Ultrashort-Pulse Laser Processing ofThick Metal Layers on a Ceramic

Substrate

Master Thesisfor

Master of Science in Lasers and Photonics

Department of Electrical Engineering and Information Technology

Chair of Applied Laser Technology

Ruhr-University Bochum

Igor Sakaev

Supervisor: Jun.-Prof. Dr. rer. nat. Evgeny Gurevich

Bochum 2015

Abstract

Ultrashort-Pulse Laser Processing of Thick Metal Layerson a Ceramic Substrate

The work investigates the feasibility of ultrashort-pulse laser processing for micromachining ofmetallic coatings, applied to the ceramic substrate by a thermal spraying method. The limitations ofthe thermal spraying technology do not allow to produce a coating layer with thickness less thenseveral 10 μm. Therefore, the term thick metal layer is used to distinguish the investigated processfrom the ablation of thin films deposited on a ceramic substrate. For more precise definition, theprocessed layer may be considered thick, when its thickness is comparable with the laser beam spotsize. The purpose of the investigation was to prove the possibility of fabrication of the embeddedthermocouple sensor by ultrashort-pulse laser processing in conjunction with thermal spraying,substituting the conventional thin film sensor technology based on vacuum deposition techniques. The experimental laser processing of test samples, coated using thermal spraying methods, iscarried out using ultra-fast fiber amplified laser system Tangerine. The beam guidance and focusingis performed by SCANcube 7 scanning galvanometer. The modern state of the theory of metalsprocessing using ultrashrot laser impulses is reviewed and applied to the particular technologicaltask, in order to establish the optimal laser radiation and process parameters. As a prerequisite for the processing task, the ablation threshold fluence of the laser pulse for thethermal spraying coated Nickel alloys, used to form a thermal junction, is estimated experimentally.The obtained estimation is very rough and has an order of magnitude of 0.1 J/cm2, similar to theablation threshold of a bulk Ni and Nickel super-alloys. Then, the experimental microfabrication ofnarrow bimetallic stripes on a ceramic substrate, resembling the features of the thermocouple sensorstructure, was performed. From the comparison of the microfabrication results at different focusinggeometries, the major effect influencing the morphology of the produced structure was found to benon-uniform ablation rate and crater/channel growth. This effect is typical for the micromachiningof structures with a high aspect ratio and is well described for the case of deep multi-pulsedrilling/cutting, where it leads to the formation of a tapering crater/channel. The main reasons ofthis effect are the Gaussian spatial distribution of the laser fluence and the oblique incidence on thecrater/channel sidewalls. The resulting morphology of the produced micro-stripe is qualitativelydescribed as a trapezoidal shape, with its characteristic dimensions defined by the spatial laserfluence distribution. Another important effect affecting the morphology of the structure is assumedto be the low fluence finishing, which is responsible for the high quality of the stripes sidewallsobtained in the experimental microfabraction. Finally, an attempt was made to fabricate a prototype of the fully functional thermal junction,consisting of the contact pads and the interconnecting bimetallic micro-stripe. The main obstacle,preventing a successful reproducible fabrication of the thermocouple resembling structure, wasfound to be the appearance of micro-cracks in the interconnecting stripes at the vicinity of theabrupt variations in the substrate thickness. Regardless, one of the fabricated thermal junctions wasfunctional and produced thermal voltage of 14 mV. Although, the possibility of the proposedmethod of the embedded thermocouple sensors microfabriction was principally proven, furtherinvestigations are required to improve the process, in order to achieve a true industrial feasibility.

Table of Contents

1. Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1

1.1. Laser direct patterning of metal films deposited on a ceramic substrate . . . . . 1 1.2. Motivation: microfabrication of embedded thermocouple sensors . . . . . . . . . 2

1.2.1. Thin film thermocouple technology . . . . . . . . . . . . . . . . . . . . . . . . . . . 2 1.2.2. Thermal spraying technology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3 1.2.3. Fabrication of embedded thermocouple by thermal spraying

method in conjunction with ultra-fast laser micromachining . . . . . . . . . . . . . 3

2. Ultrashort-Pulse Laser Processing of Metals . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6

2.1. Basics of ultra-fast laser material interactions . . . . . . . . . . . . . . . . . . . . . . . . . 6 2.1.1. Minimization of the heat affected region as a major advantage

of the ultra-short pulse laser processing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7 2.1.2. Threshold fluence and its dependence on the pulse width . . . . . . . . . . 8 2.1.3. Ablation depth per pulse. Low and high fluence regimes . . . . . . . . . . . 9 2.1.4. Thermodynamic trajectories of ultra-fast phase transitions.

Spallation and phase explosion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14 2.1.5. Implications on laser processing strategies . . . . . . . . . . . . . . . . . . . . . 17

2.2. Influence of the beam shape and other optical parameters . . . . . . . . . . . . . . 18 2.2.1. Ablation in a Gaussian beam . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 18 2.2.2. Non-uniform ablation rate and channel/crater formation . . . . . . . . . . 19 2.2.3. The effect of the beam polarization . . . . . . . . . . . . . . . . . . . . . . . . . . . 20

2.3. Influence of the ambient gas atmosphere . . . . . . . . . . . . . . . . . . . . . . . . . . . . 21 2.3.1. Beam distortion due to the non-linear effects and optical

breakdown . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 21 2.3.2. Contamination of the ambient gas by the ablated particles and

plasma shielding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 22 2.4. Influence of the pulse repetition rate and pulse-to-pulse interactions . . . . . . 25

2.4.1. Heat accumulation due to residual thermal effects . . . . . . . . . . . . . . . 25 2.4.2. Incubation effect and damage accumulation . . . . . . . . . . . . . . . . . . . . 28

2.5. Laser ablation cutting . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 29 2.5.1. Cutting velocity and pulse overlap . . . . . . . . . . . . . . . . . . . . . . . . . . . 29 2.5.2. Scaling the processing parameters . . . . . . . . . . . . . . . . . . . . . . . . . . . . 30

3. Experimental Setup and Test Samples Characterization . . . . . . . . . . . . . . . . . . 32

3.1. Laser setup for experimental material processing . . . . . . . . . . . . . . . . . . . . . 32 3.1.1. Ultra-fast fiber amplified laser system . . . . . . . . . . . . . . . . . . . . . . . . 33 3.1.2. Scanning galvanometer for laser material processing . . . . . . . . . . . . . 34

3.2. Test samples characterization: ablation threshold of the Nickel alloy thermal-sprayed coating . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 34

3.2.1. Logarithmic fluence plots with the ablation crater diameter measured using optical microscope . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 35

3.2.2. Verification of the ablation crater diameter using electron microscope imaging . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 38

3.2.3. Logarithmic fluence plot for a tighter focusing geometry and improved beam alignment . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 40

3.2.4. Validation of the estimated ablation threshold value . . . . . . . . . . . . . . 41

4. Experimental Fabrication of Bimetallic Micro-stripes on a Ceramic Substrate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 43

4.1. Microfabrication by a direct laser ablation: process design and parameters . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 43 4.1.1. Beam trajectory patterns . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 43 4.1.2. Laser radiation and processing parameters . . . . . . . . . . . . . . . . . . . . . 44

4.2. Morphological analysis of the experimental micromachining results at different focusing geometries . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 47 4.2.1. Experimental microfabrication results at poor optical alignment

conditions with a large beam spot size . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 47 4.2.2. Experimental microfabrication results at improved optical

alignment conditions with a beam tightly focused on a sample surface . . . . 50 4.2.3. General patterns in the morphology of the fabricated metallic

stripes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 52 4.3. Measurable criteria for a metallic coating removal . . . . . . . . . . . . . . . . . . . . 53 4.4. Fabrication of a structure resembling a full functioning thermocouple

sensor . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 55 4.5. Conclusions and prospect . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 58

References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 61

Chapter 1

Introduction

1.1 Laser direct patterning of metal films deposited on a ceramic substrate

Direct laser ablation is occasionally used as an alternative to photolithography/etching productiontechniques, while processing a film-substrate system. It allows in general more rapid andstraightforward processing which does not require masking, complex multiple step processes, andclean room environment. Speaking more generally, the process of material removal by ablationinduced by a scanning laser beam is referred to as an ablative direct writing or laser ablation cutting.It is often used in microfabrication where high precision is demanded. Considering the case of thefilm-substrate system processing, it is often required to produce the desired pattern on the film withminimal influence on the substrate. In contrast to photolithography/etching, in most cases thesubstrate can not serve as stop-layer, non affected by film removal process (although if the ablationthreshold of the substrate is much higher then that of the film, this situation is possible [1]).Therefore, minimizing the substrate damage becomes an issue. Many research efforts are investedin applications of laser ablation to the thin film technology for patterning of the thin films, havingthickness of several 100 nm or even less. However, there is no particular restriction to applyingdirect laser ablation patterning method to films with thickness of several 10 µm (sometimes referredas thick films), as in the present work. Although in this case most of the effects of film-like ablation,which are extensively studied for thin film applications, do not play any significant role. As a resultthe ablation has a bulk-like behaviour more typical to high precision microfabrication of thecomponents from the bulk material. From the requirements of minimal substrate damage and high precision of the pattern structure theneed of short and ultra-short laser impulses arises. The use of ultra-fast laser systems with pulseduration below characteristic electron-phonon relaxation time allows the highest possible spatialoverlap between the heat affected region and the ablated volume. Consequently, only negligibleamount of melt is formed in the ablation process and the width of heat affected zone (HAZ), whichremains after material ablation, is minimized. Owning to this advantage, these laser systems areincreasingly used in thin film patterning and in high precision quality demanding micromachiningapplications. As mentioned before, the patterning of the thick films has more similarities with ultra-fast laser micromachining, considering the physical nature of the processes involved. But theconcept of obtaining a pattern on film-substrate system by direct ablation of the film is identical tothe thin film patterning. The well known advantages of laser material processing along with highprecision, resolution and quality of the obtained pattern structure distinguish direct laser ablationfrom conventional thick film technologies. The term “micromachining” is often used to describe applications where the resolution down to themicrometer or several tens of micrometers is achieved [2]. Due to Gaussian shape of spatial energydistribution and uniform ablation threshold of most materials, resolutions of sub-micrometer scaleare also possible. However, the present work deals with the patterns with tens of micrometerscharacteristic feature size. Metal film of several tens of micrometers thickness deposited on aceramic substrate is subjected to direct laser ablation in order to obtain pattern of lines several tensof micrometers wide, isolated from one another by a ceramic gap. The aspect ratio of the desiredstructure is therefore close to 1. The use of ultra-fast fiber amplified laser system with pulseduration around 300 fs allows minimal heat affected zone, potentially leading to high structurequality and contour accuracy. Nevertheless, it is achieved not only by a mere reduction of pulseduration, but also on the basis of proper adjustment of a wide range of process and laser parameters.

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1.2 Motivation: microfabrication of embedded thermocouple sensors

1.2.1 Thin film thermocouple technology

The thin-film thermocouple (TFTC) technology was originally developed for application tosuperalloys used in jet aircraft engines for temperature measurements up to 1000°C [3]. Thistechnology was advanced through several NASA contracts and grants in order to meet the urgentneeds in aeronautic and aerospace research. Sensors in a thin-film form provide a minimallyintrusive means of measuring surface parameters, such as temperature, in hostile environments.They are used in engine systems in order to evaluate advanced materials and components andprovide experimental verifications of computational models. Unlike more conventional wire or foilsensors, thin-film sensors do not necessitate any machining of the surface, thereby leaving intact thesurface's structural integrity. The thin films are vacuum deposited directly onto the surface and havethicknesses on the order of a few micrometers. As a result, thin film sensors add negligible mass tothe part and cause minimal disturbance to the gas flow over the surface. The fabrication of thin film sensors is normally performed in clean room environment to minimizepossible contaminations. For an electrically conductive substrate (such as superalloys used in jetengines) an electrically insulating sub layer must first be vacuum deposited by electron beamevaporation or by sputter deposition. The insulation sub layer consists of several layers of vacuumdeposited ceramic coatings optimized for better adhesion and surface homogeneity (Fig. 1.1a). Thethin film thermocouple sensors are then fabricated directly on top of the ceramic insulation layereither by sputter deposition through stenciled shadow mask or by a photolithography/etchingprocess (Fig. 1.1b). As mentioned before, all deposition processes in this sequence require highvacuum from 10-4 to 10-6 Torr.Different metal films are used to form a thermal junction, and one of the most common is alumel-chromel junction (so called K-type junction). Chromel is an alloy of around 90% of Ni and 10% Cr.Alumel contains around 93-96% of Ni, 1.8-2.5% Al and the rest is Mn and Si. Alumel-cromeljunction exhibit linear behaviour of thermal voltage in wide range from 0 to 1000ºC. To fabricate athin film alumel-chromel thermocouple by vacuum deposition one has to perform electron beamevaporation of alumel and chromel targets alternately at high vacuum. The delineation of the sensoris produced by a slit mask held in proximity to the substrate. After the deposition of the first metalthe substrate has to be removed from the vacuum, the slit mask has to be realigned to delineate thesecond metal of the couple, and the substrate is then again reloaded into vacuum chamber [4].Alternatively, one can use a photolithigraphy/etching processes for sensors delineation, although inthis case the process sequence is no less complex.

Figure 1.1b: Schematic cross-section of a thin film thermocouple [4]

Figure 1.1a: Schematic diagram of layers deposited in a thin film sensor fabrication [3]

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The thickness of the thermocouple films obtained inthese processes is around several micrometers. Inprinciple, laser direct writing by a selective ablation ofthe metal film could be used for thermocouple patterningreplacing the slit mask/photolithography methods.However, this would not eliminate the need insophisticated vacuum deposition processes and only addup the costs of a high power ultra-fast laser materialprocessing system.

Figure 1.2: TFTC sensors on Space Shuttle main engineturbine blade [3]

1.2.2 Thermal spraying technology

Thermal spraying is a coating method, characterized by spraying of the coating material on thesubstrate in form of dispersed condensed particles by a gas jet. The process requires also a source ofthermal energy to melt or fuse the sprayed material. The coating material reaches the substrate inform of fine dispersed particles, melted or plasticized together, which form a laminate coating layerwith good adhesion properties. Thermal spraying is distinguished from other coating methods, such as vacuum deposition orwelding processes (surfacing), by a jet of condensed particles. Thermal spraying method is anintermediate case between the welding (surfacing), where new phase is formed from the materialsof the substrate and coating, and the vacuum deposition, where the coating is formed from theatoms and ions of the coating material, condensating on the substrate. Considering the involvedphysical processes, the boundary is not very clearly pronounced. During thermal spraying meltingof the substrate surface can occur, as well the deposition of the coating from the vapour phase [6].However, this method is clearly distinguished owning to the numerous technological advantages ithas to offer: reduced processing cost, versatility in coating materials and high processing speed [5].Consequently thermal spraying is widely used for applying protective coatings to machine partsoperating in harsh environments, such as turbine blades. In case the thin film thermocoupletechnology is employed, the protective ceramic coating can be applied directly after the sensorfabrication, embedding the sensor within the protective layer.

1.2.3 Fabrication of embedded thermocouple by thermal spraying method in conjunctionwith ultra-fast laser micromachining

The novel idea of the research group from FS-Julich was to verify the possibility of using differenttypes of thermal spraying processes through the entire sequence of the embedded thermocouplefabrication, eliminating the need in sophisticated vacuum deposition techniques and clean roomenvironment requirements. According to this concept, thermal spraying can be utilized not only toapply the final protective overcoat, but also to produce the insulative sub-layer and the film metallayers, comprising the thermocouple. This would allow to fully exploit the technological advantages

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of thermal spraying method and greatly reduces the costs and enhance the overall speed of theprocess. However, because of the nature of the thermal spraying process, which forms the coatingout of dispersed microparticles, the layers produced by this method cannot be considered thin. Infact the minimal thickness of coating produced by the single run of a thermal spray jet is around 20µm [5]. Because of the same reason, the resolution of the slit mask placed in front of the substratecannot be very high. Therefore, the delineation of the thermocouple sensor has to be produced by adifferent method. Relatively large thickness of the metal layer does not allow the effective use ofphotolithigraphy/etching techniques. Alternatively, exactly at this process stage the direct laserablation could prove effective. In more precise definition, the task would require micromachining ofthe thick metal film with ultra-fast laser system with minimal damage to the substrate (ceramicinsulative coating) to produce the desired quasi-planar pattern of the thermocouple sensor. To verify the possibility of the proposed concept, a number of generic test samples were prepared inFS-Julich by all-thermal spraying process sequence. At the first stage, an insulative ceramic coatingwas applied to a metallic substrate by use Atmospheric Plasma Spraying process, in which thesource of the thermal energy is a plasma arc. The insulative layer consists of two sub-layers:MCrAlY bond coating, widely used to reduce mismatch of thermal expansion coefficients of metalsand ceramic, and Yttria Stabilized Zirconia (YSZ) which acts as an insulator (Fig. 1.3b). Due to thealready mentioned peculiarities of thermal spraying method the roughness of the obtained coatinglayer was around 12 µm, which is very rough compared to the vacuum deposition. Then an alternatethermal spraying method – High Velocity Oxy-Fuel Flame Spraying (HVOF), was used to producethe layers of thermocouple metals (alumel and chromel), which coincide at 5 mm wide regionforming a prerequisite for a thermocouple junction. In HVOF process the source of thermal energyis a continuous gas combustion under high pressure in a combustion chamber. The combustion gasused in the process was a pure Hydrogen, which was supplied in a mixture with Oxygen to thecombustion chamber (Fig. 1.3a). The process was performed at two stages feeding alternatelyalumel and chromel powder, and applying simple rectangular mask to produce different metalcoatings on the two sides of the sample and an overlap region at the sample centre(Fig. 1.3c).

