cladding, 322 hardening quality, 35,47,48 oxidation, 131, 132, 139 reduction, 132, 144

ofrnetals,132,144 ternperature, 52, 63, 65, 67,100, 105, 131,

147, 149, 150, 165, 168, 172, 175, 178,180,182,184,189,196,201

tension driven flow, 57 Surgical (C02) laser, 284 Swirling flarnes, 393, 394

recirculating, 393 Syrnmetrical plane-plane resonator, 35

Table position, 35 Tailored blanks263, 264,266 Ternperature diffusivity, 198

field, 60, 62, 64, 79

435

ofrnelting, 61, 109, 158, 167, 171, 182, 189 of rnolten pool, 63, 64

Ternpering, 305, 308, 310, 311, 312, 313 treatments, 310

Thermal conductivity, 58, 70, 88, 90, 100, 104, 117,118,132,133,134,135,136, 137,138,140,150,171,279,389

cutting processes, 277 cycle, 53, 165 degradation, 117, 118, 126 diffusivity, 5, 61, 70, 101, 103, 106, 140, 144,

153, 171, 175 distortion, 5, 10,24,28,245,247,257 effects, 22 expansion, 86, 129,245 rnodel, 117, 130 polyrner, 297 pressure, 89 stresses, 39, 86, 129, 165

Thermally induced wavefront error, 82, 87, 247 Thermocouples, 35, 157,391,396,397,404 Thermo-plastic deformation, 160 Thickness ratio, 224, 225 3D cornponents, 235

printer, 293, 296 topography, 354,372

Three-dirnensional heat-transfer rnodel, 53, 80 rnodel for semi-infmite plate, 50

transient rnodel for finite slabs, 50 for laser chernical vapor deposition, 74

Threshold of rnelting, 178 TIG, 221, 322

welding, 41

436

Time-of-flight technique, 367, 369, 370 Titanium dioxide, 388 Tool durability, 162

steels, 305 Transient model, 48, 50, 57, 69, 70, 74 Transition zone, 308 Transmitting optics, 8, 400 Traverse speed, 35, 40, 45 Triangulation, 365, 369, 372, 373, 374

technique, 353, 356,368,369,372, 376,379 Turbulent heat fluxes, 385, 399, 404

recirculating flames, 392,403 TV-holography, 334, 338 2D laser cutting, 283

printer, 293 Two-dimensional model, 58, 60, 204 Two-layer system, 145, 146 Two-level backward Euler scheme, 205 Two-wave interferometer, 329

Ultimate resistance, 158 Ultraviolet, 42, 299 Unstable oscillator, 4

Vaporization,S, 48, 66, 75,80,109,117,131 temperature, 6, 109, 117, 118

Variable density flow, 383 Velocity field, 58,60,62,63,64,79,80 VGA-PAL converter, 364 Vibration analysis, 339, 344, 347

stability, 35

Video speckle interferometry, 327, 339, 349 system, 328

Visible emissions, 381, 385, 389

Wagner formula, 139 Waistradius, 13 Wave frequency, 40 Wavefront aberrations, 80, 81 Weight reduction, 31, 263 Weld design, 272

faults, 41 integrity, 267 pool, 7,42,47, 54,65

Welding,S, 6, 7, 8,10,12,41,44,53,77,96, 166,207,209,210,211,214,215, 216,235,238,261,271,272,274, 321,323,418

cost, 219 of aluminium, 215, 216 of steels plates, 213 quality, 35 speeds, 225, 226, 249, 267, 272 test machine, 222 torch, 228

Window transmissivity, 381, 385,389 Workpiece movements, 249, 250

X-ray diffractograms, 310 X-rays, 403

Young's modulus, 91, 94