Experimental investigation and modelling of a buoyant attached plane jet in a room

合集下载

中英文文献翻译—关于汽车喷油泵端面凸轮轨道冷锻的研究

附录附录A：Research on orbital cold forging for the edge cam of automobile fuel injection pumpABSTRACT: The experimental investigation and theoretical analysis of an orbital cold forging of an edge cam was explored. The effects of blank shape and the process parameters on the edge cam forming are discussed. Photoplastic technology was successfully applied to the simulation of an orbital cold forging process applied to an edge cam. The China-made polycarbonate (PCBA) was used as a simulation material, and a 3-D strain distribution was obtained inside the model materials, which provided theoretical guidance for optimizing process parameters on the orbital cold forging of the edge cam. The factors that caused a crack were identified and the deformation laws in orbital cold forging of an edge cam were clarified. The metal flow line, hardness, microstructures and accuracy of the orbital cold forged edge cam were found to meet the required service properties of the product.Keywords: Orbital cold forging, edge cam, photoplasticity, simulation1. INTRODUCTIONThe edge cam of the automobile fuel injection pump is a key part with complicated shape and high precision, encountered high contact stress and a high shearing stress. It is hard to be made by the conventional forming processes or machining processes, and it doesn’t meet service properties of the product. The structural character and the forming property of materials of the edge cam were analyzed. Orbital cold forging is a forging process where a fixed bottom die and a moveable/orbiting top punch come together to form complex geometric workpieces with finished dimensional accuracy in a single forging operation. The orbiting upper die rolls over a metal blank while the bottom die is raised hydraulically. The billet is kneaded into the bottom die with relatively little force to produce a Near Net Shape or finished workpiece. This process features the orbiting upper die. Due to this orbiting motion over the workpiece, the resultant localized forces can achieve a high degree of deformation at a relatively low force level. As per the adjacent diagram, the forming force required, compared to conventional forming, is much lower due to a smaller contact area.2. EXPERIMENT2.1 Experimental equipmentThe orbital forging press—Type PXWP-100C with capacity of 1600KN was employedfor forming experiment. The orbital head completes 200 revolutions per minute. The inclined angle of the orbital head is adjustable 0° to 2° and there are four orbits selectable: circle, straight, spiral and daisy.2.2 Selecting the blank shapeThe geometric shape of the blank influences greatly the filling performance and the die life of orbital cold forged parts. If the selection is incorrect, it would either damage the die or make the work piece formation unsuccessful. According to the characteristic of the geometric shape of the edge cam, three kinds of blank shape were adopted to test.Blank a) with the step and flange:The size of the protruding step of blank with the function of fixing position in the lower die is basically same as the size of the edge cam. The shape of the edge cam can be formed by the orbital forging with small deformation.Cylindrical blank b) with small step:The size of the protruding step of blank with the function of fixing position in the lower die is basically same as the size of the edge cam. The shape of the edge cam can be formed by the orbital forging with large deformation.Cylindrical blank c):By the outside diameter of the cylindrical blank to fix position in the lower die, the shape of whole edge cam can be formed by the orbital forging with large deformation.All three types of blanks can be successfully formed to the edge cam. Although the deformation of blank a) and blank b) are easier to form than the blank c), the blank a) and the blank b) require being preformed in mass production, thus it will incur high costs. In contrast, although the deformation of the blank c) is higher, its geometric shape is much simpler. Preforming is not a necessary process for blank c). It is suitable for mass production.The work hardening of material plays a very important part in orbital cold forging. So to avoid the surface cracking of the blank in orbital cold forging, the softening annealing and good lubricating for blank are very much required. The orbital cold forging process for the edge cam is as below:Blank——Spheroidizing annealing ——Baiting——Phosphating and soap treatment——Orbital cold forging.2.3 Selecting orbital cold forging ParametersIf the orbit is a circle, the angle of oscillation can range from 0° to 2°. Once the angle is decided, the movement of the rocker will be not changed. If the angle is larger, the area between the rocker and work piece surface is also decreased, so the required deformation force is rather less, the forming time shorter. While a larger angle can bring higher efficiency, the accuracy of the parts will also be affected. This type of orbit is suited forproducing axisymmetrical parts. If the spiral orbit is selected, orbital forging will be helpful for the radial and axial flow of the metal, and also has a better centre deformation; this rocker will apply action cyclically to the central area of the blank. So it is easy to form a part with a much complicated end face. Selecting a linear orbit, it will be easier to form a longer part such as hammer and chisel. If a daisy orbit is selected, the part with tooth profile such as bevel gear and jaw clutch will be formed more easily.The angle of oscillation directly influences the deformation resistance and metal flow, the deformation resistance of theorbital cold forging is calculated as formula below:Where: S—— feed per revolution, mm/r；R—— maximum radius of the orbital cold forged part;γ—— angle of oscillation;σs—— yield strength of materials;K ——influence coefficient of the friction, the inhomogeneous distribution of stress and the shape of forged part etc.For the edge cam, S=1.2 mm/r; γ= 2°, σs=700 N/mm2 (the average yield strength of material 20CrMo), K=1.8, R=32.5mm，So, P=1444KN.Thus, if the angle of oscillation is increased, the loading force