Figure 1.3b: Coating layers on thetest sample

Figure 1.3a: Schematic layout of HVOF spraying proces s [7]

Figure 1.3c: A test sample coated with alumel and chromel byHVOF method. A 5 mm wide overlap region can be identified atthe centre

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By varying HVOF process parameters, different thickness of metal layers on test samples wasobtained, ranging from 22 to 60 µm. Consequently, the thickness of the bimetallic layer in theoverlap region could reach 80 µm, meaning that a thermocouple could form an 80 µm ledge on themachine part surface. But, due to the fact that characteristic thickness of hydrodynamic boundarylayer is around 1 mm, this would still pose only minimal intrusion to the investigated operationregime. Still, the width of the thermocouple sensor has to be in the order of several tens ofmicrometers, in order to minimize the mass added to the surface and increase the temporal andspatial resolution (in direction parallel to the overlap) of temperature measurements. Thisrequirement dictates the resolution of the micromachining that has to be achieved by ultra-fast lasermaterial processing system, intended to be used for sensors delineation. In this process all theunnecessary metal coating has to be removed by direct laser ablation, leaving only thin metal stripes(several 10 µm wide) of thermocouple metals, with overlap region forming a thermal junction. Athermocouple will measure the temperature difference between the junction point and the reference,which normally will be a base of the turbine blade or other machine part of interest, considered asbeing kept at the ambient temperature (Fig. 1.4).

Figure 1.4: Schematic of thermocouple sensor circuit formed by delineation of bimetallic layer by adirect laser ablation

The primary objective of the present work is therefore, the proof of the possibility of obtaining apattern of metal lines in width of several tens of micrometers, isolated from one another by aceramic gap, by means of ultra-fast laser micromachining of bimetallic thick film-ceramic substratesystem. Hence, an experimental study of the ultra-fast micromaching process was carried out usingthe test samples provided by FS-Julich (see Fig. 1.3c) and ultra-fast fiber amplified laser system formaterial processing. As mentioned before, many studies have proved the possibility of thesuccessful application of ultra-short pulse laser ablation to the patterning of the thin films withresolution down to sub-wavelenght dimensions [8]. However, the need of obtaining a quasi-planarpattern on a thick film more typical to the field of micromachining of the micro-scale machine partsand the unusual structure of the processed material obtained in the thermal spraying processrequired experimental investigation. During this investigation the theory of ultra-short pulse laserprocessing of metals, described in the next chapter, was methodically applied to the particulartechnological task, in order to identify the optimal process and laser parameters and sketch thecontours of the future possible production process.

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overlap region

Thermal junction (T1)

Chromel

Alumel

The investigated part surface Base

Reference junction (T2)

Extension wire

Extension wire

Data acquisition system

Chapter 2

Ultra-short Pulse Laser Processing of Metals

2.1 Basics of ultra-fast laser material interactions

In this chapter the review is restricted to the ultra-short laser pulses interactions with metals,omitting the case of dielectrics and semiconductors. Although, many features of ultra-fast light-matter interactions have a universal character, this restriction is made for convenience to closelyreview the case of ultra-short pulse laser processing of metals. From a physical point of view, metalis considered a system of quasi-free electrons and a crystal lattice, in which the light-matterinteractions are dominated by a free-carrier absorption of the electrons. Then, on a time scale of acharacteristic electron-phonon interaction, the energy absorbed by the electrons from the photons istransferred to the lattice in carrier-phonon scattering processes. After this process reaches anequilibrium, the electron-lattice system can be characterised by a certain spatial temperaturedistribution, which then obeys the laws of the classical heat conduction. However, because thecharacteristic time of electron-phonon relaxation is around several picoseconds or tens ofpicoseconds, ultra-short laser pulses in picosecond or sub-picosecond range deposit all of theirenergy to the electron subsystem and vanish by the time the thermal equilibrium is established. Thisleads to the situation, in which during the time of irradiation by an ultra-short impulse the materialexists in a non-equilibrium state, which is defined by two distinct temperatures of electron andlattice subsystems (the so called two-temperature model). In the next step towards the systemthermalization, highly superheated electron gas acts as a heat source for the lattice subsystem (Fig.2.1). In this complex non-equilibrium process the thermal diffusion length for the lattice subsystemis no longer dependent on the duration of the original laser pulse, but on the electron-phononcoupling and the electron subsystem diffusivity, which are solely the processed metal parameters.For the given metal this thermal diffusion length is the absolute minimum, which corresponds to theminimum possible heat affected zone (HAZ) in the laser ablation process and the maximummicromachining precision. The progress in ultra-fast laser technology through the development of chirped pulse amplificationlead to the emergence of laser systems capable of generating ultra-short pulses of high peak powerat high repetition rates. These systems are increasingly investigated for the feasibility in variousindustrial applications, as in the present work.

Figure 2.1: Schematic drawing of the ultra-fast lasermetal interaction [9] The laser energy is absorbed by free electrons. Thisprocess is followed by an energy transfer to the latticeand a heat diffusion into the solid. γ — the electron-phonon coupling constant. Two processes proceedsimultaneously in the electron subsystem: electroncooling by electron-phonon relaxation and thediffusion of the hot electrons inside the material.

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2.1.1 Minimization of the heat affected region as a major advantage of the ultra-short pulselaser processing

Laser ablation, within the meaning of a material removal by short high-intensity laser pulses, can beachieved also by longer duration nanosecond laser pulses. In fact, the simplified condition forapplication of the technique can be written as [8]:

Δh≈max {lth , lα} (2.1)

Here Δh is the thickness of the layer ablated per pulse, lα=α−1 is an optical penetration depth and

lth is the thermal penetration depth in laser material interactions. In case of non ultra-short pulses,the characteristic thermal penetration depth is approximately equal to √D th τp , where Dth isthermal diffusivity of the material and τp is the duration of the laser pulse. Due to the strong free-carrier absorption of metals, the optical penetration depth is very short (several 10 nm) and isalways exceeded by the thermal penetration depth (expect the case of the low fluence ultra-shortpulses that will be discussed later). In fact, in metals the above condition can be adequatelyfulfilled only in case of the irradiation by ultra-short laser pulses, in which lth reaches its lowerlimit defined by the parameters of electron-lattice system. In this case, the thermal penetration depthcan be approximated as √D e τe− ph , where τe-ph is the characteristic electron-phonon relaxationtime and De is the hot electron diffusivity. Here it is worth to remember that lth is nothing elsethen the width of Gaussian-like temperature distribution. In case of a nanosecond pulse it is thedistribution of temperature of electron-lattice system by the end of pulse duration. In case of theultra-short (pico-, or subpicosecond) pulse it is the distribution of lattice subsystem temperatureafter the characteristic time of electron phonon relaxation. From this point of view it is natural toassume, that with the minimization of lth the width of HAZ and the thickness of the molten layerformed during ablation process will also tend to their minimum. This simplified concept alsopredicts that there can be no additional reduction of HAZ because of the additional decrease of thepulse duration below the electron-phonon relaxation time. Comparing between the holes fabricated by a nanosecond and an ultra-short femtosecond laserpulse one can notice an obvious difference in morphology (Fig. 2.2a,b).

Figure 2.2a,b: SEM pictures of holes fabricated in the thick steel foil by a Ti:sapphire-laser [9](a) - τp = 3.3 ns, ϕ = 4.2 J/cm2, (b) - τp = 200 fs, ϕ = 0.5 J/cm2

Around the hole, fabricated by a nanosecond impulse laser, there is a considerable width of material

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that experienced melting and resolidification. There are also pronounced burrs left by a moltenmaterial which was ejected from the hole through liquid phase expulsion during the ablationprocess. On the contrary, around the hole fabricated by an ultra-short impulse laser the width of theregion with the signs of melting and resolidification is very thin, and any signs of burrs can barelybe noticed. Again, the simplified intuitive explanation to these experimental evidences is that, incase of ultra-short laser ablation, the heat affected region almost coincides with the ablated volumeand only minor heat affected zone remains after the ablation.

2.1.2 Threshold fluence and its dependence on the pulse width

On the Fig. 2.2 apart from the pulse duration there is an another parameter ϕ which has dimensionsof energy per unit area. It is the laser pulse fluence – the energy of the laser pulse per unit surface ofthe irradiated spot. Therefore, it is the characteristic of the amount of energy deposited in thematerial by the pulse per unit surface. It can be also noticed that the fluence of the nanosecond pulsewas almost one order of magnitude higher then that of femtosecond, and yet the authors emphasizethat the fluences used for holes fabrication were the lowest necessary to obtain a significantablation, slightly above the ablation threshold fluence. The existence of well defined thresholdfluence above which macroscopic amounts of material are ejected from the irradiated surface isvery typical in short and ultra-short pulse ablation processes. The difference in the pulse fluencenecessary to initiate the ablation for nanosecond and femtosecond cases implies, that the ablationthreshold fluence has to be dependent on the pulse duration. This dependence can be analysedemploying simple energetic considerations, as was already done by a number authors [10,11].Not considering the complex thermodynamics of ultra-fast phase transitions, the condition forablation threshold can be defined as condition for significant evaporation: the energy deposited inthe target solid per unit mass exceeds the specific evaporation heat Ω. In a simplified form thiscondition can be written as:

Ω∼I thr τ p

ρ lth

(2.2)

Here Ithr is the ablation threshold intensity (assuming rectangular pulse for simplicity), ρ is the targetmaterial density and τp and lth are the pulse duration and thermal penetration depth. In case of ananosecond impulse lth∼√(Dth τ p) . Thus:

Ithr∼ρΩ√ Dthτp

; φthr∼I thr τ p∼ρΩ√Dth τ p (2.3)

These relations predict that for nanosecond pulses the ablation threshold fluence decreases with thepulse duration as the square root of the pulse width. However, it was based on assumption that theprocess of the energy transfer inside the target is governed by the thermal diffusion at theequilibrium of the lattice-electrons system. For the case of irradiation by ultra-short impulses, this isobviously not true. Intuitively, one can suppose that as in the case of thermal penetration depth,there exists a lower limit for the ablation threshold fluence, below which it no longer decreases withthe decrease of the pulse duration. Indeed, from the relation (2.3) for the threshold fluence, one cannotice that it is similar to the energy for evaporation per unit volume times thermal penetrationdepth, which gives exactly the energy per unit surface needed for a significant evaporation to occur.

φthr∼ρΩ√Dth τ p∼ρΩ lth (2.4)

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As mentioned before, for the case of ultra-short laser pulses there exists a different expression forthe thermal penetration depth, which defines the width of temperature distribution in the latticesubsystem and depends solely on the material parameters. Strictly speaking, because the ultra-fastprocess cannot be considered thermal the more precise term – effective laser energy depositiondepth, is sometimes used in this case [10]. Here the designation of lth is maintained having themeaning of qausi-thermal penetration depth in the lattice subsystem. Now it can be clearly seen thatfor the pulse width shorter then the characteristic electron-phonon relaxation time, the ablationthreshold fluence is a constant for a specific material, which doesn't depend on the pulse duration.

φthr∼ρΩ√De τe− ph (2.5)

The described behaviour of the ablation threshold fluence with the pulse width has solidexperimental evidences [12].

Figure 2.3: Pulse width dependence of the threshold fluence (damage fluence) of a gold thin film(the mirror and the grating) irradiated by laser pulses at 1053 nm [12]

2.1.3 Ablation depth per pulse. Low and high fluence regimes.

For the ultra-short pulse laser ablation the thickness of the layer ablated per pulse typically shows alogarithmic dependence on the pulse fluence. This dependence can be predicted by a simplifiedtreatment of ultra-fast laser material interaction with Two-Temperature Model (TTM). This modeluses the concept of electron and lattice quasi-equilibrium temperatures to write two separate heatconduction equation for the electrons and the lattice. In these equations electron-phonon relaxationterm acts as a heat sink for the electron subsystem and as a heat source for the lattice subsystem[13]. Using this model and making several assumptions to simplify the mathematical treatment, onecan derive the following analytical expressions for the one-dimensional lattice temperaturedistribution, established after the characteristic time of electron-phonon relaxation [14]:

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T lat (z )≃φa

Clat lαexp (−z / lα) (lα≫lth) (2.6)

T lat (z )≃φa

Clat lth

exp (−z / lth) (lα≪lth) (2.7)

Here ϕa is the absorbed laser fluence, Clat is the heat capacity of the lattice, and lα and lth are theoptical and the effective thermal penetration depth by an electronic heat diffusion. The relationbetween the optical and thermal penetration depth distinguish between two cases of the temperaturedistribution: the lager of these values defines the width of the temperature profile. As already mentioned, in metals the optical penetration depth is in the order of 10 nm and willalways be smaller then the electronic heat diffusion length, which defines a thermal penetrationdepth. However, for the electronic heat diffusion to occur, the energy distribution in the electronsubsystem has to be thermalized. It is assumed by some authors [11,14] that at sufficiently low laserpulse fluences the low density of the hot electrons prevents the development of electronic heatdiffusion process. In other words, with laser pulse fluence sufficiently low and electron-phononcoupling of metal strong enough, it is possible that all electrons will transfer their energy to thelattice, omitting the process of thermalization in the electron subsystem and the onset of thediffusive heat transfer. In this case, a ballistic transport of non-thermalized electrons can still occur,having the effect of carrying energy away from the surface [15]. The effect of the ballisticpropagation of electrons is not included in TTM and represent one of the limitations of this model,that sometimes requires to be phenomenologically extended to obtain adequate results [8].Nevertheless, it may be possible that at sufficiently low laser fluences the energy transport from theelectrons to the lattice is restricted to the optical penetration depth [14]. lth is sort of tends to zeroin this case, and the width of energy deposition inside the lattice is defined by lα , which is inorder of several 10 nm. Consequently, the width of the heat affected zone is reduced even below thelimit defined by the electronic heat diffusion and electron-lattice coupling. On the other hand, if the laser pulse fluence is high enough, it will produce sufficient electrondensity for the electron subsystem to thermalize and undergo a diffusive heat transfer before theelectrons are cooled down through the electron-phonon relaxation channel. This case is inagreement with previous considerations, in which the hot electron diffusion defines the thermalpenetration depth inside the material. The lattice temperature profile is given then by the equation2.7 [14]. To derive the dependence of ablation depth on fluence, the previously defined condition for asignificant evaporation can be used in a form: Clat T lat⩾ρΩ . Then from equations 2.6, 2.7 onecan find [14]:

Δh≃lα ln(φa

φthr) ; φthr≃ρΩ lα (lα≫lth) (2.8)

Δh≃lth ln(φa

φthr) ; φthr≃ρΩ lth (lα≪lth) (2.9)

It can be noticed that the threshold for the low fluence regime is lower than for the high fluenceregime. A very obvious result, considering that the energy deposition is confined by the opticalpenetration depth. Due to the same reason however, the increase of ablation rate with the fluence ismuch steeper for the high fluence regime. By this difference in steepness, the two ablation regimescan be well distinguished on the plots of the ablation rate versus pulse fluence on a logarithmicscale. On the Fig. 2.4 there is a typical ablation rate versus pulse fluence plot, obtained by measuring the

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depth of the ablation crater produced by 100 subsequent impulses with a white light interferometer[15]. Two distinct lines can be distinguished on the logarithmic scale corresponding to the low andhigh fluence ablation regimes. The intersection with the fluence axis gives the value of the thresholdfluence and the line steepness has the meaning of thermal and optical penetration depth accordingly.As on the Fig. 2.4, the two ablation regimes are sometimes referred as gentle and strong ablationphase. With the transition to the high fluence regime the ablation rate starts to increase more rapidly.However, in this case thermal penetration depth due to the electronic heat conduction gives rise to arelatively small (compared to nanosecond case) but still considerable amount of melt, formed in theregion, where the material temperature rises above the melting threshold. This corresponds to thesituation already discussed in 2.1.1, in which ultra-short laser ablation produces a small, but finiteand still observable, heat affected zone. At the gentle ablation phase the heat affected zone isminimized to the order of the optical penetration depth, the absolute lower limit allowed by physics.

Figure 2.4: Ablation rate vs laser pulse fluence for steel 316L [15](τp = 150 fs, λ = 775 nm)

The hole shown on the Fig. 2.2b was produced by laser pulses with the fluence 0.5 J/cm2, which formost metals is in a low fluence regime. An example of the hole drilled in a high fluence regime isshown on the Fig. 2.5. The signs of melting and resolidification and pronounced burrs are here verysimilar to the nanosecond pulse drilled hole morphology (Fig 2.2a), although a careful analysisshould reveal that the heat affected zone is smaller. It may be concluded that in in the high fluenceregime ultra-short laser pulses do not provide considerable advantages for material processing incomparison with the nanosecond ones. However, the ablation rates achieved in this regime are veryhigh, sometimes exceeding even the ablation rate of nanosecond pulses at the same energy. Incopper target the ablation rates of 70 μm per pulse were observed and it is not likely to be the upperlimit [11]. Fig. 2.6 shows the schematics of ultra-fast laser metal interaction at low fluence regime,in which the electronic heat diffusion is suppressed by a fast electron-phonon relaxation.

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Figure: 2.5: A hole drilled in a thick steel plate by a high-energy ultra-short laser pulses [11](Ep = 2 mJ, τp = 200 fs)

Figure 2.6: Schematic of ultra-short pulse laserablation in a low fluence regime [11]A low fluence ultra-short laser pulse generates alow density of the hot electrons. These electrons arerapidly cooled though the fast energy transfer tothe lattice and electronic heat diffusion doesn'tdevelop. The heating of the lattice leads to theablation confined by the optical penetration depth.

