Materials Science and Engineering A426(2006)
250–262
Nd:YAG laser butt welding of AA6013using silicon and
magnesium containing?ller powders
Reinhold Braun?
German Aerospace Center,Institute of Materials Research,K¨o ln,Germany
Received30November2005;received in revised form31March2006;accepted11April2006
Abstract
Aluminium alloy6013sheet was butt welded using a3kW Nd:Y AG laser and different?ller powders.Two kinds of?ller metals were used: gas atomised powders of the aluminium alloys AlMg5,AlSi12,AlSi12Mg5and AlSi10Mg,as well as mixtures of powders of the elements Al and Si and the binary alloys Al–5Mg,Al–5Zr,Al–5Cr and Al–10Mn.Microstructure,hardness,tensile properties and corrosion behaviour of the welds were investigated in the as-welded T4and T6conditions and after a post-weld heat treatment to the T6temper.The weld region exhibited a dendritic cellular structure in the fusion zone and a partially melted zone adjacent to the fusion boundaries.The hardness of the fusion zone depended upon the?ller metals used.Post-weld arti?cial aging to the T6temper improved the hardness,being associated with precipitation of strengthening phases,as con?rmed by transmission electron microscopy.Joint ef?ciencies achieved for post-weld heat treated joints ranged from 80to90%.Strengths of welds in the as-welded T6condition were lower due to the softened fusion zone and heat affected zone,being between 70and75%of the ultimate tensile strength of the base alloy6013–T6.Optimum tensile properties were obtained with joints made with the?ller powder AlSi12.Reasonable tensile properties were achieved for welds made with mixtures of elemental and binary alloy powders,but they did not reach those of joints made with aluminium alloy powders.The dispersoid forming elements Zr,Cr and Mn added to a mixed Al–7Si powder did not prove bene?cial effects on weld quality.When exposed to an intermittent acidi?ed salt spray fog,joints in the as-welded T4and post-weld heat treated T6conditions exhibited corrosion behaviour similar to that of the base sheet in the tempers T4and T6.As-welded6013–T4joints were susceptible to stress corrosion cracking when immersed in an aqueous solution of0.6M NaCl+0.06M NaHCO3.Sensitivity to environmentally assisted cracking was associated with grain boundary precipitates in the heat affected zone.
?2006Elsevier B.V.All rights reserved.
Keywords:Laser beam welding;Aluminium alloys;Filler powders;Mechanical properties;Corrosion behaviour
1.Introduction
Requirements of weight savings,improvement of fuel ef?-ciency and reduction of atmospheric pollution have increased the demand of the transportation industries for lightweight struc-tures[1].To meet legislative regulations and market needs and to retain competitiveness of their products,manufacturers have been forced to explore advanced materials and joining tech-niques.The use of aluminium in automotive applications is growing owing to advantageous properties of aluminium alloys, including high stiffness to weight ratio,good formability,good corrosion resistance and recycling potential[2].This has driven developments in5000and6000series aluminium alloys to expand their range of applications[2,3].Non heat-treatable5000?Tel.:+4922036012457;fax:+492203696480.
E-mail address:reinhold.braun@dlr.de.series alloys are favoured for interior structural components, while heat-treatable6000series alloys are used for outer skin panels due to their higher strength after paint baking and free-dom from L¨u ders band problems[4].The occurrence of L¨u ders lines in components during the forming process deteriorates the high surface quality required for exterior skin panels.In Europe alloy6016is the main automotive skin alloy due to its high formability,whereas alloy6111is preferred in North America because of its high?nal strength and dent resistance[5].
Laser beam welding is an attractive joining technique for alu-minium alloys,being widely used in industrial production[6,7]. The advantages are high welding speed,low distortion,manu-facturing?exibility and ease of automation.Welds are produced using CO2and Nd:Y AG laser beams.Whereas CO2lasers have the potential of welding material with high thickness or using high welding speeds due to the high beam power,the latter type of laser allows optical?bres to guide the laser to the focusing optics[8].Aluminium alloys absorb the laser more ef?ciently
0921-5093/$–see front matter?2006Elsevier B.V.All rights reserved. doi:10.1016/j.msea.2006.04.033
R.Braun/Materials Science and Engineering A426(2006)250–262251
as the laser wavelength decreases[1].The Nd:Y AG laser with a characteristic wavelength of1.06?m provides better coupling with aluminium than the CO2laser,which has a characteris-tic wavelength of10.6?m.Furthermore,the absorption of the laser beam increases drastically when a keyhole is formed due to multiple re?ections of the beam in the keyhole.Dif?culties occurring with laser beam welding of aluminium alloys are hot cracking,porosity,loss of alloying elements and grain boundary melting in the heat affected zone[1,6,9].
In the automotive industry,the use of tailor-welded blanks made from aluminium offers considerable weight and cost reduction[1,10].To achieve the targets regarding operational cost,manufacturing cost and structural weight,airframe manu-factures have adopted integral construction in the manufacturing process,replacing the conventional built-up structure[11,12]. For the aircraft Airbus A318,mechanically fastened stiffeners of lower fuselage shells have been substituted by stringers attached to the skin using CO2laser welding.This joining process is also employed during the manufacture of skin/stringer panels of the lower shells of the aircraft Airbus A380.The aluminium alloys used for lower fuselage applications are the weldable Al–Mg–Si–Cu alloys6013and6056[13].
The weldability of aluminium alloys depends on the type of alloy and is strongly affected by the chemical composition of the weld zone[1,14,15].Al–Mg–Si alloys are susceptible to hot cracking,exhibiting maximum susceptibility when containing about1wt.%Mg2Si.Solidi?cation cracking in high strength aluminium alloys can be avoided by modifying the weld pool chemistry using appropriate?ller metals and dilution ratios. Aluminium?ller alloys containing excess silicon and magne-sium are recommended for6000series alloys[15].Laser welds of the alloys6063and6082were produced without cracking using the?ller wires Al–5Si and Al–7Si with magnesium and scandium additions[16,17].The automotive alloy6111was also autogenously butt welded with full penetration and min-imum surface discontinuities[10].Sheet materials of the alloys 6013and6056considered for fuselage applications were welded using CO2and Nd:Y AG lasers,respectively,and Al–12Si?ller wires[18,19].While the application of?ller wire seems to sta-bilise the laser beam welding process[20,21],the use of?ller powder feeding equipment might allow a greater?exibility with regard to welding position and welding process interruption in comparison to restrictions associated with tolerances required for wire positioning accuracy.Furthermore,powder production processes provide an expanded chemical composition range of ?ller metals compared to that of wires produced by ingot met-allurgy.In the present work,alloy6013sheet was butt welded using a Nd:YAG laser and?ller powders containing silicon and magnesium.Microstructure,mechanical properties and corro-sion behaviour of the joints in different heat treatment conditions were investigated.
2.Experimental
The material used was a1.6mm thick sheet of the alloy6013 received in the naturally aged T4condition.To achieve peak-strength,part of the sheet was arti?cially aged at191?C for4h.Butt welds of alloy6013in the tempers T4and T6were produced in a shielded atmosphere using a3kW continuous wave Nd:Y AG laser.The shielding gas was helium at a?ow rate of10l/min. The focal length of the employed lens was150mm.The laser beam was normal to the sheet and focused1.0–1.5mm above the sheet surface.The specimens being400mm×100mm in size were tightly fastened(no gap)using a rigid clamping equipment. The edges of the specimens were milled and https://www.doczj.com/doc/a214978851.html,ser power and welding speed ranged from2300to2600W and from 3to4m/min,respectively.The weld direction was parallel to the rolling direction of the sheet.Filler materials used were gas atomised powders of the aluminium alloys AlMg5(4.5–5.5% Mg),AlSi12(10.5–13.5%Si),AlSi12Mg5(10.5–13.5%Si, 4.5–5.5%Mg)and AlSi10Mg(9.0–11.0%Si,0.2–0.5%Mg). The chemical composition ranges of the alloying elements given in wt.%are presented in brackets.The size of the powder parti-cles ranged from45to150?m.The?ller materials were added to the weld pool using a gas nozzle and argon transport gas at a?ow rate of2.5l/min.The feed rate was10g/min.The gas nozzle was positioned at a distance of10–12mm from the weld pool at an angle of55?with the specimen.The powder particle ?ow direction was opposite to the weld direction.
For welding a particular aluminium alloy,the most suitable ?ller metal has to be selected to obtain optimised weld proper-ties.Mixing powders of elements and binary alloys provide an economical way to adjust?ller metal composition.Therefore, mixtures of different powders were evaluated as?ller metals instead of highly alloyed aluminium powders.In a?rst step, powder mixtures of the pure elements aluminium,silicon and magnesium were manufactured.Proportions of each component were weighed and then mixed using a TURBULA?shaker-mixer for1–2h.Powder particles of the elements Al and Si were approximately spherical,whereas those of Mg were irreg-ularly shaped.Visual inspection indicated that a homogeneous mixture was not obtained for mixtures containing high amount of magnesium(≥4.5wt.%),probably related to the irregular shape of the magnesium particles.Therefore,these powders were additionally mixed in a planetary ball mill with agate balls for45–60min.Filler materials of the chemical composition Al–12Si,Al–0.9Mg–1Si,Al–12Si–5Mg and Al–4.5Mg–0.7Mn were manufactured(compositions given in wt.%).In a second approach,magnesium was added using gas atomised powder of the binary alloy Al–5Mg.Filler metals Al–7Si,Al–7Si–3Mg and Al–7Si–4.5Mg were produced by mixing powders of the elements aluminium and silicon and the binary alloy Al–5Mg. To study the effect of the dispersoid forming alloying elements zirconium,chromium and manganese on the weld properties, small amounts of powders of the binary alloys Al–5Zr,Al–5Cr and Al–10Mn were added to an Al–7Si mixture.The size of the powder particles was again in the range between45and150?m. Because of the spherical shape of the blend components,pow-ders in the second approach were mixed only in a shaker-mixer for1–1.5h.Chemical compositions of?ller metals produced in this way were:Al–7Si–0.2Zr,Al–7Si–0.2Cr,Al–7Si–0.4Mn, Al–7Si–0.2Zr–0.2Mn and Al–7Si–0.2Cr–0.2Mn.
Optical microscopy and scanning electron microscopy (SEM)were used to perform microstructrual examination.
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Compositional analysis of the fusion zone was carried out by energy dispersive X-ray(EDX)spectroscopy at an accelerating voltage of20kV.Point and area analyses were conducted. The microstructure of the weld region was also studied by transmission electron microscopy(TEM).TEM specimens were prepared by jet polishing using a solution of nitric acid (33vol.%)and methanol(67vol.%)at?20?C and12V.
Vickers hardness measurements were carried out on etched metallographic sections of the welds along half thickness line with a load of4.91N(0.5kp).Joints were investigated in the as-welded T4and T6conditions.Welds of6013–T4sheet were also studied after a post-weld heat treatment to the T6temper (4h/191?C).Tensile tests were conducted using transverse ori-ented tensile specimens with a gauge length of50mm and a width of12mm.The weld zone was in the centre of the speci-mens,and the thickness of the weld region was reduced to that of the base sheet by milling the weld bead.Duplicate specimens were tested.