is decreased, this is helpful for the radial flow of the metal. The maximum angle of oscillation 2° was applied for orbital cold forging of the edge cam. According to the characteristics of the edge cam’s geometric shape and the type of orbit, the circular motion of the orbital head is selected.The process trial proves the filling property of metal is better, and the clear top & bottom faces of the edge cam was obtained.The lower die of orbital cold forging is generally very highly stressed. The lower die of orbital cold forging is similar to extrusion die. The lower die of orbital cold forging for the edge cam is reinforced by two stress rings. The lower die of orbital cold forging with axially-split inset was employed so as to minimize dangerous stress concentrations.In order to prevent the curve surface and the protruding step of the edge cam from being worn down and a fracture failure, a mobile core is employed in the process, and then an ejector is used to push the forged part out after it is being formed.3. PHOTOPLASTIC SIMULATIONOrbital cold forging deformation for the edge cam was studied with a photoplastic method. The model blank of photoplastic simulation need not to be split before forming. So the defects which the approximate portion of other experimental simulation methods is toolarge in studying the large deformation are avoided for photoplastic method. The photoplastic method can directly give us a set of the difference lines of equal principal strain, and a set of the direction tracks of principal strain in the model. It has many advantages such as: the strain diagram was directly perceived through the senses, and it’s reality, high measuring precision and high sensitivity. It is convenient for data collection, and provides a good way to study the plastic deformation which reflects the real situation.The China made polycarbonate (PCBA) was used as a simulation material. The blank size of the model material (PCBA) was same as the blank size of the edge cam. The blank of the model material was directly orbital cold forged. The full-field strain distribution for orbital cold forging deformation was indirectly obtained by the similarity relationship.The full-field strain distribution which is on the section with height of 17mm along the Z axial direction of the edge cam for orbital cold forging.From the strain distribution diagram, it was observed in the deformation of the edge cam, that the εZ of intermediate zone is compression strain, the radial strain εr and the tangential strain εθ of intermediate zone are tension strain where r≈10mm. This is identical with ordinary cylinder upsetting, and the fluctuation of strain value is low. But the fluctuation of strain value is higher where the intermediate zone r>10mm. It indicates that the deforming zone is inhomogeneous deformation. The max.radial strain εr was attained and the lar ger tangential strain εθ was also attained where r=25mm. As a result, the cracks occur easily in this zone. This is consistent with the cracks occurrence of orbital cold forging of the edge cam. The strain distribution of each Z axis section and deformation rules can be seen from the strain distribution diagram in orbital cold forging of the edge cam, such as the cracks occur easily in the maximal tension strain zone. The deformation homogeneity and detailed deforming of each deformation zone can be explored by the fluctuation of each strain value. It provides experimental basis for keeping defects of deformation cracks from happening, and fixing quantify datum for process experimental analysis. Thereby correct metal deformation laws were obtained in orbital cold forging of the edge cam.4. RESULT AND ANALYSISUpon the completion of optimizing the process parameters, orbital forging of the edge cam analysis and experimental trials, the qualified orbital forged edge cam was obtained.The microstructure of blank (Spheroidizing annealed condition) is composed of ferrite and pearlite.After blank forging, the grain is elongated and distorted. It takes on the obvious fiber structure, and the crystal boundary and the slip line can’t be identified. Thus the grains we re fragmentized in orbital forging of the edge cam, the amount of recrystallizing nucleus wereincreased in the subsequent heat treatment.The interior zone of the edge cam is free from any defects. But the metal flow line of the machining made edge cam is cut off. The brinell hardness of blank (annealed condition) is 130-135HB on average, after deformation, as result of work hardening, the highest hardness of the edge cam reaches to 287HV, a 110% increase. The hardness distribution of the edge cam isn’t u niform. The hardness of the edge cam is higher where the strain is higher. This is consistent with experimental simulation and physical measurement. It indicates that the mechanical property of the edge cam by orbital cold forging is better than the machining made edge cam. The work hardening of the material caused by orbital cold forming increases tensile strength and hardness of the edge cam. The wear resistant property of the edge cam was greatly improved.5. CONCLUSIONSUpon the completion of the research on the orbital cold forging of the edge cam, the following conclusions are drawn:(1) The complicated 3D curved parts with finished dimensional accuracy such as the edge cam of automobile can be formed in a single forging operation.(2) The photoplastic technology can be applied to simulation of orbital cold forging process of the edge cam.(3) The photoplastic technology gives us a theoretical basis for further exploration on orbital cold forging process, as well as the selection of the optimization process parameters.(4) The metal flow line, hardness and metallographic structure of the orbital forged edge cam meet the service performance, which is better than machining made part.(5) In comparison with hot precision forging, the orbital cold forging of the edge cam can save 1/3 of the material consumption, and the productivity increases by over five times. In comparison with machining, the orbital cold forging of the edge cam can save 2/3 of the material consumption, and the productivity increases by over ten times.附录B：关于汽车喷油泵端面凸轮轨道冷锻的研究摘要:对试验研究和理论分析的端面凸轮机构轨道冷锻进行了探讨。