The gentle ablation regime can be clearly distinguished by the pronounced structure of surfaceripples, which have a periodicity of the order of the laser wavelength and an orientationperpendicular to the polarization direction of the laser light (Fig. 2.7). The ripples originate from theinterference between the incident laser light and the tangential wave produced by laser scattering onsurface defects [8]. On the contrary, in the strong ablation regime the ripple structure is washedaway by a layer of the re-solidified melt produced in the heat affected zone (Fig. 2.8). The theory of the two ablation regimes was a heuristic interpretation of the results given by Two-Temperature model, developed by the authors in a number of works [9,11,14] in order to explain theexperimental observation of the existence of two ablation regimes, different in ablation rate vsfluence dependances and in ablation crater morphologies. This theory was proven to be in a goodagreement with experimental results [14,15]. There are however alternative explanations, linkingbetween the existence of different ablation regimes and the thermodynamic processes leading toablation and different mechanisms of material removal. These explanations are briefly presented inthe next section.

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Figure 2.7: SEM images of craters produced in steel 316L by 0.5 J/cm 2 laser pulse in gentleablation regime [15]

(a) – 5 impulses, (b) – 10 impulses, (c) – 15 impulses, (d) – 100 impulses

Figure 2.8: SEM images of craters produced in steel 316L by 4.32 J/cm 2 laser pulse in strongablation regime [15]

(a) – 5 impulses, (b) – 10 impulses, (c) – 15 impulses, (d) – 100 impulses

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2.1.4 Thermodynamic trajectories of ultra-fast phase transitions. Spallation and phaseexplosion.

The condition of a significant vaporization, introduced in a previous section, simply implies that thepulse has enough energy to induce phase transition of a certain amount of material. However asrecent studies reveal, the exact mechanism of these transitions can be quite different. Thecharacteristic timescales of the involved processes should again be considered. The typicaltimescales of the processes taking place in a target, following the irradiation by the laser pulse, aredepicted on the Fig. 2.9 [16]. It can be noted that for a nanosecond pulse the irradiation time is ofthe same scale as thermal and structural effects. Therefore, with a sufficient pulse energy theevaporation will take place already during the irradiation, and the thermal diffusion will define thetemperature profile before the ablation front. For pico- and sub-picosecond pulses the irradiationtime is much shorter then the time needed for evaporation/ablation processes to develop. Therelevant parameter in this case, however, is again not the pulse duration, but the electron-phononrelaxation time τe-ph, the time needed for equilibration between electrons and phonons, andestablishing a lattice temperature profile. For a nanosecond pulse τp≫τe−ph equilibrium betweenelectrons and phonons prevails throughout the heating stage (electron and lattice temperatures areequal) and phase changes can be regarded as slow thermal processes involving quasi-equilibriumthermodynamics [16]. In contrast, in case of the ultrashort pulses the material is driven into a highlynonequilibrium state, before any structural effects or phase transitions occur, on condition thatτe− ph≪τm . τm – the time needed for structural modifications. Otherwise (τe−ph>τm) , the

process will have non-thermal nature, i.e. photochemical ablation. In metals however, this conditionis generally fulfilled.

Figure 2.9: Typical timescales of the processes involved in laser material interactions [16]

Another important characteristic timescale, linked with the structural modifications, is the timerequired for a mechanical relaxation (expansion) of the heated volume, τs≈lth /C s , with Cs – thespeed of sound in the target material. When max {τp , τe−ph}≤τs , the heating and melting takeplace under a nearly constant volume condition, causing a build up of compressive stresses [17]. The consideration of processes time scales leads to a conclusion that if the above conditions arefulfilled, the heating of material with ultra-short laser pulse can be considered an isohoric (constantvolume) process. Therefore, the thermodynamic trajectory of this process can be sketched on thetemperature-density phase diagram. It should be noted however, that the target is not assumed to

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reach thermodynamic equilibrium as a whole, but rather, the thermodynamic equilibrium can beassumed to take place locally, on scales smaller then the optical penetration depth [16]. Thethermodynamic trajectory of the process consists of: heating at constant volume up to a very highnear critical temperature, quasi-adiabatic expansion process through a solid-liquid coexistenceregion. Depending on the heating intensity (dependent on the laser fluence) and maximumtemperature reached in the isohoric heating process, different expansion trajectories are possible.On the Fig. 2.10a the reached temperature is that of the superheated solid. The following expansionleads to material melting, possibly reaching the liquid-vapor metastable region. If that happens, arapid nucleation takes place and the overheated material decomposes into a mixture of vapor andliquid droplets. This mechanism of material ejection has been called phase explosion. However, theejection of the material can take place even at lower heating intensities without significant materialvaporization. In this case the liquid layer formed due to material melting is ejected by the laserinduced stresses, generated at the surface region of the target under conditions of stressconfinement. This mechanism of material ejection is called spallation by analogy with the termused to describe the dynamic fracture, that results from the reflection of a shock wave from a backsurface of a sample. The separation and ejection of the liquid layer takes place by the nucleation,growth, and coalescence of multiple voids in a subsurface region of the target. The appearance ofthe voids coincides with the arrival of the unloading stress wave that propagates from the surfaceand increases its strength with depth [17].

Figure 2.10a: Phase explosion/spallation processtrajectory on temperature-density phase diagram

The thermodynamic trajectory on Fig. 2.10b corresponds to a very high laser fluence resulting inexceptionally intense heating. Melting occurs at the very beginning of the relaxation process, andthe material then expands in a super-critical fluid state, in a process called fragmentation. Spallationand phase explosion are the most interesting material ejection mechanisms however, since thetransition between them can be identified with the transition from a gentle to strong fluence regime.The existence of the spallation and phase explosion phenomena is supported by the experimentalobservations [18] and by Molecular Dynamics Simulations [17]. These simulations were also ableto provide the fluence dependence of the total amount of material removed from the target (totalablation yield) and the number of ejected individual (vapor-phase) atoms (Fig. 2.11). Two abruptthreshold-like transitions can be noticed: the abrupt increase of the total yield on the spallationthreshold, and the abrupt increase of the vapor phase atoms with onset of phase explosion. The totalyield of the ablated material does not increase with the fluence increasing above phase explosionthreshold, and this seems to be in contradiction with the experimental observations mentioned in theprevious section (Fig. 2.4). However, the simulations predict increase of the depth of molten layer

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Figure 2.10b: Fragmentation process trajectory on temperature-density phase diagram [16]

in the phase explosion regime, and the authors suggests that the experimentally observed increase inablation rate can be explained by the contribution of melt expulsion to the ablation yield [17].Indeed, if the fraction of the vapor phase increases, it should apply a recoil pressure on the moltensurface layer. The explanation to the increase of the melt thickness during phase explosion is theinteraction of the vapor with the target surface long after the explosive decomposition, whichprolongs the melting time and therefore the melting depth. Moreover, some ejected droplets of amolten material can be redeposited on the target surface and pushed back by the same recoilpressure of the expanding vapor phase. On the contrary, spallation of the molten layer interrupts theheat conduction process and leads to the decrease in melting depth, comparing with a pure meltingprocess. This can be regarded somewhat similar to the HAZ minimization mechanism described in2.1.1.The assumption of melt expulsion taking place in the phase explosion regime is consistent with theobserved drilled holes morphology at high laser fluences (Fig. 2.5). At low fluence regime thespallation mechanism results in ejection of thin molten layers (decomposing to nano-scale dropletsafter-wards) with minimal melting depth. Thus, the theory of transition from spallation to phaseexplosion is able to provide adequate explanation to the experimental observations without the needof assumption on the energy confinement within the optical penetration depth at low laser fluences,

Figure 2.11: Total ablation yield (a)and number of individual (vaporphase) atoms in the ablation plume(b) as functions of the absorbedlaser fluence [17]

Results are for simulations of a bulkNi target irradiated with 1 ps laserpulses, with an ablation plumeanalyzed at 500 ps from thebeginning of the simulation.

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2.1.5 Implications on laser processing strategies

Summarizing the results of laser material interaction theory, which have the out-most importancefor the micromachining applications, the ablation rate and the ablation precision should beconsidered. The ablation rate is known to increase with the laser pulse fluence up to considerablyhigh values, but on transition from low to high fluence regime (from spallation to phase explosion)the amount of melt participating in the process significantly increases. If the melt layer thickness isdefined to be the smallest feature size for micromachining, it can be concluded that the precisionand the achievable resolution decrease with the increasing fluence upon transition to the strongablation phase. The full potential advantage of the ultra-short pulses for the precisionmicrofabrication can be exploited only at low laser fluences, with spallation as a primary materialejection mechanism. The material removal rate at this regime is however very low, potentiallymaking the fabrication process very time consuming. From these considerations the concept of low-fluence finishing may be developed. The high fluence regime can be used for a rough machining torapidly sketch the coarse contours of the fabricated structure, and the low fluence regime – for post-processing to satisfy the precision requirements. The use of this strategy is sometimes very natural and straightforward, as for example in drilling ofthrough-holes. To maximize the process speed, drilling is preformed using high energy pulses(around 1 mJ). And yet, the quality of the hole side walls is excellent (for femtosecond laser pulses).The reason to this is the spatial intensity distribution of the laser pulse. After the drilling process isfinished, the central high intensity part of the pulsed laser beam propagates through the hole withoutabsorption, and side-wings of the beam with low intensity interact with the hole sidewalls at lowfluence regime. This results in “polishing” of the initially rough hole side surface bearing the signsof resolidified melt ejected from the hole in a rapid high fluence drilling process. The resultingmorphology of the side wall is shown on the Fig. 2.12 after a single run of the cutting beam andafter five more runs [19].

Figure 2.12: SEM image of a cut through 125 μm Mg foil produced by femtoscond laser pulses [19]

The described effect already belongs to the influences of optical parameters and the spatial pulseenergy distribution on the process of microfabrication. More of the similar effects are going to bebriefly reviewed in the following section.

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2.2 Influence of the beam shape and other optical parameters

2.2.1 Ablation in a Gaussian beam

The process parameters, whose influence was analysed in the previous section, were the pulsefluence and pulse duration. The influence of the pulse duration was for simplicity limited to the fact,that below certain threshold (electron-phonon relaxation time) the regime of ultra-short pulseprocessing is entered. The fluence of the laser pulse influences the ablation rate andmicromachining precision. However, the pulsed laser radiation propagates in space as a beam andthe fluence of the laser pulses has a certain spatial distribution, depending on the beam shape.Because of this non-uniform distribution of the fluence in space the ablation rate and quality can bequite different in different regions of the irradiated spot, the fact that will have a huge impact on themicromachining process (as in the example of through-holes drilling). Moreover, the ablation willonly take place in the region where the ablation threshold fluence is exceeded. Thus, the ablationcrater diameter has to be dependent on the beam shape. This dependence can be analysedquantitatively. Modern laser system are able to generate high quality beams, whose spatial intensity distributioncan be well approximated by a Gaussian function. The temporal profile of the pulse can alsoconsidered to be Gaussian. Thus, the pulse intensity can be written as:

I (r ,t)=I0 exp(−2r 2

w02 )exp(

−2t 2

τp2 ) (2.10)

where I0 is the peak intensity, w0 is the beam waist and τp is the pulse duration, defined at the (1/e2)intensity contour. r is the radial coordinate of distance from propagation axis. The fluence is anintegral characteristic of the pulse defined as the energy density per unit surface and can be obtainedby integration of the intensity over a pulse duration (extended to the infinity).

φ(r )=∫−∞

+∞

I (r , t)dt=φ0 exp(−2 r2

w02 ) (2.11)

where φ0=I 0 τ p√π/2 is the peak fluence. The integration over a spatial coordinate will give thetotal energy of the Gaussian beam pulse:

Ep=∫0

∞

φ(r)2π r dr=∫0

∞

φ0 exp(−2r2

w02 )2π r dr=

πw02

2φ0 ; φ0=

2 Ep

πw02 (2.12)

If the condition for the ablation is set to φ≥φthr , the diameter of the ablated spot can be obtainedfrom 2.11:

D2=2w0

2 ln(φ0

φthr) (2.13)

Thus, the squared diameter of the ablation crater is proportional to the logarithm of the peakfluence. The equation 2.13 can provide an insight on how sub-wavelength structures can be obtained withdirect laser ablation (Fig. 2.13). Although, in order to achieve this, the pulse fluence has to exceedthe ablation treshold only slightly and the ablation rate is expected to be very low.

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Figure 2.1 3 : Gaussian beam focused on the target and its spatial intensity profile. The ablated spotis smaller then the beam waist [2]

Based on this logarithmic dependence, a very useful technique for estimation of the ablationthreshold and/or the beam waist can be employed [20]. Plotting the experimental values of theablation crater diameter squared versus the pulse energy on a logarithmic scale and making a linearfit, allows to evaluate the ablation threshold energy from the intersection with the abscissa axis andthe beam waist from the line slope. The threshold fluence of the pulse can be then obtained from therelation 2.12.

2.2.2 Non-uniform ablation rate and channel/crater formation.

When a laser beam with a Gaussian spatial profile is incident on the flat target, the distribution ofthe absorbed energy per unit surface (fluence) will show the same dependence. Following theresults of ablation theory discussed in a previous section, the ablation rate can be considered mainlythe function of the fluence, and proportional to the logarithm of a ratio between the incident fluenceand the ablation threshold. Hence, the depth of the ablation crater may be expected to be initiallynon-uniform tapering towards the bottom. This effect builds up with each subsequent pulse used todrill the hole or channel by method of multi-pulse processing, further increasing the tapering anglewith drilling depth. Once the quasi-conical shape of the crater has been formed, the increase of theirradiated surface area due to the oblique incidence on the side walls can lead locally to aconsiderable reduction in the fluence [21]. At some point this reduction of fluence due to theincrease of the irradiated surface will case the absorbed fluence to fall below the ablation threshold.Once this happens the profile of the side walls remains rather fixed, and only the tip of the channelcontinues to grow in depth due to the perpendicular incidence of the beam. Figures 2.14, 2.15 show the results of numerical calculations for the case of nanosecond pulsesusing the model, which takes heat transfer into account. The geometrical dependants of heatconduction leads to the increase of threshold fluence at the narrowing tip, which eventually stopsthe ablation and interrupt the drilling process. For the drilling with ultra-short laser pulses, withnegligible thermal effects, this is not likely. Besides, the model does not take into account multiplereflections from the walls of the channel, whose effect concentrate the energy at the centre of thechannel, tending to widen and deepen the tip. Finally, non-linear distortions of the beam profile(sections 2.3.1, 2.3.2), can cause energy re-distribution inside the channel and allow fabrication ofholes of rather cylindrical shape [19]. Nevertheless, the tendency for the formation of taperingchannels narrowing to the bottom and the non-uniform ablation rate is quite pronounced especiallyfor the low fluence processing.

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Figure 2.14: Correlation between the beamprofile E (r) and the normally incident intensityEn for a numerically calculated hole geometry[21](Eth – ablation threshold fluence)

2.2.3 The effect of the beam polarization.

Polarization direction of the laser beam is known to have a strong influence on the processingefficiency of conventional laser cutting applications. If the cutting plane coincides with the plane ofpolarization (for a linearly polarized beam), the absorbance on the oblique cut surface can increaseup to 20 times due to the angular dependence of the reflectance for a p-polarized light (for CO2laser cutting) [22]. For the femtosecond laser pulses the reflectance may be significantly reduceddue to the non-linear absorption [23]. Nevertheless, the polarization effects on the ultra-short pulselaser processing can be observed sometimes, and one of the experimental evidence is the deviationfrom a cylindrical shape of the hole outlet in the ultra-short pulse through drilling [24]. The picturesof the outlet hole together with the polarization directions are shown on the Fig. 2.16. From thecomparison of the direction of the hole widening and the polarization vector, it may be concludedthat this effect cannot be attributed to the enhanced absorbance. Instead, it may be the evidence ofthe key role of the multiple reflections in the deep drilling process. The direction in which the holesare wider is perpendicular to the polarization plane, therefore the reflectance from the obliqueincidence is higher on its side walls according to the classical Fernel relations. Enhanced multiplereflections allow the propagation of the energy through the channel to the outlet. On the contrary, inthe direction of the polarization vector most energy is absorbed inside the channel and the holeoutlet remains under-processed. This unwanted effect is usually avoided employing circularpolarization of the drilling beam. This observation shows, that the polarization effects on thereflectance from oblique surfaces still can have an influence on the ultra-short pulse processing.

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Figure 2.15: Numerical calculations results for the temporal evolution of the hole profile and thecorresponding absorbed intensities for the drilling of aluminium [21]

Figure 2.16: Outlet of the hole drilled in steel invacuum conditions with 170 fs Ti:sapphire laserpulses [23]The direction of the polarization vector of the drillingbeam is shown besides.

2.3 Influence of the ambient gas atmosphere

Previous discussion was mainly limited to the case of the laser processing of metals in a vacuumenvironment. The holes shown on figures 2.2, 2.5 were drilled in a vacuum cell at a pressure below10-4 mbar. Thus, the influence of the ambient gas atmosphere was excluded from consideration. Inpresence of the ambient gas new phenomena can be observed: non-linear interactions of the laserradiation with a gas atmosphere and optical breakdown and plasma ignition in front of the target.These phenomena are clearly capable to affect the processing significantly, and one of the reasonsto study them is to understand the limitation of the processing in an ambient air atmosphere (whichis most simple from technological point of view).

2.3.1 Beam distortion due to the non-linear effects and optical breakdown

The non-linear interactions of the intense focused beams with an air medium include self-scattering,self-phase modulation and multi-photon ionization. The abrupt ionization of the ambient gasproduces a microplasma with non-uniform refractive index, further facilitating the beam distortionand spectral modifications. Since for the electrons with the energy of a few eV the characteristictime between collisions with the ambient gas molecules is on the order of 1 ps, an avalanchebreakdown cannot develop during sub-picosecond pulse irradiation. This implies that, for the gasionization with ultrashort laser pulses, the multiphoton process is dominant [25]. Probably due tothis fact, the screening effect of the produced microplasma is relatively weak. But the enhancementof scattering effects can be rather strong for sufficiently high intensities (Fig. 2.17).As can be seen from the Fig. 2.17, the threshold fluence for this phenomena is rather low –noticeable scattering starts already below 10 J/cm2. When the threshold is significantly exceeded,the beam starts propagating within a wide scattered cone, the tip of which is located in front of thefocal plane of the focusing lens. At fluences below 5 J/cm2 these effects may be expected not to playa significant role. Another possibility of keeping the irradiation intensity below the threshold is toemploy picosecond pulses instead of femtosecond ones. It has been shown that for the pulses with

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duration of several picoseconds the scattering effect is rather low even at high pulse fluences [26].Therefore, it may be advantageous to some point to use picosecond pulses for microprocesing,provided they are still shorter then the characteristic electron-phonon relaxation time of theprocessed metal.