The corrosion behaviour was studied using modi?ed salt spray tests according to ASTM standard G85,Annex A2.Pan-els of50mm×100mm in size were exposed to an intermittent acidi?ed salt spray fog for2weeks.Before exposure,the panels were cleaned in an aqueous5%NaOH solution at room tem-perature for5min,dipped in nitric acid for30s and rinsed in deionised water.The stress corrosion cracking(SCC)behaviour was investigated performing constant load tests under perma-nent immersion conditions in an aerated aqueous solution of 0.6M NaCl+0.06M NaHCO3.Flat tensile specimens of width 6.5mm and gauge length15mm were used,with loading axis in the transverse direction.The fusion zone in the centre of the tensile specimens was machined,reducing the thickness to that of the base sheet.Before testing,the SCC specimens were ultra-sonically cleaned in alcohol and degreased in acetone vapour. The failure criterion was fracture.The maximum exposure time was30days.Duplicate specimens were tested for each stress level.
3.Results and discussion
3.1.Joints welded using aluminium alloy?ller powders
Full penetration joints were produced by laser beam weld-ing using aluminium alloy?ller powders.The welds exhibited a good weld shape with approximately round beads(Fig.1). Visual inspection as well as metallographic examinations did not reveal hot cracks.The use of highly alloyed?ller powders avoided weld pool compositions close to the critical Mg2Si con-tent at which maximum solidi?cation cracking susceptibility occurs.Porosity was the main weld defect observed in the fusion zone.Fig.2shows metallographic sections of the fusion zone. The fusion zone exhibited a?ne cellular dendritic solidi?cation structure with formation of mainly equiaxed grains in the weld centre.This dendritic structure was generally found with laser welded6000series aluminium alloys[1].Fine equiaxed grains reduce the solidi?cation cracking susceptibility and improve the mechanical properties of the welds[14].Low melting point seg-regates are distributed over a larger grain boundary area in
?ne Fig.1.Macrographs of6013alloy joints laser beam welded using aluminium alloy?ller powders:(a)AlSi12and(b)AlSi12Mg5.
grained material,and strains are accommodated more uniformly by equiaxed grains.A partially melted zone was observed adja-cent to the fusion boundaries(Fig.2b).Its width was typically two to three grains.Partially melted zones occur as a result of heating up the area surrounding the fusion zone to tempera-tures between the eutectic temperature and the liquidus of the alloy[14].During weld thermal cycles,grain boundaries con-taining eutectic phases locally melt in the region adjacent to the fusion boundaries.For the ternary Al–Mg–Si system,the eutec-tic alloy,liquid→?-Al+Mg2Si+Si,was reported to solidify at the eutectic temperature of559?C with the composition 4.6wt.%Mg,13.2wt.%Si and82.2wt.%Al[22].During fusion welding,aluminium alloys can be susceptible to liquation crack-ing being related to liquation of low melting point components in the partially melted zone.In laser welded aluminium alloy joints, however,liquation cracking has rarely been observed due to the low heat input and small heat affected zone associated with this welding process[1].The eutectic phase re-solidi?es on cool-ing.Adjacent to the grain boundaries,a precipitate free zone is observed caused by solute rejection on freezing[23].The precipitate free zones occur preferentially on the grain sides ori-ented towards the fusion zone(Fig.2b).These features observed in the heat affected zone of laser beam welded Al–Mg–Si–Cu sheet adjacent to the fusion boundary are referred to as oriented dispersoid-free zones[24].Their formation is probably caused by strain or diffusion induced grain boundary migration.
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253
Fig.2.Optical micrographs of6013alloy joint welded using aluminium alloy ?ller powder AlSi12Mg5,showing(a)the dendritic structure of the fusion zone and(b)the partially melted zone.
Results of EDX analysis of the weld solidi?cation structure are given in Table1.To determine the composition of the fusion zone of joints made with different?ller powders,an area of 50?m×40?m was scanned(Fig.3).The average concentra-tions of magnesium and silicon in the fusion zone increased with increasing content of these alloying elements in the?ller metals.Magnesium and silicon were also enriched in the?-Al dendrite cores,being available for precipitation hardening.The interdendritic regions consisting of several phases contained sig-ni?cant amounts of the alloying elements silicon,magnesium, copper and manganese as well as the impurity element iron.An increased iron concentration was measured in the fusion zone compared to that in the base sheet,probably due to high amount of impurities in the?ller powders.The multicomponent inter-metallic compounds in the interdendritic regions were primary and second phase particles as well as eutectic phases.Possible phases that can form in Al–Mg–Si alloys encompass Al12(Fe, Mn)3Si,Mg2Si,Al5Cu2Mg8Si6,Al2Cu,Al8FeMg3Si6[25]and the?-FeSiAl5phase being very hard and brittle with a detri-mental effect on mechanical properties[26,27].Similar EDX Table1
Results of EDX analyses of chemical composition of the base material and the fusion zone of6013alloy joints welded using different aluminium alloy?ller powders(concentration in wt.%)
Al Mg Si Cu Mn Fe Base alloy96.950.870.61 1.180.300.10 Filler powder AlMg5
Fusion zone96.02 1.470.81 1.080.340.28 Dendrite97.52 1.490.530.47
Interdendritic region90.18 1.05 1.38 5.430.51 1.45 Filler powder AlSi12
Fusion zone95.310.67 2.670.730.380.25 Dendrite97.330.54 1.730.39
Interdendritic region84.18 1.368.82 3.07 1.00 1.56 Filler powder AlSi12Mg5
Fusion zone95.21 1.10 2.10 1.010.340.23 Dendrite97.670.790.950.60
Interdendritic region83.87 1.678.04 4.710.69 1.02 Filler powder AlSi10Mg
Fusion zone95.32 1.00 1.81 1.170.350.36 Dendrite97.640.80 1.040.53
Interdendritic region88.03 1.05 4.25 5.040.55 1.09 analyses results were found in the fusion zone of T-shaped welds of6056sheet laser welded using an Al–12Si?ller wire[19].
Hardness values determined in the fusion zone of alloy6013 joints made with different?ller powders are presented in Fig.4. For welds in the as-welded T4and T6conditions fairly simi-lar values of the Vickers hardness HV0.5were measured.The lowest hardness values were observed for joints made with the ?ller powder AlMg5.The hardness increased using silicon-rich ?ller powders.The highest value was determined for welds made with the?ller powder AlSi12Mg5.This might be related to an increasing amount of brittle eutectic phases and constituent par-ticles when using the latter highly alloyed?ller metal.Similar results were found for butt and T-shaped welds made with?ller powders rich in silicon and magnesium[28,29].The hardness
in Fig.3.Scanning electron micrograph of the fusion zone of6013alloy joint welded using aluminium alloy?ller powder AlSi12Mg5.
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250–262
Fig.4.Vickers hardness HV0.5in the fusion zone of 6013alloy joints made with different aluminium alloy ?ller powders in the as-welded T4and T6conditions and after a post-weld heat treatment (pwht)to the T6temper.
the fusion zone increased after a post-weld heat treatment owing to precipitation strengthening,as con?rmed by transmission electron microscopy.TEM micrographs revealed needle-shaped precipitates (Fig.5).These precipitates corresponded to the ? -phase,which mainly contributes to the strength in 6000series aluminium alloys [30].In copper containing Al–Mg–Si alloys,the Q -phase provides additional strength [31].Both ? -and Q -phases are coherent with the matrix,aligned along the 100 direction of aluminium.For alloy 6111in the peak-aged temper,approximately 20%of the total volume fraction of precipitates was found to be Q
.
Fig.5.Transmission electron micrograph showing strengthening precipitates in the fusion zone of 6013alloy joint made with aluminium alloy ?ller powder AlSi10Mg in the post-weld heat treated T6temper.
Fig.6shows hardness pro?les across the welds made with different aluminium alloy ?ller powders,taken in the midsection of the joints.In the as-welded T4and T6conditions,joints made with AlMg5exhibited a hardness drop in the fusion zone below the T4level of the parent sheet.The degradation might be associ-ated with vapourisation loss and dilution signi?cantly reducing the concentration of the alloying element silicon in the fusion zone,and,therefore,precipitation of hardening phases was not suf?cient to recover the full strength level of the naturally aged microstructure.For welds made with silicon-rich ?ller metals,the hardness in the fusion zone was similar to or higher than that of the base alloy 6013–T4.Hardness values measured close to the fusion boundaries did not reveal a heat affected zone in as-welded 6013–T4joints.In accordance with results presented in Fig.4,the hardness pro?les con?rmed the increase in strength of the fusion zone resulting from post-weld arti?cial aging to the T6temper.The deep drop in the weld made with AlSi12Mg5was probably related to porosity.(Although the shape of the indent did not indicate penetration in a pore,subsurface pores reduce the dent resistance resulting in an increased indentation depth and thus in a lower hardness value.As shown in Fig.1b,the fusion zone of welds made with AlSi12Mg5contained many pores.)For joints in the as-welded T6condition,the hardness decreased in the heat affected zone from the peak strength level to the T4level.A narrow hardness plateau was observed adja-cent to the fusion boundary,indicating that the precipitates of the peak-aged parent sheet dissolved due to the heat input of welding.Subsequently,this region hardened by natural aging.Similar hardness pro?les were observed in laser beam welded joints of 6013–T6extrusions [32].
Transverse tensile properties of joints made with different aluminium alloy ?ller powders are presented in Table 2.For as-welded 6013–T4welds,joint ef?ciencies (ratio of the ultimate tensile strength of the weld to that of the base alloy)were in the range from 80to 90%.The lowest values were measured for joints welded using ?ller metals with high magnesium content.These welds exhibited also reduced ductility.After a post-weld heat treatment to the T6temper,the ultimate tensile strength increased for all joints.The improvement was low for welds made with https://www.doczj.com/doc/a214978851.html,ing silicon-rich ?ller powders,strength increase was quite equal,indicating that a high amount of silicon was necessary to provide signi?cant precipitation of strengthen-ing phases.Joint ef?ciencies approached 90%for post-weld heat treated welds made with silicon-rich ?ller powders.In the as-welded T6condition,joints exhibited ultimate tensile strengths being rather similar to those in the as-welded T4condition.This was in accordance with the hardness drop in the fusion zone and heat affected zone (Fig.6).Joint ef?ciencies between 70and 75%were obtained referred to the strength of the base alloy 6013–T6.The fracture elongation measured using a 50mm gauge length was reduced for joints in the as-welded T6condi-tion,because strain was concentrated in the weld region due to the lower strength there.
SEM fractographs of 6013welds revealed a predominantly dimple like fracture mode (Fig.7).Failure occurred in the fusion zone,often close to the fusion boundaries.On the fracture surfaces,pores and secondary cracks were observed.Porosity
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Fig.6.Hardness pro?les across6013alloy welds in the as-welded T4and T6conditions and after a post-weld heat treatment to the T6temper.The joints were welded using aluminium alloy?ller powders(a)AlMg5,(b)AlSi12,(c)AlSi12Mg5and(d)AlSi10Mg.
included macropores with a diameter ranging from200to 600?m and micropores being50?m and less in size.Macro-porosity was probably associated with process instability and keyhole collapse[9,33].These large pores were particularly found in welds of Al–Mg alloys[33]and in joints made with?ller alloys containing a high amount of magnesium[29,34].Micro-porosity was caused by hydrogen rejected from the solid metal on solidifying[15].Potential hydrogen sources were moisture contaminants and hydrated oxides on the surface of the parent sheet and?ller powders.In the present work,no special surface treatments were employed,either to the sheet or to the?ller met-als.Owing to their high speci?c surface area,?ller powders can supply a large amount of hydrated oxides.Therefore,hydrogen induced microporosity typically occurred in joints laser beam
Table2
Transverse tensile properties of1.6mm thick sheet of base alloy6013in the tempers T4and T6and6013alloy joints made with different aluminium alloy?ller powders in the as-welded T4and T6conditions and after a post-weld heat treatment(pwht)to the T6temper
Temper Filler powder Ultimate tensile strength(MPa)Fracture elongation(%)Joint ef?ciency(%) Base material
T434528.9
T639712.3
Laser beam welded joints
T4as-welded AlMg5282 5.182
T6pwht AlMg53100.378
T6as-welded AlMg52760.770
T4as-welded AlSi123168.992
T6pwht AlSi123620.691
T6as-welded AlSi123000.476
T4as-welded AlSi12Mg5285 4.783
T6pwht AlSi12Mg53420.286
T6as-welded AlSi12Mg53050.577
T4as-welded AlSi10Mg3017.787
T6pwht AlSi10Mg361 1.091
T6as-welded AlSi10Mg2880.473
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Fig.7.Scanning electron fractographs of6013alloy joints welded using alu-minium alloy?ller powders(a)AlSi10Mg and(b)AlSi12Mg5,showing pores and secondary cracks,respectively.
welded using?ller powders[20,28,35].Proper surface prepara-tion of the parent material and protection of the?ller metals from humid environments can reduce pore formation in welds[16,33]. Secondary cracks observed on the fracture surfaces were prob-ably associated with hot cracking(Fig.7b).These small cracks might be caused by liquation of low melting point eutectics along the grain boundaries in the dendritic solidi?cation structure of the fusion zone[36]
.Fig.8.Scanning electron fractograph of6013alloy joint welded using mixed ?ller powder Al–12Si–5Mg,showing microporosity.
3.2.Joints made with mixed?ller powders
Transverse tensile properties of joints made with aluminium alloy powders and corresponding mixtures of elemental pow-ders are summarised in Table3(values of joints welded using the?ller alloys AlMgSi1and AlMg4.5Mn are taken from[35]). Ultimate tensile strength and fracture elongation were higher for welds made with aluminium alloy powders.Whereas reduction in tensile properties was slight for joints made with the elemental powder mixture Al–12Si,the degradation was signi?cant using powder mixtures containing magnesium.This might be associ-ated with insuf?cient mixing due to the irregular shape of the magnesium particles and high levels of oxide and hydroxide contamination on the surface of elemental Mg powder.On the fracture surfaces of tensile specimens taken from joints made with powder mixtures containing Mg particles,numerous small pores were observed(Fig.8).This porosity probably caused premature fracture resulting in degraded ductility and strength. Similarly,porosity was found to reduce fatigue strength to about one-half of that of the parent material[37].
Results of the EDX analysis of welds made with?ller pow-der mixtures are listed in Table4.Gas atomised powders of the binary alloys Al–5Mg,Al–5Zr,Al–5Cr and Al–10Mn were
Table3
Transverse tensile properties of6013alloy joints made with different aluminium alloy?ller powders and powder mixtures in the as-welded T4condition and after a post-weld heat treatment(pwht)to the T6temper
Filler powder Type T4as-welded T6pwht
Ultimate tensile strength(MPa)Fracture
elongation(%)
Ultimate tensile
strength(MPa)
Fracture
elongation(%)
AlSi12Alloy powder3168.93620.6 Mixed powder2977.13450.5
AlSi12Mg5Alloy powder285 4.73420.2 Mixed powder250 2.22650.1
AlMgSi1Alloy powder291 6.23430.3 Mixed powder247 2.42690.3 AlMg4.5Mn Alloy powder259 3.22820.2 Mixed powder2080.82460.2
R.Braun/Materials Science and Engineering A426(2006)250–262257 Table4
Results of EDX analyses of chemical composition of the fusion zone of6013alloy joints welded using different?ller powder mixtures(concentration in wt.%)
Al Mg Si Cu Mn Fe Filler powder Al–7Si
Fusion zone96.350.68 1.440.930.330.28 Interdendritic region89.060.80 3.88 4.070.63 1.58 Filler powder Al–7Si–3Mg
Fusion zone95.69 1.13 1.29 1.230.360.31 Interdendritic region87.79 1.24 2.96 5.900.64 1.47 Filler powder Al–7Si–4.5Mg
Fusion zone96.29 1.070.93 1.060.360.28 Interdendritic region88.07 1.52 3.36 4.080.76 2.21 Filler powder Al–7Si–0.2Zr
Fusion zone96.180.70 1.31 1.250.330.23 Interdendritic region87.58 1.08 4.66 3.650.88 2.15 Filler powder Al–7Si–0.2Cr
Fusion zone96.390.71 1.27 1.040.330.26 Interdendritic region89.640.72 3.68 4.180.57 1.21 Filler powder Al–7Si–0.4Mn
Fusion zone96.450.69 1.17 1.030.400.25 Interdendritic region87.680.88 3.68 4.20 1.12 2.44 Filler powder Al–7Si–0.2Zr–0.2Mn
Fusion zone96.360.74 1.27 1.010.370.24 Interdendritic region88.450.83 4.33 3.230.96 2.21 Filler powder Al–7Si–0.2Cr–0.2Mn
Fusion zone96.520.67 1.330.900.330.25 Interdendritic region87.650.83 4.53 3.630.98 2.38
added to an elemental powder mixture Al–7Si.The reduced silicon content of7wt.%corresponds to that of the standard aluminium?ller alloy AA4010.With increasing magnesium content in the?ller metal,an increased magnesium concentra-tion was measured in the fusion zone and in the interdendritic regions of the solidi?cation structure.Zirconium and chromium were not detected either by area analysis of the fusion zone or by point analysis of interdendritic phases,probably due to their limited(0.2wt.%)additions.Concentration of manganese was rather similar in all powder mixtures,being slightly increased in the interdentritic regions of joints welded using?ller metals containing manganese.
Fig.9shows hardness pro?les across alloy6013welds made with the?ller powder mixture Al–7Si–0.2Zr–0.2Mn in different heat treatments.For joints in the as-welded T4and T6condi-tions,hardness in the fusion zone dropped below the T4level of the base alloy.In the post-weld heat treated T6temper,the increase in strength was moderate,approaching the level of 6013–T4parent material.The pro?les were similar to those mea-sured for joints made with the?ller alloy AlMg5(Fig.6a).The hardness drop in the fusion zone was observed for all joints made with mixed Al–7Si powder and powder mixtures Al–7Si containing Zr,Cr and Mn.Hardness values measured in the fusion zone are listed in Table5for joints made with different mixed?ller powders in the as-welded T4condition and after a post-weld heat treatment to the T6temper.Hardness did not depend on the?ller metals used,except for welds made with Al–7Si–4.5Mg.For the latter welds,higher hardness was mea-sured in the T4and T6tempers.Again,this might be associated with an increased amount of brittle intermetallic and eutectic phases when using highly alloyed?ller powders.The addition of7wt.%Si was to small to increase the strength in the natu-rally aged fusion zone to the level of the6013–T4base alloy. However,suf?cient amounts of the alloying elements Mg and Si were present in solid solution to cause precipitation hardening after a post-weld heat treatment,as indicated by the increase in strength of joints in the post-weld heat treated T6
temper. Fig.9.Hardness pro?les across6013alloy joints made with mixed?ller powder Al–7Si–0.2Zr–0.2Mn in the as-welded T4and T6conditions and after a post-weld heat treatment to the T6temper.
258R.Braun/Materials Science and Engineering A426(2006)250–262
Table5
Vickers hardness in the fusion zone and transverse tensile properties of6013alloy joints made with different?ller powder mixtures in the as-welded T4condition and after a post-weld heat treatment(pwht)to the T6temper
Filler powder Temper Vickers hardness(HV0.5)Ultimate tensile strength(MPa)Fracture elongation(%) Al–7Si T4as-welded81±1.6269 3.8
Al–7Si T6pwht103±3.93150.2
Al–7Si–3Mg T4as-welded83±5.3273 4.6
Al–7Si–3Mg T6pwht109±9.03240.2
Al–7Si–4.5Mg T4as-welded89±8.9269 3.7
Al–7Si–4.5Mg T6pwht114±10.73160.2
Al–7Si–0.2Zr T4as-welded82±3.2275 3.9
Al–7Si–0.2Zr T6pwht104±4.33120.2
Al–7Si–0.2Cr T4as-welded82±3.3273 3.9
Al–7Si–0.2Cr T6pwht106±2.43210.2
Al–7Si–0.4Mn T4as-welded83±2.4275 4.0
Al–7Si–0.4Mn T6pwht102±5.43180.2
Al–7Si–0.2Zr–0.2Mn T4as-welded82±3.0279 4.1
Al–7Si–0.2Zr–0.2Mn T6pwht103±2.53200.1
Al–7Si–0.2Cr–0.2Mn T4as-welded82±3.3256 2.7
Al–7Si–0.2Cr–0.2Mn T6pwht103±4.73190.2
Ultimate tensile strength and fracture elongation of joints in the as-welded T4and post-weld heat treated T6conditions did not depend on composition of the mixed?ller powders used(Table5).Tensile properties of joints made with the ele-mental powder mixture Al–7Si were lower than those obtained for joints welded using the powder mixture Al–12Si(Table3).
A decrease of silicon content in binary Al–Si?ller metals below12wt.%did not improve strength and ductility,neither did an increased amount of silicon,as found for joints made with the aluminium alloy?ller powders Al–20Si and Al–40Si [28].Optimum joint ef?ciencies were achieved using the?ller metal AlSi12.For joints welded using the mixed?ller powders Al–12Si and Al–7Si,the increase of strength after a post-weld heat treatment to the T6temper was similar being about50MPa (Tables3and5).Therefore,the higher tensile properties of joints made with Al–12Si were not associated with enhanced precip-itation of hardening phases,but were probably related to the solidi?cation structure of the fusion zone containing a higher amount of silicon.
In accordance with results found for joints made with the?ller alloys AlSi12and AlSi12Mg5(Table2),magnesium added to the powder mixture Al–7Si did not reveal any bene?cial effect on strength.The improvement in strength of post-weld heat treated joints was about50MPa,being similar to that for welds made with mixed Al–Si powders(Table5).Thus,a higher Mg concen-tration in the fusion zone did not cause more copious hardening precipitates,but primarily resulted in Mg-rich phases in the inter-dendritic regions.The dispersoid forming elements zirconium, chromium and manganese did not exert an in?uence on tensile properties of welds.However,it has to be noted that dispersoids like Al3Zr,Al12Mg2Cr or Al12Mn2Si,which retard recrystalli-sation and grain growth in wrought aluminium products,form during homogenisation annealing of the ingots[38].Due to the high cooling rates during laser beam welding,these dispersoids were unlikely to be present in the dendritic solidi?cation struc-ture of the fusion zone of joints made with?ller powder mixtures containing small amounts of the binary alloys Al–5Zr,Al–5Cr and Al–10Mn.