Figure 2.17: Scattering and absorbtion in gases measured as a divergence from Gaussian beamprofile of 800 nm, 120 fs pulses [26]

2.3.2 Contamination of the ambient gas by the ablated particles and plasma shielding.

Plasma shielding effect plays very significant role in a conventional laser processing. In theprocessing with ultra-short pulses laser-plasma interactions play much less important role, since theexpansion of the vapor plume appears much later after the end of the pulse, and therefore cannotinteract with it (at least not with the same pulse which caused the ablation). Phenomena like lasersupported absorption wave are totally absent in ultra-short pulse ablation. However, it was alreadyseen in the previous section that in presence of an ambient gas multiphoton ionization can producemicroplasma near the focal spot of the laser beam, which leads to the minor absorption andpotentially significant scattering of the pulse energy. This effect may also to some extent beperceived as a plasma shielding. Apart from that, the term plasma shielding in context of the ultra-short pulse laser ablation is more often referred to the interaction of vapor/plasma produced by theablation products with the subsequent pulse during multi-pulse laser processing. Fig. 2.18 shows the results of pump-probe experiments that were carried out in order to investigatethe transmittance of vapor-plasma products shortly after the ablation by a femtosecond pulse. Thefirst transmission minimum is reached after about 5 ns and has almost 100% attenuation. Thesecond minimum with a reduction in transmission down to 60% follows 100 – 200 ns later. Afterthe second minimum the transmission increases to the initial value after about 2 – 3 µs [27]. The first minimum can be attributed to the free carrier absorption of plasma formed on a short timescale by multiphoton ionization of the ambient gas together with electrons and partly ionizedsublimated mass emitted from the target surface. Second minimum showed no dependence on theprobe wavelength therefore it was attributed to a Mie scattering on large conglomerates of ablatedparticles of size in order of the wavelength. It may be concluded that the initially formed plasmarelaxes on a time scale in order of nanoseconds, and afterwards the attenuation is mainly due toscattering on the ablated particles. It was also shown that the attenuation in the first minimum,attributed to a plasma absorption, is strongly dependent on the pulse fluence.

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Figure 2.18: Temporal transmission of probe pulses with a wavelength of 400 nm for the ablationof Aluminum with 200 fs pulses at a fluence of 17 J/cm [27]

The recovery of the original transmittance implies that the scattering particles were removed fromthe vicinity of the ablation crater surrounding by convective flows. In case of deep drilling ornarrow structures fabrication, these particles may reside in the vicinity of the processing regionmuch longer. The presence of electrically charged particles increases the degree of ionization inatmosphere, therefore significantly increasing the concentration of free carriers that can be heatedby a succeeding pulse, which may lead to an avalanche breakdown of the ambient gas. Thisphenomena is known as low-threshold breakdown due to contamination by the ablated particles andacts as a shielding for the incoming pulses and as an additional heat source [25, 28]. This effect can be observed and characterized by measuring optical losses in through channelsdrilled by short laser pulses in the air. The channel transmittance to the propagating pulses appearsto be strongly dependent on the pulse energy and the time interval between succeeding pulses(repetition rate). On the Fig. 2.19 the dashed line corresponds to the situation in which the beamwould propagate through the drilled channel without losses. However, the measured values of thetransmittance suggests, that at a certain relatively low energetic threshold (15-20 J/cm2) a plasma isignited, and even relatively low repetition rate of 5 Hz is sufficient to maintain the plasma andeffectively block most of the incoming radiation [29]. The plasma shielding, induced by presence of ablated nanoparticles, is especially pronounced insidelong and weakly ventilated channels. Recovery of optical transmittance occurs by deposition ofparticles on the walls, or by their removal with convective air flows. The recovery time is mostlydependent on the fluence used for the drilling, which defines the initial concentrations of theparticles, and on the geometry of the channel, which limits the possibility of the convective removalor determines the mean path for the particle to reach the wall. Employing of vacuum below 200mbar completely eliminates the shielding [29]. It should be noted, that the described experiment was performed using relatively long 300 ps pulses.For the sub-picosecond pulses the interaction time is still likely to be insufficient for an avalanchebreakdown to develop. Nevertheless, the contaminating charged particles are able to enhance theeffects of non-linear scattering and conical emission described in the previous section. Thistransformation of laser pulses in light filaments with strong divergence is likely to be responsiblefor the redistribution of laser intensity inside the channel, which allows fabrication of deepcylindrical holes, despite the limitations of the nun-uniform channel growth effect described in

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section 2.2.2 [19].

Figure 2.19: Dependences of the output energy on the incident 300 ps pulse energy for a throughopening in a 500-mm thick sample [29]

(1) - single pulse, �(2) - 5 Hz rep. Rate, (3) - 100% transmittance

Modern laser processing systems use much higher pulse repetition rate reaching several hundredkHz to maximize the productivity, and some studies actually provide a justification for using highrepetition rates to avoid plasma ignition [30]. The experimental curves on the Fig 2.20 show thatabove certain critical repetition rate value the ablation rate at atmospheric pressure is equal to that inrarefied atmosphere conditions. The authors ascribe this to the effect of rarefaction shock wavepropagating from the ablation site and creating a low density region in its wake, which lasts for acertain period of time. If the time interval between incident pulses is below certain critical value, thelow density produced by the rarefaction wave effectively increase the threshold for plasma ignition.

Figure 2.20: Dependences of the ablation rate for 0.5- mm thick steel plates on the 20 ns laser pulserepetition rate for an energy density of 100 J/cm 2 under various pressures of the surrounding air

[30]

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The authors methodically apply the theory of detonation to support their hypothesis and achievesome agreement with the experimental value of the critical repetition rate. Again, these results wereobtained for a long nanosecond pulses. However, the effect of the ambient atmosphere rarefactionshould be also beneficial for the ultra-short pulse processing, possibly reducing the degree of thenon-linear scattering and conical emission.At this point the role of the pulse repetition rate in context of laser-plasma interactions has alreadybeen discussed to much extent. In the next section other influences of the repetition rate on theultra-short pulse laser processing will be pointed out.

2.4 Influence of the pulse repetition rate and pulse-to-pulse interactions

2.4.1 Heat accumulation due to residual thermal effects

As was shown in the section 2.1, during ultra-short laser ablation the energy coupling and a materialremoval proceed on the time scale, to low for the classical equilibrium thermal conduction todevelop. This is the usual meaning of the statement, that in the ultra-short pulse ablation thermaleffects can be neglected. However, after the material ejection happened and the thin layer of meltand the heated lattice remains near a newly formed material surface, the re-solidification andthermal relaxation of the thin heat affected region proceed through conventional thermal diffusionmechanisms. The residual thermal energy relaxes through a heat conduction to the bulk material.This residual thermal energy is the energy absorbed from the laser impulse, which is left in thematerial after the ablation process. The rest of the energy is obviously removed with the ejectedmass. Simple considerations may lead to assumption that the amount of residual energy has to growwith the increasing pulse fluence. As explained in 2.1.5, in a high fluence regime during phaseexplosion the amount of vapor phase in the ablation products is increasing. Interactions of the vaporwith the target surface prolong the melting duration and can lead to the redeposition of the hotejected droplets. In the presence of the ambient gas, the ionized plasma produced nearby the targetsurface can be an additional source for residual thermal energy, which is transferred to the targetthrough radiative and/or convective mechanisms [25]. On the Fig. 2.21 K = ER/Ep , where Ep is theincident pulse energy, and ER is a residual energy measured calorimetrically. Above the fluence of 1J/cm2, this residual energy coefficient is starting to increase, signalling the transition to the highfluence regime and/or the increase of the heat transfer from plasma.

Figure 2.21: The residual energy coefficient for platinum as a function of laser fluence in variousambient gases at a pressure of 1.08 at m [25]

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Fortunately, being an equilibrium thermal effect, the residual energy relaxation can be analyzed bymeans of the classical heat transfer theory [31]. ER = K Ep can be considered as a point energysource proportional to the energy of the pulse, while K is a residual heat coupling coefficient, whichremains to be unknown time-dependent function of the process and material parameters. Thesolution of the corresponding heat transfer problem of the point energy source in a 3-coordinatespace is given by:

ΔT=ER

ρCp(4π k t )3 /2exp(

−r2

4 kt) (2.14)

where r is the distance from the ablation site, t is the time variable and the rest are the thermo-physical constants of the material. The assumption is taken of course, that these constants do notvary with the temperature. The model also does not take into account any phase changes.The equation 2.14 is the temperature field produced by the residual energy coupling from a singlelaser pulse. In order to obtain the multi-pulse temperature field, it is required to sum up thecontributions of the multiple thermal energy sources over a certain period of time. From theobtained expression the surface temperature (r = 0) can be obtained and plotted versus time (Fig.2.22a,b).Figures 2.22a,b both show the time evolution of the temperature at the location of the heat energysource for two repetition rates of 50 and 250 kHz. The difference is that in the Fig. 2.22a the energyinput is the same (5 μJ), and in the Fig 2.22b the energy input for the 50 kHz repetition rate ishigher (25 μJ). This increment in the heat source energy tries to model the situation, in which theaverage power of the processing remains constant, while the pulse energy is reduced and therepetition rate increased. It is the general strategy of increasing micromachining precision reducingthe pulse energy down to the low fluence regime, but at the same time, trying not to sacrifice theprocess productivity and therefore, increasing the pulse repetition rate. However, it can be noticedthat the heat accumulation effect may pose a fundamental limitation to this power scaling technique.The dashed line on both figures indicates the melting threshold of the material. The infinite increaseof the temperature at the time of pulse incidence is of course unphysical, and the temperaturesabove the melting threshold also do not have any physical meaning. They can merely provide anestimation of the temperature in the surrounding material on the time scales very long, compared tothe irradiation and ablation. Nevertheless, even such simple model can reveal the effect of the heataccumulation at high repetition rates: on both figures at the repetition rate of 250 kHz thetemperature rises above the melting threshold already after several pulses.

Figure 2.22a,b: Temporal evolution of the temperature on the surface of a semi-infinite body ofCrNi-steel at the location of a point source for two different repetition rate s [31]

(a) - same heat source energy of 5 µJ per pulse, (b) - same average power input of 1.25 W, i.e.with a source energy of 25 µJ at 50 kHz and 5 µJ at 250 kHz

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It should be again emphasized, that at the Figure 2.22b the average power input remains constantfor both repetition rates. It may allow to conclude, that even when the residual energy coupling isrelatively low it may potentially lead to the melting due to heat accumulation, if the repetition rate ishigh enough. Generalizing, for a particular residual thermal energy there exists a critical value ofthe repetition rate, which does not allow thermal relaxation and eventually causes material melting.The time dependent temperature increase due to the heat accumulation may be obtained by drawinga line connecting all the minimums of the multi-pulse temperature field. This will give theestimation of the temperature increase in the vicinity of the ablation site with each subsequentimpulse. The residual energy is considered to be proportional to the pulse energy, but the proportionalitycoefficient is strongly dependent on the experimental conditions and might vary from close to 0%up to 100% below the ablation threshold [31]. It may be regarded as a free parameter that has to bededuced experimentally. Another important result of the described heat accumulation theory is thatthe maximum tolerable average power at a given repetition rate, which allows to avoid heataccumulation, is proportional to the inverse square of this repetition rate. A rather trivial remarkwould be that the maximum tolerable average power increases with the heat conductivity of theprocessed material.The experimental observations of the heat accumulation effect are reported quite often [32,33].Figure 2.23 shows three holes drilled in 1-mm CrNi-steel plates with a pulse duration of 6 ps andthe same average power of 10.3 W [33]. It should be noted that due to effect of the low fluencefinishing (section 2.1.6), the residual energy from the ablation process is likely to be the same in allthree cases. The residual thermal energy due to plasma coupling however, may decrease with thedecreasing pulse energy, due to a less dense plasma formation inside the drilling channel.Nevertheless, it can be clearly seen that increasing repetition rate above a certain critical value leadsto a massive formation of melt. The signs of material being close to the melting threshold appearalready as additional structures in the inlet of the hole drilled at 85 kHz [31].

Figure 2.23: SEM pictures of holes drilled in 1 mm thick CrNi-steel plates [ 33 ]The upper pictures show the inlet, the lower the outlet of the drilling. The holes were drilled at thesame average power of 10.3 W with a pulse duration of 6 ps and a focal spot diameter of 25 µm.

Therespective repetition rates (f L ) and pulse energies (Epulse) used for the drilling process are notedbelow each pair of pictures [31].

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2.4.2 Incubation effect and damage accumulation

In section 2.1.2 the ablation threshold fluence for ultra-short pulse irradiation is shown to be thefunction of the material properties. It is consistent with experimental observations and withmolecular dynamic modelling results, in which the ablation threshold is associated with thethreshold fluence of transition from a surface melting to spallation (section 2.1.5). However, duringmulti-pulse irradiation of the target an interesting phenomena may be observed, in which theablation threshold fluence decreases with the number of the preceding incident pulse. After a certainnumber of the pulses the ablation threshold value starts to show saturation and is decreasing nolonger [10]. This effect is well known for the processing of transparent and semi-transparentmaterial where the absorption of the laser radiation is significantly facilitated after formation of anabsorbing layer due to accumulation of the absorbing defects or colour centres. These defects areformed by the preceding laser pulses whose absorption in the material is initially relatively low.Observed also during processing of metals, this effect was firstly attributed to the accumulation ofplastic deformations due to thermal loads and fatigue-like failure leading to thermo-mechanicalablation i.e. spallation. However, recent studies, performed using X-ray diffraction analyses, do notprove this mechanism [35]. Therefore the effect is still attributed mainly to the facilitatedabsorption, due to the accumulation of surface defects and an increasing surface roughness after themulti-pulse irradiation [34]. Molecular dynamics simulations predicts the generation of the strongsupersaturation of vacancies in the surface region, exposed to the multi-pulse irradiation belowspallation threshold. Indeed, below the spallation threshold the absorbed pulse energy can lead tothe melting and rapid re-solidification of a thin surface layer. The vacancies are generated at therapidly advancing solidification front and are stabilized by the fast cooling of the surface region[36]. These effects are likely to increase the absorbance of the metal surface layer and possiblyreduce the spallation threshold due to formation of nano-voids. If attributed purely to the facilitated absorption, the incubation effect is treated not as a reduction inablation threshold fluence, but as an increase of the absorbed fluence. Or in other words, thereduction of the ablation threshold related to the irradiation fluence. It is not dependent on therepetition rate, but rather on number of the preceding pulses, and is a characteristic effect in multi-pulse processing of polished surface metals, whose initial reflectance is relatively high. Thedependence on the pulse repetition rate may arise only in conjunction with heat accumulation.Indeed, in metals the increase of the surface temperature is also known to facilitate absorption.Some studies are trying to investigate the role of heat accumulation in the incubation effect [34],although the provided evidence may prove to be controversial. The incubation effect is commonly described by a phenomenological model analogous to thefatigue failure induced by the stress for N cycles:

Φthr , N=Φthr ,1 NS−1 (2.15)

where Φthr,N is the multi-pulse threshold fluence, Φthr,1 is the ablation threshold for a single pulse, N isthe number of the preceding incident pulses and S is the incubation coefficient, which quantifies thedegree of the incubation effect. As was already mentioned, after a certain number of pulses theincubation effect tends to saturate and no additional decrease in the threshold fluence is observed.To take this into account, a slight modification to the equation 2.15 was proposed:

Φthr , N=Φthr ,∞+ΔΦ thr ,1 N S−1 (2.16)

with ΔΦthr,∞ being the saturation value of the threshold fluence, which no longer changes withincreasing number of the pulses, and ΔΦthr,1 and S are the incubation effect parameters.The incubation effect can be deduced experimentally, employing the same technique for estimation

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of the ablation threshold described in 2.2.1. As threshold fluence decreases with the number ofincident pulses, the plot of the squared diameter of the ablated crater, produced by a certain pulsenumber, versus logarithm of the pulses fluence will yield the multi-pulse threshold fluence (Fig.2.24).

Figure 2.24: For a given fluence value, the diameter of the ablated crater increases with increasingpulse number [15]

2.5 Laser ablation cutting

2.5.1 Cutting velocity and pulse overlap

Up until now, a multi-pulse drilling with a fixed laser spot was extensively discussed, because it canbe perceived as an elementary operation in pulse laser processing. Laser ablation cutting (and ingeneral every micromachining operation) can be regarded as a gradual material ablation from thevolume that is to be removed, performed by a successive drilling of the holes, which overlapbetween each other. Of course, these holes are not drilled one after another through all the requireddepth, but the material is removed layer by layer by a laser beam, moving relatively to theprocessed part. Craters ablated by each single pulse overlap and the cut is formed, increasing indepth with each pass of the beam, i.e. the processing cycle. A rough estimation of the cutting depthafter single pass can be obtained by:

Δhpass=D f repΔhpulse

vrel

(2.17)

Here D is the ablation diameter, frep is a repetition rate, and vrel is the speed of the laser beammovement relatively to the work-piece – the cutting velocity. A uniform distribution on the ablationdiameter was assumed, neglecting the Gaussian distribution of the absorbed fluence and the effectof the non-uniform ablation rate (section 2.2.2).