Fractographic examinations revealed considerable macro-porosity on the fracture surfaces of tensile specimens taken from welds made with mixtures of elemental and binary alloy powders(Fig.10a).As already mentioned,macropores
were Fig.10.Scanning electron fractographs of6013alloy joints welded using mixed ?ller powders(a)Al–7Si–0.2Zr and(b)Al–7Si–0.2Cr–0.2Mn,showing macro-porosity and secondary cracks,respectively.
R.Braun /Materials Science and Engineering A 426(2006)250–262259
probably related to process instabilities and keyhole collapse [9,33].Thus,the use of powder mixtures caused a less stable laser welding process in comparison to feeding aluminium alloy ?ller powders,probably related to local inhomogeneities in the composition of the weld pool.The fracture surfaces of joints made with ?ller metals containing Al–5Mg powder particles did not exhibit pronounced microporosity,on the contrary to joints welded using powder mixtures with additions of elemental magnesium (Fig.8).Therefore,if Mg-rich powder mixtures are considered as appropriate ?ller materials,spherical particles of a binary Al–Mg alloy should be added rather than elemental Mg powder.Many secondary cracks were found on the fracture sur-faces of joints made with powder mixtures containing zirconium,chromium and manganese,especially when binary alloy Al–5Cr particles were mixed with the ?ller material (Fig.10b).Addi-tions of the dispersoid forming transition metals might increase hot tearing susceptibility,in particular,when chromium contain-ing powder was added.However,the mechanism of formation of hot tears promoted by these small additions of Al–5Cr,Al–5Zr and Al–10Mn is not understood and dif?cult to explain,because chromium and zirconium were not detected in the fusion zone by EDX analysis.
Titanium and zirconium are often added to the ?ller met-als as grain re?ners,because these elements can promote an equiaxed grain structure in welds [14,15].Similarly,scandium additions exceeding a critical level of ～0.4wt.%resulted in a highly re?ned uniform grain structure in the fusion zone of 7000series aluminium alloys [39].Small primary Al 3Sc parti-cles were formed in the liquid acting as ef?cient nucleants for the formation of new grains.The weldability of the Al–2.2Li–2.7Cu alloy being highly susceptible to hot tearing was enhanced by alloying additions of titanium and zirconium [40].The improve-ment of weldability was associated with re?nement of the grain structure and alteration of the shape and distribution of the eutectic phase in the weld.For high strength Al–Zn–Mg alu-
minium alloys,transition metal alloying elements were found to have an effect on solidi?cation cracking susceptibility [41].Whereas sensitivity increased with increasing amount of cop-per and chromium,welds containing manganese and zirconium were less susceptible.To improve the solidi?cation cracking resistance of Al–Zn–Mg alloys,optimum additions of Mn and Zr were found to be 0.68–0.72and 0.12–0.16wt.%,respectively.3.3.Corrosion behaviour of 6013joints
Fig.11shows the appearance of panels of the base material and joints made with the ?ller powder AlSi10Mg after 2weeks of exposure to an intermittent acidi?ed salt spray fog (ASTM G85,Annex 2).According to visual inspection,base alloy 6013was susceptible to pitting in both tempers T4and T6.For the peak-aged sheet,a more uniform corrosion attack was observed.The surface appearance of panels of joints in the as-welded T4and post-weld heat treated T6conditions was similar to that of the base material in the different tempers.Metallographic sections of base alloy panels revealed pitting for the naturally aged sheet.The maximum depth of the predominantly horizontal pits was 150?m.In the T6temper,the base alloy was susceptible to intergranular corrosion and pitting with a maximum depth of attack of 235?m.Fig.12shows metallographic sections of panels of as-welded 6013–T4joints after 2weeks of cyclic salt spray testing.The prevailing corrosion attack was pitting.In the heat affected zone,intergranular corrosion was also observed at a distance of 650–1000?m from the fusion boundary (Fig.12b).Similar to base alloy 6013–T6,panels of joints post-weld heat treated to the T6temper exhibited intergranular corrosion and pitting (Fig.13).Weld beads suffered pitting.An aggravation of the corrosion attack in the weld region was not observed.The difference in chemical composition between weld bead and parent material did not cause severity of corrosion forced by galvanic
coupling.
Fig.11.Appearance of panels of 6013base alloy in the tempers T4and T6and 6013alloy joints made with aluminium alloy ?ller powder AlSi10Mg in as-welded T4condition and post-weld heat treated (pwht)T6temper.The panels 50mm ×100mm in size were exposed to an intermittent acidi?ed salt spray fog for 2weeks.
260R.Braun /Materials Science and Engineering A 426(2006)
250–262
Fig.12.Metallographic sections of 6013alloy joint made with aluminium alloy ?ller powder AlSi12Mg5in the as-welded T4condition,showing (a)pitting and (b)intergranular corrosion in the heat affected zone.The panel was exposed to an intermittent acidi?ed salt spray fog for 2weeks.
Panels of joints in the as-welded T6condition,which were exposed to an intermittent salt spray fog,exhibited a ～2.5mm wide zone adjacent to the fusion boundaries being not cor-roded [28].These corrosion free zones were also observed for as-welded 6013–T6joints which were immersed in an aque-ous chloride–peroxide solution [42].The microstructure of the heat affected zone of joints in the as-welded T6condition cor-responded mainly to a naturally aged alloy,as indicated by hardness pro?les (Fig.6).Owing to the difference in corrosion potential,the region adjacent to the fusion boundary was prob-ably cathodically protected by the adjoining peak-aged parent material exhibiting a more active corrosion potential.
However,
Fig.13.Metallographic section of 6013alloy joint made with aluminium alloy ?ller powder AlSi12Mg5in the post-weld heat treated T6temper,showing inter-granular corrosion and pitting.The panel was exposed to an intermittent acidi?ed salt spray fog for 2weeks.Table 6
Time-to-failure data of 6013alloy joints made with different aluminium alloy ?ller powders in as-welded T4and T6conditions and after post-weld heat treat-ment (pwht)to the T6temper Filler powder Temper Initial stress (MPa)Time-to-failure (h)AlMg5T6as-welded 1702×>720AlSi12
T4as-welded 100110,229T6pwht
2002×>720T6as-welded 1702×>720AlSi12Mg5
T4as-welded 100100,235,420T6pwht
2002×>720T6as-welded 1702×>720AlSi10Mg
T4as-welded 100109,110T6pwht
2002×>720T6as-welded
170
2×>720
Transverse ?at tensile specimens were continuously immersed in an aqueous solution of 0.6M NaCl +0.06M NaHCO 3under constant load conditions.
corrosion free zones were not observed with as-welded 6013–T6coupons,which were alternately immersed in 3.5%NaCl solu-tion [28].Cathodic protection was not ef?cacious when the galvanic coupling was interrupted during specimen drying.No signi?cant effect of the ?ller metal on the corrosion performance of alloy 6013welds was found.Cyclic acidi?ed salt spray test-ing indicated similar corrosion behaviour for panels of joints welded using different aluminium alloy powders and powder mixtures.
Time-to-failure data of 6013joints are listed in Table 6,obtained from constant load tests using an aqueous chloride–bicarbonate solution.Flat tensile specimens of joints in the as-welded T4condition failed within 10days of exposure at an applied stress of 100MPa in transverse direction.Failure occurred in the heat affected zone ～1mm away from the fusion boundary,regardless of the ?ller powder used.Fractographic examination revealed an intergranular fracture (Fig.14).No fail-ure was observed for joints in the post-weld heat treated T6and as-welded T6conditions at applied stresses of 200and 170MPa,respectively.Because of reduced strength of the fusion zone
and
Fig.14.Scanning electron fractograph of 6013alloy joint made with aluminium alloy ?ller powder AlSi10Mg in the as-welded T4condition,showing intergran-ular stress corrosion cracking.The specimen failed after 109h of exposure to an aqueous chloride–bicarbonate solution at an applied stress of 100MPa.
R.Braun/Materials Science and Engineering A426(2006)250–262
261
Fig.15.Transmission electron micrograph of the heat affected zone of6013 alloy joint in the as-welded T4condition,showing grain boundary precipitates. heat affected zone,SCC specimens of joints in the as-welded T6 condition were tested at a lower load level.At higher stresses, joints failed due to severe pitting and intergranular corrosion near the fusion boundaries,as found with6013joints welded using the?ller alloy AlSi12[28].Base alloy6013in the tempers T4 and T6was immune to stress corrosion cracking.Flat tensile specimens of naturally aged and peak-aged sheet did not fail at applied transverse stresses of170and270MPa,respectively, when immersed in an aqueous chloride–bicarbonate solution for 30days.
Transmission electron microscopy revealed grain boundary precipitation in the heat affected zone of6013joint in the as-welded T4condition(Fig.15).As proved by EDX analysis in the transmission electron microscope,the precipitates were enriched with silicon and magnesium,probably being the Mg2Si phase.These grain boundary particles were not observed in nat-urally aged parent material.Their formation was induced by the heat input during laser beam welding.The precipitates pro-moted intergranular stress corrosion cracking.When panels of welds in the as-welded T4condition were exposed to an acidi?ed salt spray fog,the heat affected zone was sensitive to intergran-ular corrosion,whereas naturally aged parent sheet exhibited only pitting(Fig.12).Thus,SCC failure occurring in the heat affected zone of as-welded6013–T4joints might be caused by stress assisted intergranular corrosion.The SCC sensitivity was eliminated in the arti?cially aged microstructure probably due to coarsening of the precipitates during post-weld heat treating to the T6temper.
4.Conclusions
1.Butt welds of6013alloy sheet were produced using
a Nd:YAG laser and different aluminium?ller powders
enriched with silicon and magnesium.
2.The main weld defect was porosity in the fusion zone.Hot
cracking was only indicated by secondary cracks observed on the fracture surfaces of tensile specimens.
3.Concentrations of the alloying elements silicon and mag-
nesium being in solid solution were suf?cient to cause precipitation strengthening in the fusion zone during a post-weld heat treatment.Joint ef?ciencies approached90%for joints post-weld heat treated to the T6temper.
4.Alloy AlSi12was found to be the most appropriate?ller
powder to weld alloy6013sheet.Optimum tensile proper-ties were obtained for welds made with this?ller metal. 5.Reasonable strength values were measured for joints welded
using mixtures of?ller powders.However,joint ef?ciencies did not approach those achieved for welds made with alu-minium alloy?ller metals,primarily related to the addition of elemental magnesium powder to the mixtures.
6.Joints welded using mixed Al–12Si powder exhibited higher
tensile properties than those made with mixed Al–7Si pow-der.The optimum amount of silicon in Al–Si?ller powders was found to be near the eutectic composition.
7.Strengths of joints made with mixed Al–7Si powder did
not improve when magnesium or the dispersoid forming elements zirconium,chromium and manganese were added to the powder mixtures.
8.According to visual inspection,panels of6013welds
exposed to an intermittent acidi?ed salt spray fog suffered pitting.Corrosion attack was similar to that observed for the base material in the tempers T4and T6.
9.Metallographic examinations of panels revealed intergran-
ular corrosion and pitting for joints post-weld heat treated to the T6temper.With welds in the as-welded T4condi-tion,pitting was mainly observed;the heat affected zone, however,was also sensitive to intergranular corrosion. 10.As-welded6013–T4joints were found to be susceptible to
stress corrosion cracking in the heat affected zone when exposed to an aqueous solution of0.6M NaCl+0.06M NaHCO3.The SCC sensitivity was associated with grain boundary precipitation,as con?rmed by transmission elec-tron microscopy.
Acknowledgment
The author would like to thank Mr.G.Roth of DLR,Insti-tute of Technical Physics,Stuttgart,for carrying out laser beam welding.
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CY-1530F-500W 光纤激光切割 技 术 方 案 武汉楚域光电科技有限责任公司 2016.06
前言 光纤激光切割机是利用光纤激光发生器作为光源的激光切割机。 光纤激光器是国际上新发展的一种光纤激光器输出高能量密度的激光束，并聚集在工件表面上，使工件上被超细焦点光斑照射的区域瞬间熔化和气化，通过数控机械系统移动光斑照射位置而实现自动切割。 光纤激光器是国际上新发展的一种新型光纤激光器输出高能量密度的激光束，并聚集在工件表面上，使工件上被超细焦点光斑照射的区域瞬间熔化和气化，通过数控机械系统移动光斑照射位置而实现自动切割。同体积庞大的气体激光器和固体激光器相比具有明显的优势，已逐渐发展成为高精度激光加工、激光雷达系统、空间技术、激光医学等领域中的重要候选者。 激光切割是现今人们所掌握的各种切割技术中最好的切割方法，激光切割的优势在于: 热变形小、切割精度高、噪声小、无污染、易于实现自动切割、虽然初期投资大(劣势)，但加工成本比机械加工要少50% 。激光切割作为一种先进的制造技术，具有应用范围广、工艺灵活、加工精度高、质量好、生产过程清洁以及便于实现自动化、柔性化、智能化和提高产品质量、劳动生产率等优点。 光纤激光器是近几年激光领域里极其关注的热点，在加工领域光纤激光器有迅速替代传统的YAG、C02激光器的趋势。人们普遍认为,中功率光纤激光器将是第三代最先进的工业加工激光器。光纤激光器具有许多独特的优点：光束质量好;体积小,重量轻,免维护;风冷却简单易操件;运行成本低,可在工业环境下使用;寿命长、加工精度高、速度快;电能转化效率高,可以实现智能化、自动化、柔性化操作等。从整个激光技术的发展来看，光纤激光代表了激光的发展方向和趋势，其在工业、国防等领域有着重要的应用前景。
44瓦超高功率808 nm半导体激光器设计与制作 仇伯仓，胡海，何晋国 深圳清华大学研究院 深圳瑞波光电子有限公司 1. 引言 半导体激光器采用III-V化合物为其有源介质，通常通过电注入，在有源区通过电子与空穴复合将注入的电能量转换为光子能量。与固态或气体激光相比，半导体激光具有十分显著的特点：1）能量转换效率高，比如典型的808 nm高功率激光的最高电光转换效率可以高达65%以上 [1]，与之成为鲜明对照的是，CO2气体激光的能量转换效率仅有10%，而采用传统灯光泵浦的固态激光的能量转换效率更低, 只有1%左右；2）体积小。一个出射功率超过10 W 的半导体激光芯片尺寸大约为0.3 mm3, 而一台固态激光更有可能占据实验室的整整一张工作台；3）可靠性高，平均寿命估计可以长达数十万小时[2]；4）价格低廉。半导体激光也同样遵从集成电路工业中的摩尔定律，即性能指标随时间以指数上升的趋势改善，而价格则随时间以指数形式下降。正是因为半导体激光的上述优点，使其愈来愈广泛地应用到国计民生的各个方面，诸如工业应用、信息技术、激光显示、激光医疗以及科学研究与国防应用。随着激光芯片性能的不断提高与其价格的持续下降，以808 nm 以及9xx nm为代表的高功率激光器件已经成为激光加工系统的最核心的关键部件。高功率激光芯片有若干重要技术指标，包括能量转换效率以及器件运行可靠性等。器件的能量转换效率主要取决于芯片的外延结构与器件结构设计，而运行可靠性主要与芯片的腔面处理工艺有关。本文首先简要综述高功率激光的设计思想以及腔面处理方法，随后展示深圳清华大学研究院和深圳瑞波光电子有限公司在研发808nm高功率单管激光芯片方面所取得的主要进展。 2.高功率激光结构设计 图1. 半导体激光外延结构示意图
光纤激光切割原理 7轴联动工业机器人光纤激光切割机与五轴机床CO2激光切割机对比 三维切割系统的技术优势： 1.因为采用了业内最高精度的史陶比尔机械手，本体较轻，切割速度快，在小弧度的精细切割和大边的高速切割方面具有明显优势，实际切割速度可以达到18米/分钟而无抖动，综合加工效率是其他品牌机械手组合的两倍，性价比高，还可以节约一组的耗材和人工，后期可以少追加设备也能满足产能要求。还可24小时持续工作。一次性投入相对较少，在一个很短的折旧期内(两班8小时工
作制)，史陶比尔机器人激光解决方案就可回收投资。同时能耗少，体积小，维护需求低。 2.切割精度高。采用史陶比尔专利齿轮减速系统JCS和JCM，独一无二的驱动技术，确保了无可匹敌的轨迹控制精度和速度。即使是要求极高的小圆，或复杂立体几.何图形的加工，也可精确和快速完成，从而提升您的产品品质。系统重复定位精度高达±0.05M，完全可以满足钣金件行业的精度需求。可切割直径小至2MM的小圆，切割效果圆滑美观，目测无形变和毛刺。 3.切割幅面大，实际死角小。选配臂长2.01米的机械手，除了实现直径达3米的半球形三维加工区域外，还可实现较大的二维平面切割，配合我公司配套生产的可移动工作台2.5mX5m(2m的运动行程)，可实现2mX5m的二维平面切割。 4. 根据实际需要选配离线编程软件，可读取UG,SOLIDWORK等三维作图软件导出的vda,igs,x_t,sldprt,prt,stp,ipt,par等格式的数模，修改后直接生成切割轨迹，代替人工示教，简单易用。 5. 工业控制理念，模块化设计，全系统的防护等级为IP55,机械手防护等级更是高达IP65，系统集成度高，故障少，抗冲击振动，抗灰尘，无须光学调整或维护，真正适合于工业加工领域的应用用于恶劣的激光环境。结构坚固，动态性更佳。而其他同类产品为简单集成，设备的稳定性较差。 6.系统的工艺性和易用性较好。简单而功能强大的史陶比尔激光专用标准软件LasMAN基于Windows操作系统，用户界面简单友好，集成了机器人运动控制、激光控制、数据处理和产品管理等功能。友好的人机界面，模块化的设计，使得操作者仅需经过简单的培训即可达到系统产能最大化，同时也易于集成。这就大大降低了对操作工人的要求，降低了对工人的管理难度。 性能指标: 激光功率: 200W/300W/400W/500W/1000(根据工件材质和料厚可选) 激光波长:1070NM 工作区域:半径2米的半球形工作区域(选配半径2米的机械手) 切割速度:0-18米/分钟(根据功率大小和工件材质与厚度可调) 供电电源:三相交流380V 用电功率: <8KW(根据选配激光器功率大小而定) 冷却方式:风冷/水冷(根据选配激光器功率大小而定) 切割头焦距：5-7英寸(根据工件厚度可选)
固体激光器原理-固体激光器 固体激光器发展历程 固体激光器发展历程 固体激光器用固体激光材料作为工作物质的激光器。1960年，梅曼发明的红宝石激光器就是固体激光器，也是世界上第一台激光器。固体激光器一般由激光工作物质、激励源、聚光腔、谐振腔反射镜和电源等部分构成。 这类激光器所采用的固体工作物质，是把具有能产生受激发射作用的金属离子掺入晶体而制成的。在固体中能产生受激发射作用的金属离子主要有三类：(1)过渡金属离子；(2)大多数镧系金属离子；(3)锕系金属离子。这些掺杂到固体基质中的金属离子的主要特点是：
具有比较宽的有效吸收光谱带，深圳市星鸿艺激光科技有限公司专业生产激光打标机，激光焊接机,深圳激光打标机,东莞激光打标机比较高的荧光效率，比较长的荧光寿命和比较窄的荧光谱线，因而易于产生粒子数反转和受激发射。用作晶体类基质的人工晶体主要有：刚玉 、钇铝石榴石、钨酸钙、氟化钙等，以及铝酸钇、铍酸镧等。用作玻璃类基质的主要是优质硅酸盐光学玻璃，例如常用的钡冕玻璃和钙冕玻璃。与晶体基质相比，玻璃基质的主要特点是制备方便和易于获得大尺寸优质材料。对于晶体和玻璃基质的主要要求是：易于掺入起激活作用的发光金属离子；http://具有良好的光谱特性、光学透射率特性和高度的光学均匀性；具有适于长期激光运转的物理和化学特性。晶体激光器以红宝石和掺钕钇铝石榴石为典型代表。玻璃激光器则是以钕玻璃激光器为典型代表。
工作物质 固体激光器的工作物质，由光学透明的晶体或玻璃作为基质材料，掺以激活离子或其他激活物质构成。这种工作物质一般应具有良好的物理－化学性质、窄的荧光谱线、强而宽的吸收带和高的荧光量子效率。 玻璃激光工作物质容易制成均匀的大尺寸材料，可用于高能量或高峰值功率激光器。但其荧光谱线较宽，热性能较差，不适于高平均功率下工作。常见的钕玻璃有硅酸盐、磷酸盐和氟磷酸盐玻璃。80年代初期，研制成功折射率温度系数为负值的钕玻璃，可用于高重复频率的中、小能量激光器。 晶体激光工作物质一般具有良好的热性能和机械性能，窄的荧光谱线，但获得优质大尺寸材料的晶体生长技术复杂。60年代以来已有300种以上掺入各种稀土金属或过渡金属离子氧化物和氟化物晶体实现了激光振荡。常用的激光晶体有红宝石(Cr:Al2O3,波长6943
DB44 ICS31.260 L51 广东省地方标准 DB44/TXXXX—2014 光纤激光切割机通用技术规范 General technical specifications of fiber laser cutting machine 点击此处添加与国际标准一致性程度的标识 （工作组讨论稿） （本稿完成日期：2014-7-2） XXXX-XX-XX发布XXXX-XX-XX实施