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Figure 2.25: S ection sketch of the ablation cutting process [36]

In order to compare the degrees of the overlap, a parameter may be introduced, describing thespatial overlap between two subsequent laser pulses — a pulse overlap:

PO=1−vrel

2w0 f rep

(2.18)

It should be noted that for convenience, in this definition a beam waist diameter was used instead ofthe ablation diameter. This difference should be kept in mind for the applications, in which thedifference between these diameters can be noticeable. In order to achieve a quasi-linear cut profile with a Gaussian shape beam, a high pulse overlap maybe required. The effect of the non-uniform ablation rate tending to narrow the drilled holes shouldbe also considered. On the contrary, keeping pulse overlap low enough may prevent the heataccumulation effects from developing. Indeed, if only several impulses hit the same target spot, thismay yet be not enough for the material to reach the melting threshold, even if the repetition rate isabove critical value (section 2.4.1). A common compromise between the linearity of the cut, aprobability of the residual thermal effects to occur and the processing speed is the pulse overlap of75% [36].

2.5.2 Scaling the processing parameters

In order to maximize the precision and productivity of the laser micromachining, two alreadymentioned general strategies may be implemented.The first is the low fluence finishing, in which processing in high fluence regime is used to obtain apreform structure with low degree of precision. The productivity at this stage is very high due tomaximized ablation rate in a strong ablation regime. After that, the final structure of the requiredprecision is obtained performing a low fluence post-processing.The second strategy is working with low fluences through all the processing stages, but maximizingthe laser pulse repetition rate, to keep the average power high. Simple considerations may reveal,that for keeping the same pulse overlap, the cutting velocity should increase accordingly. Modernbeam guidance scanning systems known as scanning galvanometers are able to realize relativelyhigh scanning velocity, guaranteeing the desired accuracy, which allow to increase the pulserepetition rate up to several MHz. However, the upper value of the repetition rate is limited by theheat accumulation and plasma shielding effects. Designing a process using several hundreds kHz,one should already take these effects into account. From the figure 2.18 it can be seen that severalmicroseconds after the ablation the succeeding impulse may yet experience some degree of

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scattering on the ablated particles, which corresponds exactly to the several 100 kHz repetition rate.The heat accumulation effect is also likely to be observed at several 100 kHz repetition rates and asshown in section 2.4.1, can be very pronounced even at low residual heat inputs. Determination ofthe limiting values often requires an experimental investigation.

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Chapter 3

Experimental Setup and Test Samples Characterization

3.1 Laser setup for experimental material processing

In this section the experimental setup, used for the laser processing of test samples, will be brieflyreviewed. The setup consists of a high power ultra-fast fiber laser capable of generating high energyultra-short pulses required for material processing applications, standard optical components fixedon the optical table for beam guidance and delivery to a scanning galvanometer, which consists oftwo galvanometer mirrors and an f-theta lens objective. The beam is focused by the objective on thesample, positioned on XYZ manually moving stage. Since, the scanning galvanometer allows thebeam guidance on XY plane with high speed and accuracy, the manual XY translation of the stagewasn't used, expect for convenience. However, the movement of the stage in Z direction had to beused in order to properly position the sample in the focal plane of the lens objective. Varying thetranslation in Z direction, a beam size of the spot on the sample could be changed. The laser and thescanner control units were connected to a PC through the appropriated electronic interfaces, andlaser parameters and scanning process were controlled by a specialized software (Fig. 3.1). All system components were mounted on the optical table. The beam delivery from the laser outputto the scanning galvanometer was accomplished by means of standard optical components. AGalilean beam expander was used to collimate the output beam and a system of high reflectivitymirrors to guide the beam into the scanning galvanometer. To ensure a proper beam input into thescanning galvanometer aperture a very fine adjustment of the mirrors alignment was necessary.Failing to adjust the beam alignment properly, could lead to high losses of the optical power in thescanner and/or a considerable astigmatism/deformation/ellipticity of the output beam. The outputpower of the laser was scaled with an external acousto-optical modulator (AOM) installed in thehead of the laser system, and measured by an optical power meter.

Figure 3.1: Schematic diagram of the experimental setup for test samples processing

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x

yz

Laser control and power supply

Ultrafast fiber amplified laser system

Scanning galvanometer

Mirror systemBeam expander

Scanner control unit

XYZ motion stage

Test sample

3.1.1 Ultra-fast fiber amplified laser system

The laser system used in the experiment was a high power ultra-fast fiber laser Tangerine(Amplitude systems). It is an all-fiber laser based on an ultra-fast mode-locked fiber oscillator andan ultra-short pulse fiber amplifier, as well as on the chirped pulse amplification concept. An ultra-fast Yb-doped fiber oscillator, with passive mode-locking achieved by a saturable absorber mirror,produces chirped short pulses at the high repetition rate of several tens MHz. This repetition rate isreduced by a first AOM modulator used as a pulse picker. The picked pulses at a reduced repetitionrate are injected in a Yb-doped fiber amplifier. The fiber amplifier used in the system is a diodepumped large mode area photonic crystal fiber (LMA PCF), which allows the fundamental mode topropagated inside relatively large Yb-doped fiber core area, ensuring high absorbance of thepumped radiation (propagating in the cladding) and low non-linearities. The output beam is stilldiffraction limited with very high beam quality. The injected pulse is amplified, while travellingthrough the diode pumped Yb-doped core. Lower picking repetition rate of the pulse picker allow toobtain higher pulse energies. Due to the non-linear effects in the oscillator and amplifier fibers theamplified short pulse posses a frequency chirp. Therefore, it is sent to the grating compressor,allowing to further reduce its duration down to a femtosecond time scale. Since the non-linear chirpis dependent on the pulse energy, the compressor delay stage has to be adjusted according to thepicked repetition rate. On the compressor output there is an additional external AOM which allowsan additional control of the output laser radiation parameters. Changing the amplitude and thefrequency of the RF signal, applied to the AOM, allows to further reduce the pulse repetition rateand to change the diffraction efficiency, thus controlling the energy of the output pulse. Finally, amechanical shutter is positioned in front of the laser radiation, required to be open in order to obtainthe output emission.

fosc fPP fPP fPP fmod

Figure 3.2: Schematic flowchart of pulse picking and amplification

Since Yb-doped silica is used as an active medium, the working wavelength of the laser is 1030 ±10 nm. The specialized software, used for the control and monitoring of the system parameters,allowed to set the repetition rate of the pulse picker, as well as the frequency and diffractionefficiency of the external AOM. Due to the later explained considerations, the experiments wereperformed at 200 kHz pulse repetition rates, and both modulators were set to this value. The systemquality test reports provided the measured laser radiation parameters at several repetition ratefrequencies including 200 kHz, as well as the position of the compressor delay stage, required toobtain ultra-short pulses with duration below 500 fs. In fact, the measured FHWM of the pulse wasreported to be 290 fs. The measured beam quality was also very high and conformed thespecifications of M2 < 1.3. Based on these reports, the produced laser pulses were assumed to beseveral 100 fs in temporal width and to have a perfect Gaussian TEM00 spatial beam profile.The transmission efficiency of the external AOM (i.e. the diffraction efficiency) can be adjustedfrom 0 to 100%. It allows to change the amplitude of the RF signal delivered to the external AOMand, therefore, the amplitude of the diffracting acoustic wave. Varying this value through thecontrolling software allows to scale the laser output power and the pulse energies. Though here onecould not relay on the preceding measurement reports, since significant power drifts were known tooccur. An optical power meter was used to measure the actual value of average radiation power at a

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OscillatorPulse picker Amplifier Compressor External

AOM

certain AOM efficiency. The power was measured at the output of the focusing objective of thescanning galvanometer, to account for the losses on optical elements.

3.1.2 Scanning galvanometer for laser material processing

The scanning galvanometer SCANcube 7 (ScanLab GmbH) consists of two mirror galvanometersand a telecentric f-theta objective, and is especially designed for laser material processingapplications using high power near-infrared laser. The rotation of the mirror galvanometers guidesthe beam focused by the telecentric objective in two coordinates quasi-planar motion with highspeed of several m/s and more. If required, it allows to realize a 2.5-D processing concept movingthe focal plane of the f-theta objective. However, the task of focusing the beam on the samplesurface was performed by adjusting the vertical position of the manually controlled motion stage.The fine adjustment pitch of the stage was around 50 μm. Since, the diameter of the input apertureof the scanner is 7 mm and the focal length of the objective is 63 mm, the Rayliegh length could beapproximated as several 100 μm long. Therefore, positioning the sample surface precisely at thefocal plane to obtain the minimal beam waist radius of around 10 μm without any imaginginstrumentation, was a quite difficult experimental task. The chosen approach to this task was toreduce the output power to a low value using external AOM and slowly move the stage, whileobserving the target sample surface for signs of laser material interactions due to the increasingfluence, e.g. a spark of the low-threshold breakdown or a surface melting/oxidation. With pulseenergies sufficiently low these interactions were supposed to take place only in the vicinity of thefocal plane with energy density (fluence) high enough. The obtained beam waist diameter was thencalculated post-factum from the logarithmic plot of the squared ablation crater diameter versusimpulse energy (the procedure is described in details in the next section). Using this method beamwaist radii down to approximately 15-20 μm were obtained on the sample surface.The scanning galvanometer was controlled by the laserDesk ScanLab software through RTC-4 PC-interface. Using this software the desired beam pass can be plotted in a simple graphic redactor, andthe number of the runs (i.e. processing cycles) and the scanning velocity can be defined. The laserdesk software is also able to communicate with the laser control unit through the appropriateinterface. When the program run is executed, the RF signal is sent to the external AOM. Otherwise,the AOM is off and the beam is almost completely blocked by the output aperture. Thereby, thesoftware controls also the on/off-switching of the laser radiation. It is necessary not only for thestart and the end of the process, but also for the jumps between separate geometrical features of thebeam pass, required to produce the specific pattern on the processed material. Clearly, whilejumping between distinct lines, cut in a substrate processed by the ablative direct writing, orbetween several holes drilled in separate locations, the laser radiation has to be switched off.

3.2 Test samples characterization: ablation threshold of the Nickel alloythermal-sprayed coating

From the section 2.1 a conclusion can be made that, the most important material property for theultra-short pulse material processing is the ablation threshold fluence. It is needed to determine thelaser power required for the material removal in gentle or strong ablation regime, whichever isconsidered appropriate for a specific processing strategy. It also can be perceived as a measure ofthe minimum power, required for the processing of a given material. Therefore, in order define thematerial processing parameters, it is crucial to obtain an experimentally verified value for thematerial ablation threshold. The properties of chromel and alumel coatings, which had to be micromachined by an ultra-shortpulse laser, were considered to be uncertain to some degree. Chromel and alumel are both Nickel

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alloys containing more then 90% of this metal, but the properties of the thermal-sprayed coatingcould not initially be expected to be similar to the Nickel bulk material. Indeed, produced by finedispersed particles melted and/or plasticized together, the coating layer might significantly vary indensity from the bulk material. Moreover, the high concentration of microstructure defects and,possibly, nano-voids could significantly reduce the ablation threshold (similar to the defectaccumulation mechanism in the incubation effect). Therefore, the ablation threshold fluence ofalumel and chromel thick film coating, produced by the HVOF thermal spraying, had to be verifiedexperimentally.A method for an experimental estimation of the ablation threshold fluence was described in section2.2.1. The equation 2.13 can be written as:

D2=2 w0

2( ln Ep− ln E thr) (3.1)

with Ethr having the meaning of the ablation threshold pulse energy at the given focusing geometry.After plotting D2 versus Ep on a logarithmic scale and preforming a least square fitting to theequation 3.1, the slope of the line can be used to obtain the beam waist radius w0. Then, usingequation 2.12 the logarithmic plot can be rescaled to show the values of the laser fluence instead ofthe pulse energy. From this rescaled logarithmic plot the value of the ablation threshold can bedirectly obtained from an intersection with the abscissa axis. As was initially meant in [20], thismethod is also valuable, because it allows to estimate the spot size of the laser beam without anyimaging instrumentation. The rescaling calibration from the pulse energies to the pulse fluence isnecessary to establish a universal scale of values, allowing to compare different ablationexperiments independent of the focusing geometry. The measurements of the ablation crater diameter were performed using an optical microscope. Dueto the extremely rough surface of the coating layer it was impractical to measure the diameter of thecrater produced by a low number of impulses. 100 and 1000 pulses were used instead, in order toproduce a well defined ablation crater. However, the borders of the crater were still very difficult torecognize on the optical image, and the measured diameters had to be verified by an electronmicroscopy imaging. Using different pulse number for ablation threshold estimation, might alsoallow in principle to examine the existence of the incubation effect. The number of the pulses incident on the spot was defined using laserDesk software by setting theresidence time of the beam on the point-type mark. At the repetition rate of 200 kHz the residencetime of the beam to irradiate the spot by 100 pulses was 500 μs, and for 1000 pulses – 5000 μs.These point marks were produced on the sample surface with several different external AOMefficiency values, and the average laser power was measured by the optical power meter.

3.2.1 Logarithmic fluence plots with the ablation crater diameter measured using opticalmicroscope

The logarithmic fluence plots were obtained for six ablation crater series produced with increasinglaser power at the constant repetition rate. Three ablation crater series were drilled by 1000subsequent pulses irradiation, and three by 100 pulses. The 1000 and 100 pulse series were locatedat the opposite sides of the test sample (Fig. 1.3c). Each of the crater series was produced in adifferent coating type, namely: in alumel coating, in chromel coating and in bimetallic layer, inwhich the chromel coating resides on top of the alumel. Thus, the possible difference in ablationthreshold of the alumel and chromel coatings could become evident. The ablation threshold of thebimetallic coating could also be slightly different due some unexpected effects of thermal sprayingprocess. The ablation crater observed in the optical microscope was found to posses a significantellipticity, therefore two diameters had actually to be measured for each crater. Plotting each of thediameters squared versus logarithmic pulse energy/fluence revealed the ellipticity of the beam

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waist. The main reason of this was assumed to be a poor alignment of the beam input into thescanning galvanometer aperture, which was later improved. Another possible contribution to theellipticity could be made by the external AOM, which might apply some distortion to the beam atlow values of the transmission, used for the ablation threshold measurements. The diameter of each ablation crater was measured three times in a row, and the average value wascalculated. The entire error analysis was omitted, initially considering the method to be only arough estimation.

Figure 3.3: Logarithmic fluence plot for the diameters of the ablation craters drilled by 100subsequent impulses

Fig. 3.3 shows the plot of squared diameters of the ablation craters, produced by a 100 pulsesirradiation, after rescaling to a logarithmic fluence axis. “a” and “b” are the two axis of the ellipticalcrater measured and plotted separately. It can be noticed that the logarithmic fit lines correspondingto each axis lie very close together and, due to high measurement errors, can be considered asrepresenting the same beam waist radius of the elliptical spot. Although, there is some minordifference in the focal distance of the spot on alumel, chromel and bimetallic layers due to thedifferent layer thickness, it is smaller then 50 μm, and is not likely to influence the calculated beamwaist more than measurement errors. Therefore, the slopes of the logarithmic fit lines of each axiscan be averaged, and the average beam waist radii can be calculated: wa≈25±3μm and

wb≈35±3μm . Here the errors are calculated from the standard deviation of the slope valuesand cannot be considered very reliable, only giving the rough order of precision of the estimationsmade. It can be noticed that all six logarithmic fit lines roughly converge to a same intersection with theabscissa axis. The calculations show that indeed, the intersection coordinates for the three coatingtypes have only minor differences, probably well below the possible measurement errors. This can

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0,1 10

500

1000

1500

2000

2500

3000

3500

4000

4500

5000

f (x) = 2679 ln(x) + 6414

f(x) = 1288 ln(x) + 2365

f(x) = 2347 ln(x) + 5631

f(x) = 1372 ln(x) + 2541

f(x) = 2318 ln(x) + 5301

f(x) = 1017 ln(x) + 1765

a^2 Alumel100 Logarithmic (a^2 Alumel100) b^2 Alumel100 Logarithmic (b^2 Alumel100)

a^2 Chromel100 Logarithmic (a^2 Chromel100) b^2 Chromel100 Logarithmic (b^2 Chromel100)

a^2 Bimetal100 Logarithmic (a^2 Bimetal100) b^2 Bimetal100 Logarithmic (b^2 Bimetal100)

ϕ [J/cm^2]

D^2

[μm

^2]

be considered an experimental evidence, of the ablation thresholds of the alumel, chromel andbimetallic layer being equal or very similar. The differences in intersection coordinates for the linescorresponding to “a” and “b” axis diameters cannot have physical meaning, and are considered tobe purely the result of the measurements inaccuracy. Therefore again, the intersection values can beaveraged, and the average ablation threshold fluence of a thick film Ni-alloy coating for a 100pulses irradiation can be calculated: φthr ,100≈0.12±0.07J /cm2 . Fig. 3.4 shows the logarithmic fluence plot for the ablation craters produced by a 1000 pulsesirradiation. It can be noticed that on this plot, the values of the logarithmic fit slopes differ alreadymore significantly, especially for the “a” axis. It is arguable, whether this difference is due to themeasurement errors, or due to the beam waist radius varying with the different layers thickness.Assuming the measurement errors to be the major influence and employing the same averagingstrategy, the average beam radii calculated from these plot are: wa≈21±6μm and

wb≈34±5μm . Considering the inaccuracy of the estimation, the results are similar to the beamwaist radii obtained from the 100 pulses irradiation plot. A remark should be made that, the 100 and1000 pulses ablation craters were drilled on the opposite sides on the samples, and the heighdifference between them could be quite significant, due to the non-planar back surface of thesample adjacent to the motion stage.