目 次 前言.....................................................................................................................................................................IV 1范围.. (1) 2规范性引用文件 (1) 3术语和定义 (1) 4产品型号与构成 (4) 4.1产品型号 (4) 4.2产品构成 (5) 5技术要求 (5) 5.1工作环境要求 (5) 5.2技术参数 (6) 5.3外观质量 (6) 5.4制造质量 (6) 5.5装配质量 (6) 5.6附件和工具 (6) 5.7电气系统 (6) 5.8数控系统 (6) 5.9气动、冷却和润滑系统 (7) 5.10安全防护 (7) 5.11寿命 (8) 5.12可靠性要求 (8) 5.13噪音要求 (9) 6检验方法 (9) 6.1检验条件 (9) 6.2技术参数检验 (9) 6.3外观质量检验 (9) 6.4制造质量检验 (9) 6.5装配质量检验 (9) 6.6附件和工具检验 (9) 6.7电气系统检验 (9) 6.8数控系统检验 (9) 6.9气动、冷却和润滑系统检验 (9) 6.10安全防护检验 (10) 6.11寿命检验 (10) 6.12可靠性检验 (10) 6.13噪音检验 (10)
一．安全规程 1. 严格按照激光器启动程序启动激光器，调光，试切工件。 2. 操作者必须经过培训，熟悉切割软件，设备结构，性能，掌握操作 系统有关知识。 3. 按照规定穿戴好劳动防护用品，在激光束附近必须佩戴符号规定的 防护眼睛。 4. 设备开动时操作人员不得擅自离开岗位，如的确需要离开时应停机， 断开急停按钮下使能。 5. 不加工时应关掉激光器或光闸。 6. 保持激光器，激光头，机床及周围场地，有序，无油污，工件，材 料，废料按规定摆放整齐。 7. 气罐使用，应严格遵守气罐监察规程。开启气罐时，操作者必须站 在出气口侧面。 8. 维修时，必须严格遵守高压规程。关掉电源，检查设备故障原因。 9. 必须严格遵守消防安全规程。要将灭火器放在随手可及的位置。 10. 在未弄清某一种材料是否能够使用激光照射或切割时，不要对其 加工，以免产生烟雾和蒸汽的潜在危险。 11. 上下材料时要小心，确保人身安全。 二．操作规程 1 工作前 1.1 按规定穿戴劳动保护用品: 1.2 检查设备周围有无不安全因素: 1.3 必须按照《激光切割机日常点检一览表》进行检查，并填写记录； 符合要求后方可开始工作。 1.4 检查设备各部位是否正常，包括风、水、电、光路; 1.5 激光的开启、调整和关闭应根据操作手册进行。 1.6 激光切割机的开关机顺序及注意事项必须按照《大族激光切割机的操 作规程》进行操作。 1.7 开机空运转，确认设备无异常后方可工作。 1.8 所规定的各种气体入口压力不得超过如下数值: 1.8.1 氮气Max. 3.0MPa （30 bar ） 1.8.2 氧气Max. 1.0MPa （10 bar 1.8.3 空气Max. 0.8MPa （8 bar ） 2 工作中 2.1 设备如在如下情况下不得进行操作: 2.1.1 超长(4000mm) ，超宽(2000mm) ，超厚(20mm 以上); 2.1.2 板材不平度大于30mm以上: 2.1.3 程序不合理引起碰撞切割头时: 2.1.4 割件质量极差( 如锈蚀严重等)； 2.1.5 操作员认为切割易出现问题时; 2.2 确保各项加工参数正确，设备的移动体不与工装、工件、余料发生碰撞;
ASLD 功能介绍 1 热透镜分析 1.1小泵浦源的网格细化 在某些激光器中，泵浦光源的光束尺寸相对于激光晶体而言非常小。在对这样结构的系统进行模拟分析时需要使用到局部自适应的网格划分方法来准确反映系统参数。或者通过加大整体的网格数量来进行模拟计算，但是这样的方法不仅需要更多的计算时间，而且计算结果的准确性也很难保证。ASLD所使用的是第一种方法，在激光晶体的中心部分加入一个进行更精细划分的网格区域来准确包含模拟小泵浦源的激光系统。 1.2快速有限元分析（FEA） ASLD借鉴并使用了有限元分析的现代概念和算法。其中包含初始分析的计算方法和半粗化多重网格算法。初始分析及其相对应使用的算法能够保证ASLD在系统包含实时动态FEA仿真并且包含大量的有限元时依然能够快速计算出仿真结果。半粗化多重网格算法能够保证ASLD在对包含超长激光晶体的系统仿真任务中更快的计算出所需的结果。
1.3实时动态热透镜效应分析 实时动态热透镜效应分析在对包含闪光灯泵浦和脉冲泵浦的激光器进行仿真时具有明显的优势。该功能能够仿真出闪光灯泵浦光源每次打开和关闭的过程中晶体内部的温度和结构变化以及使用脉冲泵浦时不同的脉冲周期对系统的影响。 1.4FEA边界条件 ASLD包含的图形化用户界面能够帮助用户快速的设置FEA分析的边界条件该功能允许用户对激光晶体进行多个部分的划分并设置相应的条件。例如下图中所展示的就是利用该功能对激光晶体进行不同部分的冷却设置界面。 1.5自动抛物线拟合 在进行高斯模式分析的过程中通常需要对结构和温度数据进行多次抛物线拟合。通常这样的工作需要设计人员手工完成，该逆过程较为繁琐也容易出错。在ASLD中我们提供了针对结构和温度数据进行自动拟合的功能，该功能简便易用，并且能够在实时动态有限元分析过程中起到重要的作用。
第9卷第21期 2009年11月1671 1819(2009)21 6532 04 科学技术与工程 ScienceTechnologyandEngineering 2009 Sci Tech Engng 9 No 21 Nov.2009 Vol 通信技术 半导体激光器驱动电路设计 何成林 (中国空空导弹研究院,洛阳471009) 摘要半导体激光驱动电路是激光引信的重要组成部分。根据半导体激光器特点,指出设计驱动电路时应当注意的问题,并设计了一款低功耗、小体积的驱动电路。通过仿真和试验证明该电路能够满足设计需求,对类似电路设计有很好的借鉴作用。 关键词激光引信半导体激光器窄脉冲中图法分类号 TN242; 文献标志码 A 激光引信大部分采用主动探测式引信,主要由发射系统和接收系统组成。发射系统产生一定频率和能量的激光向弹轴周围辐射红外激光能量,而接收系统接收处理探测目标漫反射返回的激光信号,而后通过信号处理系统,最终给出满足最佳引爆输出信号。由此可见,激光引信的探测识别性能很大程度上取决于激光发射系统的总体性能,即发射激光脉冲质量。而光脉冲质量取决于激光器脉冲驱动电路的质量。因此,半导体激光器驱动电路设计是激光引信探测中十分重要的关键技术。 图1 驱动电路模型 放电,从而达到驱动激光器的目的。 由于激光引信为达到一定的探测性能,通常会要求激光脉冲脉宽窄,上升沿快,一般都是十几纳秒甚至几纳秒的时间。因此在选择开关器件时要求器件开关速度快。同时,由于激光器阈值电流、工作电流大 [1] 1 脉冲半导体激光器驱动电路模型分析 激光器驱动电路一般由时序产生电路、激励脉冲产生电路、开关器件和充电元件几个部分组成,如图1。 图1中,时序产生电路生成驱动所需时序信号,一般为周期信号。脉冲产生电路以时序信号为输入条件。根据其上升或下降沿生成能够打开开关器件的正激励脉冲或负激励脉冲。开关器件大体有三种选择:双极型高频大功率晶体管、晶体闸流管电路和场效应管。当激励脉冲到来时,开关器件导通,
激光切割机光纤 光纤激光切割机较CO2激光切割机的优势： 1) 卓越的光束质量：聚焦光斑更小，切割线条更精细，工作效率更高，加工质量更好; 2) 极高的切割速度：是同等功率CO2激光切割机的2倍; 3) 极高的稳定性：采用世界顶级的进口光纤激光器，性能稳定，关键部件使用寿命可达10万小时; 4) 极高的电光转换效率：光纤激光切割机光电转换效率达30%左右，是CO2激光切割机高3倍，节能环保; 5) 极低的使用成本：整机耗电量仅为同类CO2激光切割机的20-30%; 6) 极低的维护成本：无激光器工作气体;光纤传输，无需反射镜片;可节约大量维护成本; 7) 产品操作维护方便：光纤传输，无需调整光路; 8) 超强的柔性导光效果：体积小巧，结构紧凑，易于柔性加工要求。 当然了，与二氧化碳激光切割机相比，光纤的切割范围相对狭窄。因为波长的原因，其只能切金属材料，对非金属不容易被其吸收，从而影响其切割范围。 与YAG激光切割机相比的优势： 1)切割速度：光纤激光切割机的速度是YAG的4-5倍，适用于大量加工与生产 2)使用成本：光纤激光切割机的使用成本比YAG固体激光切割更少 3)光电转换效率：光纤激光切割机的光电转换效率是YAG的10倍左右 相应的光纤激光器的价格较高，所以光纤激光切割机价高比之YAG激光切割机要高出不少，但比二氧化碳激光切割机要低很多。但其性比价确实三者中最高的。 三维测量方式 1)将被测物体置于三坐标测量空间，可获得被测物体上各测点的坐标位置，这项技术就是三坐标测量机的原理。三坐标测量机是测量和获得尺寸数据的最有效的方法之一，可以替代多种表面测量工具，减少复杂的测量任务所需的时间，为操作者提供关于生产过程状况的有用信息。 2)三维激光扫描仪是通过发射激光来扫描被测物，以获取被测物体表面的三维坐标。三维激光扫描技术又被称为实景复制技术，具有高效率、高精度的测量优势。有人说，三维激光扫描是继GPS技术以来测绘领域的又一次技术革命。三维激光扫描仪被广泛应用于结构测量、建筑测量、船舶制造、铁路以及工程的建设等领域，近些年来，三维激光扫描仪已经从固定朝移动方向发展，最具代表性的就是车载三维激光扫描仪和机载三维激光雷达。
光纤激光切割机常见问题 1.概述 三维光纤激光切割机是由专用光纤激光切割头、高精度电容式跟踪系统、光纤激光器以及工业机器人系统组成，对不同厚度的金属板材进行多角度、多方位柔性切割的先进设备。三维机器人激光切割机设备广泛应用于金属加工、机械制造及汽车零部件制造等对3D工件有加工需求的生产中。 2.三维光纤激光切割机器人 (1)三维激光切割原理激光通过激光器产生后，由反射镜传递并通过聚集镜照射到加工物品上，使加工物品(表面)受到强大的热能而温度急剧增加，使该点因高温而迅速的熔化或者汽化，配合激光头的运行轨迹从而达到加工目的。 (2)光纤的选择根据金属板材的厚度不同，选用不同的光纤激光器功率，三维切割光纤激光器的功率一般分200W、300W、400W、500W与1000W等多种规格;对不同功率的激光器配备不同的冷却系统，以保障激光器的正常工作。同时要根据机械臂的工作半径和加工工件的大小选定合适长度的操作光纤传输激光，以满足客户切割要求。 (3)辅助气体的要求三维光纤激光切割机采用的辅助气体是99.99%的氧气，这样对切割的精度、速度和切割的断面效果有很大的帮助。 3.三维光纤激光切割机器人的特点 (1)柔性高尤其适合小批量的三维钣金切割。 其高柔性主要表现在两个方面： 第一，对材料的适应性强，激光切割机通过数控程序基本可以切割任意板材。 第二，加工路径由程序控制，如果加工对象发生变化，只须修改程序即可，这一点在零件修边、切孔时体现得尤为明显。由于修边模、冲孔模对其他不同零件的加工无能为力，而且模具的成本高，所以目前三维激光切割有取代修边模、冲孔模的趋势。一般来说，三维机械加工的夹具设计及其使用比较复杂，但激光加工时对被加工板材不施加机械加工力，这使得夹具制作变得很简单。此外，一台激光设备如果配套不同的硬件和软件，就可以实现多种功能。 总之，在实际生产中，三维激光切割在提高产品质量、生产效率，缩短产品开发周期、降低劳动强度、节省原材料等方面优势明显。因此，尽管设备成本高、一次性投资大，国内还是有很多汽车、飞机生产厂家购进了三维激光加工机，部分高校也购进了相应设备进行科研，三维激光技术势必在我国制造业中发挥着越来越大的作用。 (2)光纤激光切割机器人优缺点第一，用工业机器人代替五轴机床，两者都能进行空间轨迹描述实现三维立体切割。工业机器人的重复定位精度比五轴机床稍低，约为±100μm，但这完全可以满足汽车钣金覆盖件和底盘行业的精度要求;而采用工业机器人大大降低了系统的成本造价，减少了耗电系统费用和系统运行维护费用，减少了系统的占地面积。 第二，光纤激光相比传统激光，具有更好的切割质量，更低的系统造价，更长的使用寿命和更低的维护费用，更低的耗电。关键是光纤激光器的激光可以通过光纤传输，方便与工业机器人连接，实现柔性加工。 第三，本系统唯一的缺陷是只能加工金属工件，不能加工非金属工件。这是因为本系统采用的是光纤激光，其波长为1064nm，相对于波长为10640nm的CO2激光，不易为非金属材料所吸收。
第一台激光器——红宝石固体激光器摘要：本文主要回顾了第一台激光器的研制历程，介绍了红宝石激光器的工作原理和它的发明者梅曼先生。 一、发展历程 1917年，爱因斯坦(Einstein)在气体平衡计算的工作中，发现在自然界存在着两种发光形式:一种是自发辐射，一种是受激辐射。前者指的是自然光的发光形式，而第二种正是产生激光的基础理论。激光的定义就是:“利用辐射的受激辐射实现的光放大”( Light amplification by the stimulated emission of radiation )。爱因斯坦的观点被当时的第一次世界大战的枪炮声所淹没，对于受激辐射这一重妥概念的意义没有被人们及时认识到. 1921年，发明磁控管，从此开始了微波的研究。 1927年，狄拉克(Dirac)根据感应辐射的属性提出创制星子书瞬浮的建议。 1934年，克赖克汤和威廉}i} i}i}于振荡器发现了电磁波和分a:.的相互作用。这是最旱期的电磁波谱学实验。 30年代，一些科学家建立的量子力学理论，使爱因斯坦的这两种发光形式的物理内容得到更为深刻的阐明。同时，近代光谱学的发展，也为激光光的出现奠定了的理论基础. 1944年，扎沃依斯基发现了电子的顺磁共振，打下了对微波波段电子顺磁能级研究的基础. 1945年第二次世界大战结束以后，大扰物理学家问到大学工作，在大学里建起了强大的新设备.他们开始着手进行微波波谱学山研究。当时，韦伯(Webber )、法布里肯特、巴索夫(tacos)和普罗霍洛夫(11po1。二。。)以及汤斯("l}ow'nes)等科学家分别提出了用受激辐射获得放大的设想。这是激光理论发展的重要起点. 1946年在美、英两国几乎同时发现氨谱线中的精细结构和超精细结构。 关于波谱学最显著的成果是发现氢原子谱-的兰姆位移。这是哥伦比亚大学的兰姆( Larnb)和另一同事的共同成果。他们曾具休地论述了观测净受激发射(负吸收)的可能性，明确指出了粒子数反转能够在何种状态实现，并针对一定的入射波，粗略计算了它的增益。 作为激光的物理基础—受激辐射早在1917年就为人所知.可是，从1917年到1950年30多年来，在实验上却一直没有人去证明这个过程的存在.人们以为，要想在小于一亿分之一秒的时间里进行原子受激发射的宏观观察是难于做到的。但在后来激光器制成后.实验工作并不象人们最初所设想的那样艰难。从1940年观察到离子数反转到激光器，这中间仅仅一步之差，可是这“一步”却一直走了20年. 人类对电磁波的利用和无线电技术的发展，使社会和生产急需把这种利用由无线电波段向微波波段扩展，这就导致了微波放大理论及其器件的产生. 1951年，美国的汤斯提出了利用受激辐射获得放大的原理首先获得微波放大的设想.同年，普塞耳(I'urcell)和庞德(Pound)用核磁共振所进行的一次实验，造成了粒子数反转，进一步确认了受激辐射过程，给微波放大器的产生带来了希望。其后，汤斯进行了两年半的艰苦工作，干1953年末和果尔登(Gordon )、
光纤切割机工作原理 三维光纤激光切割机对产品切割质量有影响的因素一般有以下几点：1. 激光器功率大小 2. 辅助气压的大小和纯度 3. 切割头焦距的大小 4. 加工工件的厚度 5. 机器人的示教轨迹或离线编程的数模定位精度6. 机器人的路径重复精度、重复定位精度和速度 三维切割系统的技术优势： 1.因为采用了业内最高精度的史陶比尔机械手，本体较轻，切割速度快，在小弧度的精细切割和大边的高速切割方面具有明显优势，实际切割速度可以达到18米/分钟而无抖动，综合加工效率是其他品牌机械手组合的两倍，性价比高，还可以节约一组的耗材和人工，后期可以少追加设备也能满足产能要求。还可24小时持续工作。一次性投入相对较少，在一个很短的折旧期内(两班8小时工作制)，史陶比尔机器人激光解决方案就可回收投资。同时能耗少，体积小，维护需求低。 2.切割精度高。采用史陶比尔专利齿轮减速系统JCS和JCM，独一无二的驱动技术，确保了无可匹敌的轨迹控制精度和速度。即使是要求极高的小圆，或复杂立体几.何图形的加工，也可精确和快速完成，从而提升您的产品品质。系统重复定位精度高达±0.05M，完全可以满足钣金件行业的精度需求。可切割直径小至2MM的小圆，切割效果圆滑美观，目测无形变和毛刺。 3.切割幅面大，实际死角小。选配臂长2.01米的机械手，除了实现直径达3米的半球形三维加工区域外，还可实现较大的二维平面切割，配合我公司配套生产的可移动工作台2.5mX5m(2m的运动行程)，可实现2mX5m的二维平面切割。 4. 根据实际需要选配离线编程软件，可读取UG,SOLIDWORK等三维作图软件导出的vda,igs,x_t,sldprt,prt,stp,ipt,par等格式的数模，修改后直接生成切割轨迹，代替人工示教，简单易用。 5. 工业控制理念，模块化设计，全系统的防护等级为IP55,机械手防护等级更是高达IP65，系统集成度高，故障少，抗冲击振动，抗灰尘，无须光学调整或维护，真正适合于工业加工领域的应用用于恶劣的激光环境。结构坚固，动态性更佳。而其他同类产品为简单集成，设备的稳定性较差。 6.系统的工艺性和易用性较好。简单而功能强大的史陶比尔激光专用标准软件LasMAN 基于Windows操作系统，用户界面简单友好，集成了机器人运动控制、激光控制、数据处理和产品管理等功能。友好的人机界面，模块化的设计，使得操作者仅需经过简单的培训即可达到系统产能最大化，同时也易于集成。这就大大降低了对操作工人的要求，降低了对工人的管理难度。 性能参数 耗电耗材： 系统耗电：<8KW(根据选配激光器功率大小而异)
固体激光器设计步骤 1。如何选择激光介质？（就是你做一台激光器选择怎样的介质，比如是Nd：YAG还是Nd：YLF或者是Nd：YVO4，如果你要求单脉冲能量高重复频率较低的话就选择Nd：YLF为好，如果需要选择单脉冲能量高而重复频率低的话最好选择Nd：YVO4，当然Nd：YAG虽然两者都能，但是他也有其局限性；同时也还要考虑现在这些介质实验室和产品分别能做到多大功率）； 2。泵浦结构如何选择？（泵浦结构主要包括端泵、侧泵、面泵等等，那么选择这些结构的理由又分别是什么呢？考虑那些因素呢？） 3。冷却方式如何选择？（这个问题需要结合泵浦结构进行讨论，因为他限制了泵浦结构，是水冷还是风冷，是空气自然冷却还是用TEC进行冷却或者用热管进行冷却，你见过那些冷却方式，最好能提供参考文献） 3。腔结构如何选择？（腔大体上分稳定腔和非稳定腔，那么选择何种腔结构的原因分别是什么呢？每一种腔结构又包括具体很多结构，比如稳定腔又包括很多中，比如共焦、共心、平凹等等，选择的理由又是什么呢？对腔设计还要考虑热透镜效用的补偿，那么如何进行热透镜补偿呢？） 4。如何进行选模？（模式的选择我们大体可以分为两类，一类是横模的选择，另一类是纵模的选择，那么选择横模就设计到如何方式光阑，光阑放置在什么位置，这两者又与腔结构和腔长有关系，因为选择横模是依靠不同的模式具有不同的衍射损耗而实现的；而单纵模的选择就包括插入布鲁斯特片、光栅、棱镜、双折射波片、标准具等方法；那么这些方法各有什么有缺点呢？我们如何来选取这些方法呢？） 5。整体结构如何选择？（这儿的整体结构主要指就仅仅一个振荡器还是利用MOPA结构，如果使用MOPA结构，那么要使用多少级，级间效用怎么消除） 以上讨论的事实上还限制在实验室的阶段，要进行工程化不仅需要进行市场分析，还需要对器件进行合理的安排和包装，还要考虑到客户使用方便，系统功能稳定等等。我们目前先考虑上面的这几个问题。
光纤激光切割机器人
三维激光切割机 公司根据前期大量的市场调研，结合汽车钣金覆盖件和底盘件的行业特点，现推出工业机器人+光纤激光器的组合进行三维切割，耗材耗电总费用控制在每小时20元内，彻底有效的解决了上述问题。 首先，用工业机器人代替五轴机床。两者都能进行空间轨迹的描述实现三维立体切割，工业机器人的重复定位精度比五轴机床稍低，约为±100uM，但这完全可以满足汽车钣金覆盖件和底盘件行业的精度要求了。而采用工业机器人切割效率高，相当于传动五轴激光切割机床切割速度的两倍，大大降低了系统的成本造价，减少了耗电系统费用和系统运行维护费用，减少了系统的占地面积。 其次，用光纤激光器代替CO2激光器。光纤激光技术是近几年高速发展的激光技术，相比传统激光，具有更好的切割质量，更低的系统造价，更长的使用寿命和更低的维护费用，更低的耗电。关键是光纤激光器的激光可以通过光纤传输，方便与工业机器连接，实现柔性加工。 总之，采用工业机器人+光纤激光器的组合进行加工，修边冲孔等工艺一次完成，切口整齐无需后道工艺再处理，大大缩短了工艺流程，降低了人工成本和投入，也提高了产品档次和产品附加值。LasMAN专用激光软件的使用，支持通过数模直接生成切割轨迹，抛弃了繁杂的人工示教，更加适合小批量多批次的维修市场、新品试制和非标定制等一些个性化的切割需求。而且，投资高柔性高效率的激光切割设备，来代替昂贵的冲压设备和剪裁设备，可以更加灵活的更换产品，把握市场。