Figure 3.4: Logarithmic fluence plot for the diameters of the ablation craters drilled by 1000subsequent impulses

The average ablation threshold fluence for a 1000 pulses irradiation calculated from the Fig. 3.4will be: φthr ,1000≈0.18±0.16J /cm2 . The difference in the intersection coordinates is stillconsidered to be the consequence of the measurement errors only. As expected, the applied methodis very inaccurate, and was only able to provide an order of magnitude of the ablation threshold. Itmay be concluded, that the ablation threshold fluence is equal or similar for all three coating types,

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500

1000

1500

2000

2500

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3500

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4500

5000

f (x) = 2676 ln(x) + 5540

f(x) = 1194 ln(x) + 1951

f(x) = 1953 ln(x) + 4337

f(x) = 640 ln(x) + 735

f(x) = 2230 ln(x) + 4884

f(x) = 873 ln(x) + 994

a^2 Alumel1000 Logarithmic (a^2 Alumel1000) b^2 Alumel1000 Logarithmic (b^2 Alumel1000)

a^2 Chromel1000 Logarithmic (a^2 Chromel1000) b^2 Chromel1000 Logarithmic (b^2 Chromel1000)

a^2 Bimetal1000 Logarithmic (a^2 Bimetal1000) b^2 Bimetal1000 Logarithmic (b^2 Bimetal1000)

ϕ [J/cm^2]

D^2

[μm

^2]

and its order of magnitude is 0.1 J/cm2. With this kind of accuracy the incubation effect wasimpossible to observe. It also may be assumed that either 100 pulses is the onset of the saturation,above which the increasing pulse number no longer influence the ablation threshold, or the surfaceof thermal-sprayed coating is initially so rough and rich with microstructure defects, that thedamage accumulation mechanism is irrelevant.

3.2.2 Verification of the ablation crater diameter using electron microscope imaging

Measuring the diameter on the optical microscope images can be very challenging for the ablationcrater drilled with both high and low number of pulses. The borders of the crater can be difficult torecognize due to a very rough surface of the coating produced by the thermal spraying. Indeed, ifthe initial surface roughness is around 10 μm, the depth of the crater drilled by 100 pulses is of thesame order of magnitude. The crater drilled by 1000 subsequent pulses has a very pronounced deepchannel at the centre of the ablation spot, but near the border the crater is still very shallow. Thereason to this is the non-uniform channel growth effect described in section 2.2.2. The result is that,although the deep central cavity of the crater is very striking, its actual borders can still be barelyrecognized. With Scanning Electron Microscope (SEM) imaging the situation is quite different. SEM imagesallow to observe the rippled structure produced by an interference with a scattered surface wave(section 2.1.4). On the images made by an optical microscope the ripples appear as a dark vagueregion, probably due to the decreased reflection from the periodic surface structure. On SEMimages the borders of the rippled region can be very well recognized, and can be considered theborders of the ablation crater.

Figure 3.5: Ablation crater produced by 100 pulses with 11 μ J energy, showing a welldistinguishable ripple structure left after the material removal

An interesting feature of the surface ripples is that their direction is perpendicular to the polarizationof the laser beam [37]. Thus, SEM images can also provide the information on the polarizationvector orientation of the linearly polarized laser beam. However, the effects of polarizationdependent absorption are not likely to play a significant role in the described crater drillingexperiments, and the ellipticity of the ablation craters can rather be attributed to the poor optical

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alignment, as mention before. Another important observation is the absence of a significant meltformation, which would wash away the structure of ripples (Fig. 2.8). This allows to conclude, thatno heat accumulation effects occur at the repetition rate of 200 kHz, and the amount of melt, formedin the ablation process itself, is negligible.Fig. 3.6a,b show the craters produced by 1000 subsequent pulses. These craters are characterized bya deep cavity at the centre, but the rippled structure is still observable and can be used to distinguishthe borders of the material ablation. The narrow and deep central cavity and shallow craters bordersare the prominent illustrations to the non-uniform ablation rate effects. The crater drilled in thebimetallic layer (Fig. 3.6b) shows slightly different morphology at the centre. It remains unclear,whether this is due to some differences in the ablation process, or to purely imaging effects insidethe deep crater narrowing to the end.

Figure 3.6a: Ablation crater produced by 1000 pulses with 11 μJ energy in the chromel coating

Figure 3.6b: Ablation crater produced by 1000 pulses with 11 μ J energy in the bimetallic coating(chromel on top)

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Fig. 3.7 shows a comparison between the logarithmic fits made using data from SEM and opticalmicroscopy images (dashed lines). Some lines look quite similar, while for the other the slope of theSEM data fit is somewhat higher. The diameters measured using SEM images are always larger dueto the clearly distinguishable borders of the craters. Assuming the ablation threshold to be the samefor all coating types and the absence of the incubation effect, the average ablation threshold basedon SEM crater measurements is: φthr

SEM≈0.14±0.08J /cm2 . It may be concluded that using SEM

instead of the optical microscope alone cannot improve the accuracy of the method.

Figure 3.7: Logarithmic fluence plot for the diameters of the ablation craters m easured with SEMand with an optical microscope

3.2.3 Logarithmic fluence plot for a tighter focusing geometry and improved beamalignment.

As mentioned before, the method of plotting squared diameters of a region damaged by laser pulseversus a logarithmic pulse fluence was initially used as a technique for measurements of the pulsedGaussian-beam spot size, which posed a considerable experimental problem in times, when a photo-diode array resolution was not high enough. A significant drawback of the experimental setup, usedin this work, was the lack of the beam imaging capabilities. Therefore this “outdated” in situmeasurement technique was the only means to obtain an information on the beam profile. In theprevious sections the beam spot was characterized as having an elliptical shape with axis radii:

wa∼20÷25μm , wb∼30÷35μm . Later on, the setup was readjusted using some practical

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0,1 10

1000

2000

3000

4000

5000

6000

f (x) = 647 ln(x) + 1047

f(x) = 3529 ln(x) + 8412

f(x) = 1891 ln(x) + 3579

f(x) = 1307 ln(x) + 2655

a^2 Bimetal100 SEM Logarithmic (a^2 Bimetal100 SEM)a^2 Bimetal100 Opt Logarithmic (a^2 Bimetal100 Opt)a^2 Bimetal1000 SEM Logarithmic (a^2 Bimetal1000 SEM)a^2 Bimetal1000 Opt Logarithmic (a^2 Bimetal1000 Opt)b^2 Bimetal1000 SEM Logarithmic (b^2 Bimetal1000 SEM)b^2 Bimetal1000 Opt Logarithmic (b^2 Bimetal1000 Opt)a^2 Chromel1000 SEM Logarithmic (a^2 Chromel1000 SEM)

ϕ [J/cm^2]

D^2

[μm

^2]

techniques to improve the beam alignment for smooth perpendicular entry into the scanninggalvanometer aperture, and adjusting the vertical position of the motion stage for tighter focusing ofthe beam. Then, another series of the ablation craters with increasing pulse energy was produced inorder to characterize the beam profile after the realignment. The craters were drilled by 500subsequent pulses and their diameters (in “a” and “b” axis) measured using an optical microscope.

Figure 3.8: Logarithmic fluence plot for the diameters of the ablation craters drilled by 5 00subsequent impulses after the optical setup realignment

The slopes of the logarithmic fits for “a” and “b” axis on the Fig. 3.8 differ only slightly, meaningthe spot no longer possesses a significant ellipticity. It seems, the improved optical adjustment washelpful in eliminating this undesired effect. The average beam waist radius is then approximatelyequal to 16 μm. The error of this estimation can assumed to be similar to that in the previous section– around ±5 μm. Even though, the large uncertainty of the obtained values prevents from makingunambiguous conclusions, the results of the laser processing experiments described later show, thatthe spot size of the beam was significantly reduced in this focusing geometry.

3.2.4 Validation of the estimated ablation threshold value

It is interesting to compare the obtained estimation with other data available on the ablationthreshold of Nickel alloys. In [38] authors use the same method of logarithmic fluence plots tomeasure the ablation threshold of nickel-based super-alloy containing around 50%wt of Ni. Inorder to improve the accuracy of the crater diameter measurements, along with scanning electronmicroscopy authors employ some more advanced methods of the Atomic Force Microscopy and theWhite Light Interferometry. Owing to that, authors claim to achieve a reasonable accuracy of theestimations and were able to observe the influence of the incubation effect (Fig. 3.9). It can benoticed that, the order of magnitude of the values obtained by the authors agrees with the estimationmade in the present work.

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0,1 1 100

200

400

600

800

1000

1200

1400

1600

1800

2000

f (x) = 507 ln(x) + 900

f(x) = 559 ln(x) + 735

a^2 Chromel500 Logarithmic (a^2 Chromel500)

b^2 Chromel500 Logarithmic (b^2 Chromel500)

ϕ [J/cm^2]

D^2

[μm

^2]

Figure 3.9: Ablation threshold vs number of pulses for a nickel-based super-alloy upon irradiationby a femtosecond laser (180 fs, 775 nm, 1 kHz) [38]

It can be also interesting to use equation 2.8 to calculate the ablation threshold of pure Nickel withthe assumption of the significant evaporation conditions as the ablation onset criteria. Assuming thestandard properties of Nickel at room temperature: Ω = 6457.285 kJ/kg, ρ = 8.908 g/cm3, α(1.03μm) = 6.289·10-5 cm-1, the calculated ablation threshold fluence is φthr ,Nickel≈0.09 J /cm2 . Thefluence used in the equation 2.8 is the absorbed pulse fluence, therefore, the calculated thresholdmay correspond to a strong incubation effect leading to almost 100% absorption of the incidentradiation. The result is quite in agreement with data presented on the Fig. 3.9 for the multi-pulsethreshold fluences. Even though, the nickel constitutes only 50% of the super-alloy thethermophysical properties of the alloyed metals are quite similar. The order of magnitude of thiscalculation also agrees with the result obtained in previous sections from the logarithmic fluenceplots for alumel and chromel alloys. The equation 2.8 was derived basing on very simple energeticconsiderations, and the fact that it successfully predicts the order of magnitude of the materialablation threshold should be noted. Another key assumption for this equation to hold, is the energyconfinement on the optical penetration depth. Therefore, it may be expected to reasonably estimatethe ablation threshold of metals with strong electron-phonon coupling and low ballistic electrontransport range (section 2.1.4). Otherwise, the energy deposition depth may exceed the opticalabsorption length even at near-threshold energy densities, and the predictive power of equation 2.8would diminish [15].Finally, in the work [17], widely cited in section 2.1.5 for its findings about different materialejection mechanisms, the molecular dynamic simulations were performed using a bulk Ni targetproperties. The ablation threshold provided by these simulations (i.e. threshold for materialspallation) can be identified on the Fig. 2.11 and is equal to 0.172 J/cm2. Again, this is the value ofthe absorbed fluence.

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Chapter 4

Experimental Fabrication of Bimetallic Micro-stripes on a Ceramic Substrate

4.1 Microfabrication by a direct laser ablation: process design and parameters

4.1.1 Beam trajectory patterns

The main purpose of this chapter is to present and discuss the results of the experimentalmicrofabrication of structures, resembling the key features of the embedded thermocouple sensor(section 1.2.3), and to describe the process and laser parameters used to achieve these results. Thesekey features in the sensor structure were considered to be narrow metal stripes isolated by a non-conductive ceramic gap. To produce a metallic stripe from a continuous coating, all the unnecessarymaterial had to be ablated through the entire depth, exposing the ceramic substrate. Essentially, thisis a micromachining by a laser ablation cutting (section 2.5.1). The laser beam, moving relatively tothe processed sample, produces a cut in width of the order of the spot size and depth approximatelygiven by the equation 2.17. A number of adjacent cuts, distanced by an order of the spot diameterfrom each other, will remove the required material volume with a sufficient number of processcycles. This basic idea of laser machining was used in order to design the laser beam trajectories ina graphic redactor of the laserDesk software. Two main types of the beam trajectory patterns, usedto produce an isolated narrow metal stripe, are shown on the Fig 4.1a,b:

Figure 4.1a,b: Beam trajectory patterns used in the experimental microfabrication

Δx parameter is the so called scanning interval, which, as already mentioned, has to be in the orderof the scanning beam spot size. The term “scanning” refers here to the used beam guidance system– the scanning galvanometer. The fast-moving laser beam, guided by the rotating galvanometermirrors, scans the part, producing the overlapping cuts merging into the removed volume. Hence,the later used term scanning velocity has the same meaning as a cutting velocity, the velocity of therelative movement between the beam and the work-piece. The scanning interval has to be small

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Δx Δx

60 μm

100 μm

100 μm

a) b)

enough to allow the cuts to merge into a continuous gap of the removed material, and has a similarmeaning to the pulse overlap in the beam scanning direction. The direction of the scanning isalways the same, i.e. from top to bottom along the lines on the Fig 4.1. When the beam reaches thebottom of the lines trajectory, the laser radiation is switched off by sending the appropriatecommand signal to the external AOM, and the beam position is changed to the start of the next lineby a “jump”, i.e. fast realignment of the galvanometer mirrors. The speed of this jump can be anorder of magnitude higher then the scanning velocity. Fig. 4.1a shows a beam trajectory pattern used as a first trial in the experimental processing, tocheck the possibility of fabricating a metallic stripe of width below 100 μm. With this resultsuccessfully achieved, the pattern shown on the Fig. 4.1b was used to produce a series of evennarrower stripes (in the order of 40 μm width, taking into account the spot radius) isolated fromeach other by a 100 μm wide gap. This was considered close to the actual resolution requirementsfor the embedded thermocouple sensor fabrication.

4.1.2 Laser radiation and processing parameters

The essential processing parameters for the pulsed laser material processing are the laser generationwavelength, the laser pulse duration, the pulse repetition rate and the average output power of thelaser. The output power at the given repetition rates defines the pulse energy and at the givenfocusing geometry (and the beam shape and quality) defines also the laser pulse fluence spatialdistribution. The central wavelength of the laser generation was defined by the active medium of thelaser system (Yb-doped silica) and equal to 1.03 μm. Converting the wavelength of infrared lasersto green or even UV spectrum is known to increase the absorbance of metals [36]. However, thesame effect could be achieved by simply increasing the laser output power, and even thoughreducing operation wavelength could increase the process energetic efficiency, it was not likely toreveal any fundamentally new effects in the micromachining. Therefore, the laser operationwavelength parameter was kept constant.The laser pulse duration parameter was controlled by the compressor delay stage. Its position wasadjusted to the optimal value for a given pulse picking rate. According to the quality test reports theoptimal adjustment of the delay stage position guarantees the pulse duration below 500 fs (section3.1.1) In fact, recalling the theoretical results presented in the section 2.1, once the duration of thepulse is definitely in the ultra-fast interaction regime (i.e. below the characteristic time of theelectron-phonon relaxation), its exact value becomes less important, since all the relevant processesof laser-matter interactions are defined by the material-specific time constants. The exact value ofthe pulse duration can be still important for the observation of the non-linear scattering effects(section 3.1.1). In this work however, these effects were trying to be avoided by staying at the lowfluence regime. Therefore, the pulse duration may be considered as merely being in a femtosecondtime scale, with the exact value being inessential for the micromaching process. Changing the pulse picking frequency of the laser system allowed a wide variation of the pulserepetition rate values – from several kHz to several MHz. The optimal value of this parameter, tostart the investigations with, was chosen to be 200 kHz. The choice is far from arbitrary and isaimed to maximize the productivity of the process, while avoiding any negative pulse-to-pulseinteraction effects. The first effect which should be considered, while choosing the repetition ratefor a laser micromachining process, is a subsequent pulse attenuation by the ablation products, e.g.plasma shielding. In the experimental research described in the section 2.3.2 the recovery of theinitial optical transmission was achieved 2-3 μs after the ablation. 200 kHz corresponds to a 5 μstime interval between subsequent pulses. Other sources also report this value of the repetition rate tobe the safe margin for the manifestation of plasma shielding effects [32,34]. It is still somewhatunclear to which extent the ablative pulse fluence influences the time of the optical transmissionrecovery, but it is natural to assume, that pulses with higher fluence produce more ablation products,

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which require more time to be removed from the ablation site. The effect of plasma shielding isclosely connected to the heat accumulation, since interaction of a subsequent pulse with the ablatedparticles forms a residual heat source close to the target surface. But even in absence of thisadditional heat source (e.g. in vacuum conditions), the heat accumulation effect can manifest itselfat high pulse repetition rates. As described in section 2.4.1, the onset of the heat accumulation ismore likely to be caused by increasing the pulse repetition rate then by increasing the pulse energy.Owning to this fact, it can be very convenient for the experimental process design to fix therepetition rate and vary the pulse energy (i.e. average power at a given repetition rate), in search forthe value satisfying the processing requirements. After the desired pulse energy (and the spatialdistribution of the fluence at the given focusing geometry) is defined, the pulse repetition rate canbe attempted to increase until the limit, restricted by the plasma shielding or heat accumulationeffects.From very obvious reasons, the higher pulse repetition rate is always advantageous, since it allowsto increase the processing speed and productivity (in absence of negative pulse-to-pulse interactioneffects). Being more specifically, the higher repetition rate allows to use a higher scanning speed,and therefore reduce the processing time, at the given pulse overlap parameter. As shown in thesection 2.5.2, the pulse overlap parameter together with the pulse repetition rate defines thecutting/scanning velocity at the given focusing geometry. With the laser repetition rate fixed at 200kHz, the required pulse overlap set to 75% and the beam waist radius assumed to be at least 10 μm,the equation 2.18 yields: vrel = 1 m/s. Since the scanning velocity defined by the equation 2.18 issubject to changes with a focusing geometry, the minimum possible value estimated from thelogarithmic fluence plots in 3.2 was used. With beam waist radius more then 10 μm, the pulseoverlap is expected to be higher then 75%. However, it should be noted that the beam spot size isnot always equal to the ablation diameter, and the non-uniform ablation rate effect can lead tosignificant variation of the drilled depth along the beam waist (section 2.2.2). The diameter ofsignificant crater depth can be much smaller then the beam waist diameter. This effect shouldbecome more pronounced with the increasing depth of the cut. It is also likely to influence theminimum required scanning interval, which would allow the adjacent cuts to effectively merge intoa single ablated volume. Higher pulse overlaps and smaller scanning intervals may be required tobreak the tendency of the tapering channels formation, and certain fluctuations in the cutting depthmay be inevitable, while processing with Gaussian-beam laser pulses.At this point, it may be convenient to summarize the laser and process parameters, which were keptconstant in the following experimental investigations:

Laser wavelength: λ=1.03μmPulse duration: τp<500 fsPulse repetition rate: f rep=200 kHzScanning velocity: vrel=1m /s

The remaining varying parameters will then be: the output power and the pulse energy of the laser,the fluence of the laser pulse and its spatial distribution, and the scanning interval. The method ofvarying the laser output power and pulse energy was described in the previous chapter: by settingthe transmission efficiency of the external AOM. To establish the typical scale of values, theexperiments were conducted with the average optical power incident on the target ranging from 0.5to 4 W, which corresponds to 2-20 μJ pulse energy. At the given pulse energy, the laser fluence and its spatial distribution could be varied by movingthe stage with the target test sample relatively to the focal plane, i.e. changing the focusinggeometry. Due to lack of imaging instrumentation in the experimental setup, this parameter couldnot be measured directly, but only estimated basing on the slopes of the logarithmic fit linesdiscussed in the previous chapter. Comparing the slopes of the logarithmic fits on the Fig. 3.3 and

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3.8, it may be concluded that changing the target position corresponding to the Fig. 3.3 to thecorresponding to the Fig. 3.8 can reduce the beam waist radius from 25 to 15 μm (±5 μmapproximately). The threshold fluence for the material ablation was found to be in the order of 0.1J/cm2. For the beam waist radius of 20 μm, the pulse with 0.1 J/cm2 peak fluence has the energy ofapproximately 1 μJ. Obviously, these pulses cannot be used for material processing, since theablation crater it produce will be infinitely small. In fact, the pulse energies with ablation rates highenough allowing a reasonable processing time, still retaining a good cut quality, were found to havethe energy of 10-15 μJ (at the focusing geometry of 15-25 μm beam waist radius). At this point, asimple visualization of the spatial fluence distribution may prove useful (Fig. 4.2).