三维切割系统的技术优势： 1.因为采用了业内最高精度的史陶比尔机械手，本体较轻，切割速度快，在小弧度的精细切割和大边的高速切割方面具有明显优势，实际切割速度可以达到18米/分钟而无抖动，综合加工效率是其他品牌机械手组合的两倍，性价比高，还可以节约一组的耗材和人工，后期可以少追加设备也能满足产能要求。还可24小时持续工作。一次性投入相对较少，在一个很短的折旧期内(两班8小时工作制)，史陶比尔机器人激光解决方案就可回收投资。同时能耗少，体积小，维护需求低。 2.切割精度高。采用史陶比尔专利齿轮减速系统JCS和JCM，独一无二的驱动技术，确保了无可匹敌的轨迹控制精度和速度。即使是要求极高的小圆，或复杂立体几.何图形的加工，也可精确和快速完成，从而提升您的产品品质。系统重复定位精度高达±0.05M，完全可以满足钣金件行业的精度需求。可切割直径小至2MM的小圆，切割效果圆滑美观，目测无形变和毛刺。 3.切割幅面大，实际死角小。选配臂长2.01米的机械手，除了实现直径达3米的半球形三维加工区域外，还可实现较大的二维平面切割，配合我公司配套生产的可移动工作台2.5mX5m(2m的运动行程)，可实现2mX5m的二维平面切割。 4. 根据实际需要选配离线编程软件，可读取UG,SOLIDWORK等三维作图软件导出的vda,igs,x_t,sldprt,prt,stp,ipt,par 等格式的数模，修改后直接生成切割轨迹，代替人工示教，简单易用。 5. 工业控制理念，模块化设计，全系统的防护等级为IP55,机械手防护等级更是高达IP65，系统集成度高，故障少，抗冲击振动，抗灰尘，无须光学调整或维护，真正适合于工业加工领域的应用用于恶劣的激光环境。结构坚固，动态性更佳。而其他同类产品为简单集成，设备的稳定性较差。 6.系统的工艺性和易用性较好。简单而功能强大的史陶比尔激光专用标准软件LasMAN基于Windows操作系统，用户界面简单友好，集成了机器人运动控制、激光控制、数据处理和产品管理等功能。友好的人机界面，模块化的设计，使得操作者仅需经过简单的培训即可达到系统产能最大化，同时也易于集成。这就大大降低了对操作工人的要求，降低了对工人的管理难度。 性能参数 耗电耗材： 系统耗电：<8KW(根据选配激光器功率大小而异) 零星耗材：<0.5元/小时(包括高功率激光器水冷系统的滤芯、切割头气嘴和切割头保护镜片) 吹气费用：<6元/小时(以用纯氧辅助切割2MM内碳钢为例) 性能指标: 激光功率: 200W/300W/400W/500W/1000(根据工件材质和料厚可选) 激光波长:1070NM 工作区域:半径2米的半球形工作区域(选配半径2米的机械手) 切割速度:0-18米/分钟(根据功率大小和工件材质与厚度可调) 供电电源:三相交流380V
固体激光器原理及应用 摘要：固体激光器目前是用最广泛的激光器之一，它有着一些非常突出的优点。本论文先从基本原理和结构介绍固体激光器，最后介绍其在监测，检测，制造业，医学，航天等五个方面的应用及未来的发展方向。 关键词：固体激光器基本原理基本结构应用 1激光与激光器 1.1激光 1.1.1激光（LASER） 激光是在 1960 年正式问世的。但是，激光的历史却已有 100多年。确切地说，远在 1893年，在波尔多一所中学任教的物理教师布卢什就已经指出，两面靠近和平行镜子之间反射的黄钠光线随着两面镜子之间距离的变化而变化。他虽然不能解释这一点，但为未来发明激光发现了一个极为重要的现象。 1917年爱因斯坦提出“受激辐射”的概念，奠定了激光的理论基础。激光，又称镭射，英文叫“LASER”，是“Light Amplification by Stimu Iatad Emission of Radiation”的缩写，意思是“受激发射的辐射光放大”。激光的英文全名已完全表达了制造激光的主要过程。1964年按照我国著名科学家钱学森建议将“光受激发射”改称“激光”。 1.1.2产生激光的条件 产生激光有三个必要的条件： 1)有提供放大作用的增益介质作为激光工作物质，其激活粒子（原子、分 子或离子）有适合于产生受激辐射的能级结构； 2)有外界激励源，将下能级的粒子抽运到上能级，使激光上下能级之间产 生粒子数反转； 3)有光学谐振腔，增长激活介质的工作长度，控制光束的传播方向，选择 被放大的受激辐射光频率以提高单色性。 1.1.3激光的特点 与普通意义上的光源相比较，激光主要有四个显著的特点：方向性好、亮度极高、单色性好、相干性好。
郑州轻工业学院 课程设计任务书 题目半导体激光器原理及应用 专业、班级学号姓名 主要内容、基本要求、主要参考资料等： 完成期限： 指导教师签名： 课程负责人签名： 年月日
郑州轻工业学院半导体激光器课程设计 郑州轻工业学院 课程设计说明书题目：半导体激光器原理及应用 姓名：王森 院（系）：技术物理系 专业班级：电子科学与技术09-1 学号：540911010132 指导教师：运高谦 成绩： 时间：年月日至年月日 I
郑州轻工业学院半导体激光器课程设计 摘要 本文主要讲的是半导体激光器的发展历史、工作原理及应用。半导体激光器产生激光的机理,即必须建立特定激光能态间的粒子数反转,并有合适的光学谐振腔。由于半导体材料物质结构的特异性和其中电子运动的特殊性,首先产生激光的具体过程有许多特殊之处,其次所产生的激光光束也有独特的优势,使其在社会各方面广泛应用。从同质结到异质结,从信息型到功率型,激光的优越性也愈发明显,光谱范围变宽,相干性增强,可以说是半导体激光器开启了激光应用发展的新纪元。 关键词激光技术；半导体激光器；受激辐射；光场 II
郑州轻工业学院半导体激光器课程设计 Abstract This article is mainly about the history of the development of semiconductor lasers, working principle and applications. Semiconductor lasers produce laser mechanism, which must be established between the specific laser energy state population inversion, and a suitable optical resonator. As the physical structure of the semiconductor material in which electron motion specificity and particularity, while the specific process of producing laser has many special features, the other produced by the laser beam has a unique advantage to make it widely used in all sectors of society . From homo-junction to the heterojunction, the power from the information type to type, is also becoming increasingly apparent superiority of the laser, spectral range, coherence enhanced semiconductor lasers opened a new era in the development of laser applications. Keywords: Laser technique；Semiconductor lasers；Stimulated emission；Optical field III
大家应该都知道，光纤激光切割机使用的激光器是光纤激光器，那么，有人知道光纤激光切割机的激光器工作原理是怎样的么？今天，专注激光切割机多年的小编就来为大家普及一下。 首先，来看看什么是光纤激光器。光纤激光器（Fiber Laser）是指用掺稀土元素玻璃光纤作为增益介质的激光器，利用泵浦光照射掺有稀土元素的玻璃光纤泵浦源内，促使稀土离子吸收光子后受激发辐射与入射光子同频率，传播方向和振动方向的相干光，通过光放大从而产生高能量的激光。 光纤激光器应用范围非常广泛，包括材料加工，通信，医疗美容，科研与军事，仪器仪表与传感器；材料加工作为激光应用最大的领域，发展出激光打标，切割，焊接，熔覆，清洗，表面处理等多种应用技术，而激光切割已成为激光技术应用最广泛的领域。 由于光纤激光器采用的工作介质具有光纤的形式，其特性要受到光纤渡导性质的影响。进入到光纤中的泵浦光一般具有多个模式，而信号光电可能具有多个模式，不同的泵浦模式对不同的信号模式产生不同的影响，使得光纤激光器和放大器的分析比较复杂，在很多情况
下难以解析，不得不借助于数值计算。 光纤中的掺杂分布对光纤激光器也产生很大的影响，为了使介质具有增益特性，将工作离子（即杂质）掺杂进光纤。一般情况下，工作离子在纤芯中均匀分布。但不同模式的泵浦光在光纤中的分布是非均匀的。因而，为了提高泵浦效率，应该尽量使离子分布和泵浦能量的分布相重合。在对光纤激光器进行分析时，除了基于前面讨论的激光器的一般原理，还要考虑其自身特点，引入不同的模型和采用特殊的分析方法，以达到好的分析效果。 和传统的固体、气体激光器一样，光纤激光器也是由泵浦源、增益介质、谐振腔三个基本要素组成。泵浦源一般采用高功率半导体激光器，增益介质为稀土掺杂光纤或普通非线性光纤，谐振腔可以由光纤光栅等光学反馈元件构成各种直线型谐振腔，也可以用耦合器构成各种环形谐振腔。泵浦光经适当的光学系统耦合进入增益光纤，增益光纤在吸收泵浦光后形成粒子数反转或非线性增益并产生自发发射。所产生的自发发射光经受激放大和谐振腔的选模作用后，形成稳定激光输出。 关于光纤激光切割机的激光器工作原理，宏山激光小编就介绍到这里了，宏山激光多年
光纤激光切割机操作规程 1.范围 本规程规定了光纤激光切割机操作的一般要求、特别要求和开关机顺序。 2.一般要求 2.1.遵守光纤激光切割机安全操作规程。严格按照激光器启动程序启动激光器。 2.2.操作者须经过培训，熟悉设备结构、性能，掌握操作系统有关知识。 2.3.按规定穿戴好劳动防护用品，在激光束附近必须佩带符合规定的防护眼镜。 2.4.在未弄清某一材料是否能用激光照射或加热前，不要对其加工，以免产生烟雾和蒸气的潜在危险。 2.5.设备开动时操作人员不得擅自离开岗位或托人待管，如的确需要离开时应停机或切断电源开关。 2.6.要将灭火器放在随手可及的地方;不加工时要关掉激光器或光闸;不要在未加防护的激光束附近放置纸张、布或其他易燃物。 2.7. 开机后应手动低速X、Y、Z轴方向开动机床，检查确认有无异常情况。 2.8．出光前检查铜嘴，保持干净，无堵塞；出光后检查光斑，保持圆是均匀的。 2.9.送料前观察材料有无起拱，材料起拱严重时不可使用，以免撞上激光头，后果严重。 2.10.生产运行前要检查所有准备工作是否到位，保护气是否开启，气压是否达到。激光是否待命状态。 2.11.在加工过程中发现异常时，应立即停机，及时排除故障或上报主管人员。 2.12.下班后清理抽风斜板铁渣，清理机台灰尘，废料。 3.特别要求 3.1材料的厚度不同，使用的扩束镜与聚焦镜也不同，更换扩束镜与聚焦镜时，必须有技术部人员在场，操作者不得私自更换扩束镜与聚焦镜；更换扩束镜与聚焦镜后需要标定。 3.2设备处于自动工作运转中非常危险，绝不可进入机床工作运动区域内。任何操作过程都必须注意安全。无论任何时间进入机床运转范围内都可能导致严重的伤害。 3.3每班检查水位，要求至少浸过蒸发器盘管。 3.4机器停止工作，外界环境温度低于0℃时，务必把水放干，腔体、蒸发器、水箱、水泵内不能有积水，防止结冰损坏机器；再次生产启动机器，气温仍很低，腔体或管内仍有少许冰粒，务必用25℃~30℃温水加入水箱，启动循环系统融化冰粒，再进行生产 4.开关机顺序