Figure 4.2: The spatial distribution of the laser pulse fluence for different focusing geometries

It can be noted that since the ablation threshold fluence is in the order of 0.1 J/cm2, most of thedistribution on the Fig. 4.8 is well above the threshold, and the ablation crater should be in principlelarger then the beam spot size. However, again, the influence of the non-uniform ablation rate onthe crater morphology should not be underestimated. This effect leads to the situation, in which theouter regions of the crater stop to grow in depth after getting a certain inclination (section 2.2.2). Onthe other hand, the central part of the crater will deepen with each subsequent impulse due to theperpendicular incidence at the central tip, multiple reflections and energy redistribution due to non-linear beam distortions. To illustrate these arguments, the spatial energy distribution with w0 = 15 μm may be considered.With the ablation threshold around 0.1 J/cm2, the observed diameter of the crater (or the width ofthe line cut) should be approximately 40 μm. However, the region of 10 < r < 20 μm (for instance)will stop to grow in depth after several micrometers due to the effect of the oblique incidence. Atthe region r < 10 μm on the other hand (again, as a characteristic example), the depth of the craterwill increase with each subsequent pulse, rapidly growing approaching the beam axis. This will leadto the situation, in which the actual observed macroscopic diameter of the hole (or the width of thecut) will be approximately 20 μm, even though the signs of material ablation will be observed on amuch wider region. One can think about applications, in which the material damage in the vicinityof the cut is unacceptable. One of the possible strategies to minimize this effect may be in keepingthe ablation diameter significantly smaller then the beam waist. Indeed, if the ablation threshold onthe Fig. 4.2 would be for example 2 J/cm2, the fluence distribution along the ablation crater wouldbe more uniform, and it can be intuitively assumed that shallow region around macroscopic craterwould be much less pronounced. In practice however, this would require using low pulse energiesand/or very tight beam focusing (beyond the capabilities of the lens used in the experimental setup).

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r [ μm]

w0=15μm

w0=20μm

Epulse=14 μJφ[J /cm2

]

Fortunately, for the fabrication of thermocouple sensor some minor damage and depth variations atthe edges of the stripe can be acceptable. The more significant problem is that even at the region, in which the macroscopic depth growth isobserved, it is far from being uniform. The blind hole or cut will have a morphology of the taperingchannel with a narrow tip. In absence of the stop layer, whose ablation threshold exceeds the peakpulse fluence, this may lead to inability to produce a smooth flat surface after the removal of therequired volume depth. Here, the two spatial fluence distributions shown on the Fig. 4.2 may lead todifferent morphologies. For the distribution with w0 = 20 μm, the shallow region of the ablatedmaterial around the deep crater may be even more pronounced, since the fluence is higher at theside-wings. However, at the central region, where a macroscopic ablation depth is observed, thefluence distribution will be somewhat more uniform, comparing with the fluence profile at thetighter focusing. While this slightly defocused fluence profile is certainly likely to damage thematerial around the macroscopic cut, it may be more suitable to produce a more flat surface afterthe material removal. Obviously, due to a significantly lower peak fluence, the overall ablation ratewill be much slower then in the tighter focusing geometry (w0 = 15 μm). These arguments might be only illustrative and intuitive, however they are not merely theoreticalspeculations. The described behavior of the laser ablation cut morphology was experimentallyobserved during the attempts to achieve the reasonable quality of the cut and the uniformity of thematerial removal. The experimental evidence of the discussed effects, and the results achieved inthe experimental microfabrication of the embedded thermocouple elements, will be presented in thenext section. Clearly, these effects may be expected to become influential with the increasing depthof the channel/cut, i.e. in the micromachining of structures with high aspect ratio. For the givenprocessing task, it is the consequence of the relatively large coating thickness (similar to the beamspot size) that has to be removed. In the ablation of thin films or micromachinig with low depth ofmaterial removal, where the spot size exceeds by far the cutting depth, these effects are not likely toplay a significant role.

4.2 Morphological analysis of the experimental micromachining results atdifferent focusing geometries

4.2.1 Experimental microfabrication results at poor optical alignment conditions with alarge beam spot size

The results of the laser processing presented in this section, are performed under conditionsassociated with a poor optical alignment, characterized by logarithmic fluence plots on the Fig. 3.3,3.4 (section 3.2.1). At this focusing geometry the beam possesses a significantellipticity/astigmatism and the beam spot is quite large (for both elliptical axis). Although, the pooradjustment of the beam entry into the scanning galvanometer aperture was a significant drawback inthe experimental setup, which had to be (and was) corrected, the relatively large beam spot size wasfound to be quite acceptable for the particular microfabrication task. With a beam trajectory definedas a series of adjacent lines (Fig. 4.1a,b), the influence of the beam ellipticity on the processing wasonly marginal. The large elliptical axis “b” was coincident with the cutting direction. Therefore, itscontribution was mainly to an increase of the pulse overlap. With wb≈35μm the pulse overlapwould be slightly above 90% instead of the predefined 75%. In the direction perpendicular to thecut, the beam waist radius was smaller, around 25 μm approximately. With this beam waist size itwas possible to produce a metal stripe in width of several 10 μm between the two ablative cuttrajectories distanced from each other by 60 μm (Fig. 4.1b). The morphological comparison with thestripe produced at the tighter focusing geometry provided some insights on the role of the non-uniform ablation channel growth in the investigated micromachining process, already discussed in

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the previous section to some extent. The micro-stripe structures shown in this section were fabricated using pulse energiesapproximately equal to 12 μJ. With the focusing geometry characterized by a relatively large beamspot, increasing the pulse energy further increased the ablation diameter to values, unacceptable forthe required micromachining resolution. With lower pulse energies (~7 μJ) it was impossible toremove the metal coating layer and reach the ceramic substrate surface within a reasonableprocessing time. The possible explanation to this fact may be the decrease of the volumetricablation rate with the channel depth. Thus, the ablation diameter should be high enough to break thetendency of a tapering channel formation and allow to deepen the cut. The influence of this effect isof course dependent on the pulse overlap and the scanning interval. The scanning interval waschosen to be 20 μm, similar to the estimated beam waist radius. Fig. 4.3a shows an optical microscope image of a stripe fabricated from a continuous chromelcoating with the beam trajectory pattern shown on the Fig. 4.1a, with Δx = 20 μm. On the Fig. 4.3bthe focus of the image is shifted to the ceramic surface, exposed after the metallic layer ablation.The stripe has an average width of approximately 80 μm, while its base adjacent to a ceramicsubstrate surface is ~ 100 μm (the distance between adjacent beam trajectory lines on the Fig. 4.1a).Clearly, due to the effect of the non-uniform ablation rate, the stripe has a trapezoidal shape,widening to the bottom.

Figure 4.3a: ~ 80 μm micro-stripe fabricated by12 μJ pulses with a large beam spot

Fig. 4.4 shows the same stripe on a SEM image with a sample tilted by 19˚, to allow studying thesidewall morphology. It can be noted, that no signs of the adhesion violation between the metalliccoating and the ceramic substrate are present after the processing.To produce the stripe from a ~48 μm chromel coating, 300 processing cycles (runs of the beam overthe same trajectory) were necessary to completely remove the metallic layer and leave the ceramicsurface free of metal debris. This implies, that in a single run 160 nm of the coating material depthwas removed in average. However, since the ablation rate is non-uniform this estimation cannot bevery reliable. Firstly, the ablation rate (and therefore, depth of material removal in a single run) issubject to changes with a channel depth. Secondly, at final stages of the process several runs wererequired to remove the metal remains by low-fluence side-wings of the beam profile. These strips ofmetal leftovers tend to appear between adjacent beam trajectory lines and are again the consequenceof the non-uniform channel growth phenomena. Since the scanning interval is 20 μm, the overlapbetween adjacent cuts was not high enough to break this tendency completely. However, severalfinishing runs left a quite clean ceramic surface, as can be seen on the optical image on the Fig. 4.3b

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Figure 4.3b: same as 4.3a, but the focus shifted to a ceramic surface

and SEM image on the Fig. 4.4. It is very probably that during these runs some ablation of theceramic substrate occurred. If so, this ablation was very subtle and relatively uniform, as nogrooving cuts are observed on the ceramic surface.

Figure 4.4: SEM image of the ~ 80 μm micro-stripe fabricated by 12 μJ pulses with a large beamspot (sample tilted by 19 ˚)

To simulate the fabrication of narrower stripes required by the embedded thermocouplespecifications from the same ~48 μm chromel coating, beam trajectory shown on the Fig. 4.1b wasemployed with the same number of runs and pulse energy. The optical microscope image of theresulting series of narrow stripes can be seen on the Fig. 4.5a,b. Again, the trapezoidal shape of thestripe is evident. The average width of the stripe is approximately 35 μm, and the width of the base– 60 μm (the distance between adjacent line trajectories of the beam on the Fig. 4.1b). Themagnified SEM image of a single narrow stripe is shown on the Fig. 4.6a,b with and without the 19˚inclination.

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35.9 μm ~60 μm

Figure 4.5b: same as 4.5a, but the focus shifted to a ceramic surface

Figure 4.5a: ~35 μm micro-stripes fabricated by 11 μJ pulses with a large beam spot

Figure 4.6a: SEM image of the ~35 μm micro-stripe fabricated by 12 μJ pulses with a largebeam spot (sample tilted by 19 ˚)

The aforementioned “average width of the stripe” is the width of the quasi-trapezoidal structure as itperceived on optical microscope images. However, it is not the width of the stripe vertexundamaged by the laser radiation. Careful analysis of SEM images show, that the region at the topof the stripe with no signs of material ablation is much narrower: ~65 μm instead of 80, and ~25instead of 35. And yet, the so-called “average width” may have some physical meaning: it is thewidth beyond which the abrupt increase in depth is observed. This is consistent with a typical cutmorphology defined by the non-uniform channel growth effect discussed before: the shallow regionis abruptly replaced by a channel of macroscopic depth. Clearly, as steep as at may be, this abruptlydeepened cut sidewall is not perpendicular, but has a certain inclination. Hence, the macroscopictrapezoidal shape of the produced stripe. Examining closely the sidewall on the Fig. 4.6a, burr-like structures can be noticed, which remindthe signs of the re-solidified ejected melt. However as remembered, no significant melt formationwas observed in the multi-pulse drilling experiments (section 3.2.2). Alternatively, it might be theconsequence of the distorted elliptical spot, which performs the cutting. This assumption is based onthe later observed superior quality of the cut sidewalls, produced by a tightly focused beamcorrected for elliptical distortions.Summing up, the first fabrication experiment proved the possibility of producing a narrow metallicstripe from a continuous coating with a reasonable precision and no signs of violating the adhesionto a ceramic substrate. Improving the optical alignment and the focusing geometry, in order toeliminate the ellipticity of the beam spot and reduce the beam waist size, could potentially increasethe processing precision. But it also led to additional consequences, which had to be taken intoaccount, possibly requiring to alter some of the processing parameters.

4.2.2 Experimental microfabrication results at improved optical alignment conditions with abeam tightly focused on a sample surface

This section analyzes microfabrication results performed with an improved optical alignment andfocusing geometry, characterized by the logarithmic fluence plot on the Fig. 3.8. The beam spotfocused on a sample surface is no longer considered elliptical, and the beam waist radius is about~15 μm. The machined coating was now bimetallic, with ~48 μm thick chromel layer coated on topof ~22 μm thick alumel layer, coated on the ceramic substrate. It can be noted that the overall

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Figure 4.6b: same as 4.6a, but without 19˚ inclination

thickness of the metallic coating is now ~1.5 times higher. The energy of the laser pulse was now14 μJ. Other laser and process parameters apart from the slightly increased pulse energy and thefluence distribution, defined by the focusing geometry, remained unchanged. The beam trajectoryshown on the Fig. 4.1b was used to fabricate a series of narrow stripes, similar to the requiredbimetallic thermal junction. The removal of the bimetallic layer and exposure of the ceramic surfacerequired 300 beam runs over the same trajectory. Taking into account the increased thickness of thedouble layer coating, it may be concluded that the overall ablation rate has increased about 1.5times due to the tighter focusing and higher peak fluence.

Figure 4.7a: S eries of narrow micro-stripesfabricated by tightly focused 14 μJ pulses

Figure 4.8a: SEM image of the micro-stripeseries fabricated by tightly focused 14 μJpulses (sample tilted by 19 ˚)

Looking on the fig. 4.7b, with focus of the image shifted to a ceramic surface, the aforementionedtendency of metal leftovers, appearing at the center between the trajectories of the beam runs,become very evident. In contrast to the previous case, with a tight focusing geometry (and the same

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45.5 μm

Figure 4.7b: same as 4.7a, but the focus shifted to a ceramic surface

Figure 4.8b: same as 4.8a, but a single micro-stripe is zoomed

scanning interval of 20 μm) they become nearly impossible to remove by additional finishing runs.This is a very obvious consequence of using the same scanning interval with a smaller size beamspot. Examining the ceramic surface on SEM images, the strips of metal leftovers become even morepronounced. Moreover, there are signs of a significant damage the ceramic substrate, whichmanifests itself in a form of groove cuts along the beam run trajectory lines, where the laser fluencehas a peak. The increase of the fluence due to the tighter focusing had no doubt increased theamount of the ceramic ablation, comparing to the previous case. Even if the metal leftoversformation will be reduced, by reducing the scanning interval, the significant damage to the substratemay pose a considerable problem. If the application will require minimization of the substratedamage, reducing the pulse energy at the tighter focusing geometry should be considered. It willcome at expense of the ablation rate and the processing speed, however.From examining the SEM images it may be also evident, that the overall quality and precision ofthe microfabrication has significantly improved. The structure of the stripe looks much more“sharp” with clearly defined edges. While the trapezoidal shape still characterizes its macroscopicconfiguration, the region undamaged by a laser radiation is wider (~35 instead of ~25 μm), and theshallow region at the edge of the cut is much less pronounced (hence, the impression of welldefined edges of the structure). Moreover, the quality of the cut sidewall had significantlyimproved. It looks quite smooth and uniform with almost no signs of burrs, unlike in the previouscase, even though the peak intensity is rather high and some melt formation might be expected. Thiscan be considered a result of the low fluence finishing effect, observed and pointed out in deepholes drilling experiments (section 2.1.6). Even if some burrs are formed due to re-solidification ofthe melt, ejected in the ablation process, they are cleaned by the low fluence side-wings in severalfinishing runs. Of course, during these last runs the ceramic substrate is damaged. While the exactreasons may remain uncertain, the experimental fact is, that the quality of the cut sidewall for atightly focused beam had significantly improved. This can be attributed to the low fluence finishingeffects, and to the regular shape Gaussian spot, which might be more suited to produce a uniformablation cut, then a distorted spot with a significant ellipticity/astigmatism.

4.2.3 General patterns in the morphology of the fabricated metallic stripes

Summarizing, the general morphological pattern of the produced structures can be describedqualitatively. The so-called “stripe”, had to have a form of the metallic boss, affixed from thebottom to a ceramic bulk by the adhesion provided by the thermal spraying process. However, dueto the extensively discussed non-uniform ablation rate effects, this boss was not rectangular, butrather had a shape resembling a pyramid with a cut vertex (i.e. trapezoidal shape). Furthermore,between the steep sidewall of the cut and the undamaged vertex of the stripe, there was a shallowregion with low-pitching slope, where the ablation depth was only microscopic. The schematic cross-section of the described morphology is depicted on the fig. 4.9. Thecharacteristic dimensions of this trapezoidal shape (e.g. the slope of the sidewall, the width and thedepth of the shallow region) depends on the laser fluence spatial profile. In principle, it can beroughly predicted by the model proposed in [21], though there are many effects which aren't takeninto account (multiple reflections, non-linear beam distortion, melt formation). With high peak fluences due to the tight focusing geometry, the damage to the substrate close to thebeam axis and appearance of grooves on a ceramic surface may become inevitable. It may be alsoimportant to note again, that no signs of adhesion violation between the metallic coating and thesubstrate, or between alumel and chromel coatings were observed in the produced structures in anyof the experiments.

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Figure 4.9: Schematic transverse section of the fabricated micro-stripe

4.3 Measurable criteria for a metallic coating removal

Up to now, the criteria for a removal of the metallic layer was only visual. I.e. when the ceramicsurface was exposed and became visible with only minor metal leftovers, the layer considered to beremoved. This criteria is however very subjective and indistinct, and in case of tightly focused beampulses, with the beam radius lower then the scanning interval, it becomes also impractical (Fig.4.7b). A clear measurable criteria is needed to define the number of processing cycles required, untilthe metal coating can be considered removed. This criteria can be defined in terms of the electricalconductivity of the metallic layer – the layer is considered removed, when its conductivity drops tozero. Thus, the metal leftovers are not considered to form a layer, if they are isolated from oneanother and cannot conduct electricity. Clearly, the application requirement may be to clean theceramic surface from any metal debris for reducing the excessive weight on the part and/or betteradhesion of the protective ceramic coating, applied to the part after the fabrication of the sensor.However, the minimal requirement for a sensor functionality is, that its elements will be electricallyisolated from one another. The measurements of the electrical conductivity (or rather, the resistivity in Ω) were performedusing another beam trajectory pattern – the conductivity measurement rectangle (Fig. 4.10a). Onecontact of the voltmeter was placed at the center of the rectangle, and another – in the outer region.Working again with the tight focusing geometry, the scanning interval Δx was now set to be 10 μm. The conductivity measurement experiments were carried out in the tight focusing geometry withpulse energy 14 μJ, on the layers of alumel and chromel with approximate thickness of 47 and 33μm. The number of processing runs was gradually increased (with the resolution of 10 runs) untilthe conductivity measured by the voltmeter dropped to infinitely low values (the resistivitymeasured between the contacts raised above 200 MΩ). With both coating types the resistivityincrease was rather abrupt and occurred after 90-110 runs. The conductivity cutoff occurredabruptly for instance after 90 runs, while after 80 runs the measured resistivity was still negligible(~ 0.7 Ω). The changes in the number of runs for a zero conductivity, for measurement rectanglesproduced in different locations on the test-sample, were likely to be caused by the non-uniformity inthe thickness of the metallic coating, the consequence of the low precision of the thermal sprayingprocess. The measured conductivity dropped to zero suddenly, as soon as the mechanical cutoff

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AlumelAlumel

ChromelChromel

Ceramic substrateCeramic substrate

undamaged vertex

shallow region

steep sidewall

possible damage to the ceramic substrate

between the inner contact pad and outer metal surface occurred. This can be illustrated by themicroscope images of the rectangle gap after the number of runs insufficient for the conductivitycutoff, i.e. with the layer still behaving as a perfect conductor (Fig 4.11).

Figure 4.10a: The conductivity measurementrectangle beam trajectory pattern

Figure 4.11: Microscope images of the conductivity rectangle gap region failing to insulate theinner contact pad from the outer surface

On the images the metal leftovers, which close an electrical contact and allow electricityconduction, can be clearly seen. The morphology of these leftovers is quite interesting. These are nolonger strips extended in parallel to the beam trajectory lines. The decrease of the scanning intervalto 10 μm seemed to greatly reduce the tendency of their formation, so that they can be hardlyspotted on the processed ceramic surface. Instead, here small regions can be noticed, in which somemetal still left on the ceramic surface, covered with a mesh of tiny holes. The apparition of thismetal remains, while the rest of the layer was completely ablated, can be again attributed to the non-

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~2 mm

σσ

Δx

Figure 4.10b: A conductivity measurment rectangle on the optical microscope image produced in a ~47 μm thick alumel layer with 100 process runs

uniform layer thickness. However, the observed mesh of the holes can be the evidence that thegrowth of the ablation craters proceeds separately from each other, i.e. the pulse overlap is not highenough to break the tendency of the tapering channel formation. It is possible that along withdecreasing the scanning interval, the tighter focusing geometry also requires to decrease thescanning velocity to keep a high pulse overlap, breaking the tendency of separate growth of taperingchannels. However, it is crucial to check, weather an increase of the overlap won't lead to thereduction in the quality of the cut sidewalls.

4.4 Fabrication of a structure resembling a full functioning thermocouple sensor

At this point, the amount of the acquired experimental data already allowed the fabrication of aprototype structure, which would resemble a thermocouple in its functionality, i.e. allow themeasurement of a voltage on the thermal junction. This would require to produce an isolated contactpads in alumel and chromel layers, and a narrow stripe interconnecting between the pads throughthe bimetallic layer, acting as a thermal junction. The employed beam trajectory pattern is shown onthe fig. 4.12.

Figure 4.12: Beam trajectory pattern for the fabrication of the thermocouple resembling structure

The mesh of beam trajectory lines with a scanning interval of 10 μm removed the metal coating,leaving only the contact pad and the interconnecting narrow stripe. The identical (but verticallyflipped) beam trajectory was applied to the other type of metal coating. At the region of overlapbetween alumel and chromel coatings (bimetallic layer) the two beam trajectory pattern coincided.Thus, the bimetallic layer experienced twice more beam runs (i.e. processing cycles) then alumeland chromel layers individually. This experimental fabrication was conducted on a sample coatedby a ~47 μm alumel layer and a ~33 μm chromel layer (chromel on top). The conductionmeasurement experiments showed, that the number of runs required to remove these layer is similar(considering the inaccuracy of the thermal spraying coating process leading to the fluctuations inlayers thickness) and approximately equal to ~100 process cycles. This simple concept however, encountered some problems in realization. First experimentalattempts revealed that the number of beam runs required to remove the bimetallic layer is not equalto the sum of runs required to remove each of the layers. In fact, removal of the bimetallic layerrequired much less processing cycles. Thus, in the bimetallic region a significant ablation of theceramic substrate occurred around the interconnecting stripe. A decrease in thickness of the ceramic

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2 mm

60 μmΔx = 10 μm

the interconnecting stripe

Contact pad

coating, left after the bimetallic layer removal, could be recognized even with the naked eye. With120 process runs for each beam trajectory (i.e. 240 runs over the bimetallic layer) the ceramiccoating was completely ablated, and a super-alloy substrate was exposed. As noticed before, theimpulse energy of 14 μJ in conjunction with a tight focusing geometry is able to cause a significantablation of the ceramic. Reducing the number of runs to 100 allowed to prevent the completeremoval of the ceramic coating, and the resulting structure was already quite similar to the desiredthermocouple prototype (Fig. 4.13a,b).A large amount of metal leftovers after the alumel layer removal (which had a higher thickness of~47 μm) may be noted (Fig. 4.13a). In a closer look, these are similar to the leftovers on the Fig.4.11, i.e. small regions of metal remnants, covered with a mesh of holes. Here, the former boundaryof the bimetallic layer can be clearly recognized by absence of the metal debris. The removal of thethinner chromel layer (~33 μm) had left the ceramic surface rather clean (Fig. 4.13b). However,here the border can be recognized by small region, where the ceramic coating was completelyablated. This is the consequence of the adjustment failure, due to which the double run of 200cycles was performed on the single layer of chromel coating. Obviously, the proper adjustment ofthe beam trajectories relative to the sample require means for defining the beam starting positionwith a microscopic precision. Otherwise, these excessive overlaps due to the positioning inaccuracywill become inevitable.Unfortunately, the measured conductivity between the pads was zero. Looking closely on the Fig.4.13b, the reason can be identified straightaway – the micro-cracks appear in the interconnectingstripe at the connection to the chromel contact and at the former border between the chromel andbimetallic layer. There is also a crack at the alumel side border line, which can be seen on fig. 4.13a.Summarizing, the cracks appear at the abrupt changes in thickness of the ceramic substrate and atthe connection to one of the contact pad. The latter can be explained by a scanning galvanometeroperation peculiarity – the scanning beam experiences a certain lag after the end of the beamtrajectory. Before the laser is switched off by the external AOM, the spot at the end of the beamtrajectory is hit by excessive number of pulses. Thus, the depth of ceramic layer ablation isincreased locally at both sides of the contact, and the condition for the crack appearance mayassumed to be the same, as in two previous cases: the abrupt changes in thickness of the ceramicsubstrate in the vicinity of the micro-stripe. The investigation of the cause of these cracks and their exact mechanism is beyond the scope of thiswork. However, it is important to note, that the investigation of this effect is crucial for thefabrication of the embedded thermocouple sensors. At this point, it was the only obstacle for thefabrication of the full functioning thermocouple prototype. As the abrupt change of the substratethickness is clearly involved, this effect might be a mechanical fracture of a sort. Thermal stressesare also likely to play a role in the process. Nevertheless, one example of the functionalthermocouple prototype was actually managed to be produced. The reason were the excessive metalleftovers in the particular microfabrication experiment. These leftovers closed the contact betweenthe pads bypassing the cracks in the interconnecting stripe. While it was only an accidental benefitof the normally undesired effect, which cannot be exploited as a technological method, it providedthe opportunity to test an alumel-chromel thermal junction fabricated in the particular way, namely– by an ablative micromachining of a thick bimetallic coating. Placing the voltmeter contacts atalumel and chromel pads, the resistance of the interconnecting stripe was measured as ~30 Ω, andthe voltage measured between the pads (i.e. voltage of thermal junction) was ~14 mV. Assuming theSeebeck coefficient of the K-type chromel-alumel junction to be ~40 μV/K, the theoretic thermalvoltage of the junction at 298 K will be ~12 mV. Thus, the fabricated thermal junction seems to beintact, and might perform accordingly in the thermocouple circuit (Fig. 1.4). To provide a typicalscale of the production speed, – the fabrication of the described thermocouple prototype (with 100processing cycles) required around 20 minutes of the laser operation.

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Figure 4.13a: Alumel contact pad and the inter-contact bimetallic stripe

Figure 4.13b: Chromel contact pad and the inter-contact bimetallic stripe

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Figure 4. 14a,b,c: Micro-cracks in the interconnecting stripe(a) – micro-crack at the former chromel-bimetallic border, (b) - micro-crack at the former alumel-

bimetallic border, (c) – micro-crack at the connection to the chromel contact pad

4.5 Conclusions and prospect

The conclusions of the microfabrication experiments described in this chapter, and of the work ingeneral, can be divided to those having a technological and a fundamental significance. The mostimportant technological conclusion is that the possibility of the fabrication of the embeddedthermocouple sensor from a thermal sprayed alumel-chromel bimetallic coating by an ultra-shortpulse laser processing was proven. Thus, the process sequence proposed in the chapter 1 is inprinciple technologically feasible. The vacuum deposition process can be excluded, replaced by thethermal spraying in conjunction with an ultra-fast micromachining, producing a thick filmthermocouple sensor structure. The problem of micro-crack in the interconnecting bimetallic stripe,appearing around the thermal junction, seems to be principally solvable, with the appropriate furtherinvestigations carried out. No fundamental obstacles to the proposed processing method werefound in the experimental investigations. The adhesion of the metallic coating to the ceramicsubstrate was not violated by the surrounding material ablation. And finally, the prototype thermaljunction was fabricated, which gave a realistic value of the contact potential voltage measured atroom temperature. The fundamental conclusion is the pointed out out-most importance of the non-uniform ablationrate and channel growth effects in the microfabrication of structures with a high aspect ratio (e.g.thick film patterning) by Gaussian-shape beam pulses. The final shape of the structure, its precisionand maximum resolution, are defined by these effects. The model proposed in [21], after a certainmodifications extending it from the drilling to an ablation cutting, can prove to be very useful inpredicting the shape of the produced structure, as a function of the spatial fluence distribution andthe pulse overlap. Other effects not included in the model might have to be taken into accountphenomenologically: multiple reflections from the sidewalls, non-linear distortion of the beamprofile due to interactions with an ambient gas and the melt formation and ejection. Additionalimportant effect, strongly influencing the precision of microfabrication, is likely to be the lowfluence finishing due to the interaction of the side-wings of Gaussian-beam pulses with the cutsidewalls. It is very probable that this effect, originally pointed out for the deep hole drilling [19], isresponsible for the superior precision and cut quality in a wide range of ultra-fast micromachiningprocesses.

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b)a) c)

To give a prospect to the work, a number of vectors for the future investigations can be defined, tofurther improve the micorfabrication process, possibly achieving a production feasibility.

• The problem of micro-crack appearing in narrow metallic stripe in the vicinity of the abruptchanges in the substrate thickness should be investigated and solved. If the excessiveablation of the surrounding ceramic substrate layer will found to be the cause (a thermalshock wave in a ceramic substrate or similar phenomena) it has to be avoided or minimized.A different number of processing cycles should be used for each of the layer types – alumel,chromel and bimetallic, in order to remove the metallic layer with only minor substratedamage. An accurate positioning with a microscopic precision should be used to define thestarting position of the beam, to avoid the excessive number of processing runs on thethinner layers.

• Adding to the previous concept of the minimal substrate damage, the use of a large spotdefocused beam at the final process stage may be investigated. Reducing the pulse energy atthe same time, might allow to lower the peak fluence below the ablation threshold of theceramic. Thus, a concept of stop-layer may be realized. The defocused low fluence beamwill be used to clean the substrate surface of metal leftovers and debris. However, this is notlikely to be achieved in practice. The ceramic surface is very rough and cleaning it from themetal particles, implanted in a thermal spraying process, will require some minor ceramicablation. The low fluence defocused beam might be adjusted to effectively “polish” theceramic surface in a gentle ablation regime. The advantage of using a partially defocusedbeam is the minimization of the non-uniform ablation rate effects. It may allow to leave thesubstrate surface clean and flat after the metallic layer removal, if the application demands it(e.g. for the good adhesion of the applied second ceramic protective coating). Developingthese concepts will require an in-depth investigation of the ultra-short pulse ablation of theYSZ ceramic coating, at least similar to the done for the Ni-alloy coating in the chapter 3.

• For the effective removal of the Ni-alloy layers on the other hand, moderate energy tightlyfocused pulses were found to be suitable. The beam waist radius of ~15 μm and the scanninginterval 10 μm seems to give satisfactory results. The option of increasing the pulse overlap(by reducing the scanning velocity) should be investigated, since this might counteract thenon-uniform crater growth effect and tendency for a tapering channel formation, greatlyincreasing the material removal rate. The quality of the cut sidewalls should be carefullymonitored however, for any signs of possible negative effects of the high overlap (e.g. heataccumulation).

• The use of higher pulse energies should be investigated, since it will provide an increasedablation rate and processing speed. Nonlinear effects due to interactions with an ambient air,which will undoubtedly be intensified in this case, might have a positive effect of flatteningthe beam profile and redistributing the energy more uniformly (beam filamentation). Withhigher pulse energies the amount of melt formation will increase, and so the role of the meltejection in the ablation process, but the low fluence finishing effect is still likely to keep thehigh cut quality. There might be a contradiction however with a previous concept of lowsubstrate damage, since finishing runs of high energy pulses are going to damage thesubstrate very significantly. It may be needed to return to the concept of the two stageprocess: the high energy pulses will rapidly produce a rough preform, and low energy (andfluence) pulses will provide the low fluence finishing, improving the cut quality andcleaning the ceramic surface from metal debris.

• Flat-top beam profile might be utilized to practically eliminate the nonuniform ablation rateeffects. The use of a uniform beam profile will greatly simplify many of the processingconsiderations. The shape of the fabricated stripe will much less resemble a trapezoid, andits transverse cross-section can become very close to rectangular. While this concept seems

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to promise many advantages, it is unclear whether they are worth the increase in acomplexity of the optical system required. The use of optical elements, transforming thebeam profile into a uniform, will also limit the maximum pulse energy, and therefore theablation rate.

• At the final stage of the process optimization, with the optimum energy and fluenceparameters already fixed, a power scaling by increasing the pulse repetition rate may beemployed. The scanning velocity will also have to be increased to keep the same pulseoverlap, and the productivity of the process will be enhanced. This strategy to achieve ahigher processing speed is limited by negative effects of pulse-to-pulse interactions – heataccumulation and particle shielding. The critical heat accumulation will manifest itself in areduction of the cut quality and overall processing precision due to a massive surfacemelting. The influence of the particle shielding can be recognized by a decrease in theablation rate, due to subsequent pulses attenuation. The repetition rate, at which one of thiseffect will be observed, may be considered a critical limiting value. It should be kept inmind, that this critical repetition rate value strongly depends on the pulse energy. Therefore,the optimization of both parameters for the maximum processing speed, still retaining a highprocessing quality, may become iterative. If the strategy of the two-stage processing, withhigh fluence main part and low fluence finishing, will be incorporated, the maximumallowed repetition rate at the finishing stage will obviously be significantly higher.

• The fabrication of the thermocouple prototype on a small and macroscopically flat testsample is an idealized case. The actual industrial process will have to produce thedelineation of the sensor structure on the particular machine part (e.g. turbine blade), with acurved surface and dimensions exceeding the field of view of the scanning optical system.Therefore, an industrial setup for thermocouple microfabrication has to include an auto-focusing system, and a motor-driven motion stage. The development of such system and theappropriate interfaces for its different components, can be a complex investigation task byitself.

• Finally, a functional thermocouple sensor prototype (including reference junctions) shouldbe fabricated and characterized for a thermal response. A calibration function of the device,as well as the response time and sensitivity, have to be measured as a function of differentgeometric parameters of the thermal junction stripe (e.g. the average width, trapezoidality ofthe shape, sidewalls quality, etc...). This will help to understand, what can be the desiredlaser pulse fluence profile, in order to produce the shape of the structure, conforming to theapplication requirements.

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