Modelling, simulation and experimental investigation

ORIGINAL ARTICLEModelling,simulation and experimental investigation of cutting forces during helical milling operationsChangyi Liu &Gui Wang &Matthew S.DarguschReceived:21September 2011/Accepted:23January 2012/Published online:18February 2012#Springer-Verlag London Limited 2012Abstract The kinematics of helical milling on a three-axis machine tool is first analysed.An analytical model dealing with time domain cutting forces is proposed in this paper.The cutting force model is established in order to accurately predict the cutting forces and torque during helical milling operations as a function of helical feed,spindle velocity,axial and radial cutting depth and milling tool geometry.The forces both on the side cutting edges and on the end cutting edges along the helical feed path are described by considering the tangential and the axial motion of the tool.The dual periodicity which is caused by the spindle rotation,as well as the period of the helical feed of the cutting tool,has been included.Both simulation and experiments have been performed in order to compare the results obtained from modelling with experiments.Keywords Helical milling .Hole machining .Cutting forces .Analytical model .Time domainNomenclature a e i ,a e *Radial cutting depth of side cutting edge andend cutting edge (millimetres)a p i ,a p *Axial cutting depth of side cutting edge and endcutting edge (millimetres)D m Milling tool diameter (millimetres)F Cutting force (newtons)f va Axial component of helical feed speed (millimetres per second)f vt X –Y plane component of helical feed speed (millimetres per second)f za Axial component of helical feed rate per tooth (millimetres)f zt X –Y plane component of helical feed rate per tooth (millimetres)h i ,h *Instantaneous undeformed chip width of side cutting edge and end cutting edge (millimetres)K rc ,K tc ,K ac Cutting force coefficients of radial,tangential and axial direction (newtons per square millimetre)K re ,K te ,K ae Cutting force coefficients of edge effect (newtons per millimetre)K *vc ,K *nc Tangential and normal cutting force coefficients of end cutting edges (newtons per square millimetre)K *ve ,K *ne Tangential and normal cutting force coefficients of edge effect (newtons per millimetre)P Pitch of the helix feed trajectory N m Flute number of the milling toolv Velocity of milling tool or velocity of a point of the cutting edge (millimetres per second)t Time (seconds)βHelix angle of the milling tool fluteθAngular of motive direction and X –Y plane of a point of the cutting edge (radians)ϕϕj Relative rotational angle of milling tool and the cutting tooth j (radians)Φst ,Φex Cut-in and cut-out relative rotational angle of the cutting toolΦB Diameter of the hole (millimetres)ΦODiameter of the helical feed trajectory in X –Y plane (millimetres)C.Liu (*)Nanjing University of Aeronautics &Astronautics,Nanjing,Jiangsu,Chinae-mail:liuchangyi@G.Wang :M.S.DarguschCAST CRC,School of Mechanical and Mining Engineering,The University of Queensland,Brisbane,Queensland,Australia G.Wange-mail:gui.wang@.au M.S.Dargusche-mail:m.dargusch@.auInt J Adv Manuf Technol (2012)63:839–850DOI 10.1007/s00170-012-3951-4ΩSpindle rotating angular velocity(radians per second)Ωh Helix feed rotating angular velocity(radians per second)1IntroductionHelical milling has been applied to generate boreholes by means of a milling tool to some difficult-to-cut materials. This innovative method was found to facilitate hole making in AISI D2tool steel in its hardened state,resulting in an enhancement in cutting tool life and the ability to machine H7quality holes with a surface finish of0.3μm Ra[1].The operation has also been applied to hole making in composite-metal compounds as a substitute for drilling operations.The impact of the axial and tangential feed per tooth on the process forces[2]has been investigated. Employing helical milling to aluminium with minimum quantity lubrication has shown an improvement in geometri-cal accuracy and a reduction in burr formation,lower cutting temperature and a smaller cutting force compared to drilling operations[3].The prediction of cutting force through modelling and simulation is an important research area in order to improve process ling is the most complex machining operation.Previously in the literature,machining mechanisms have been derived from a general model[4,5]and applied to the specific application,for example,five-axis milling, three-axis milling,peripheral milling,face milling and plunge milling.Modelling peripheral milling is a fundamental requirement in order to model more complex milling operations.A theoretical model based on the oblique cutting principle and cutting force coefficients has been developed in order to predict the cutting forces during peripheral milling[6–8].Considering the helical flute(or side cutting edge)of the milling cutters,an attempt to accurately simulate milling forces including the effects of engaged flute length and the number of engaged flutes caused by the radial and axial depths of cut has been previously presented[9].A common approach to facilitate the modelling of this complex situation including the milling tool geometry and the interaction with the workpiece involves analysing the cutting forces on axial discrete milling tools,then integrating these force elements.The intersection of the tool path swept envelope with the workpiece Z-buffer elements has been used to find the contact area between the cutter and the workpiece. An axial slice cutting tool discrete mechanistic model was used to estimate the cutting force vectors[10].Cutter entry and exit angles,along with the immersion angles,were used as boundary conditions in order to predict cutting forces when flank milling ruled surfaces with tapered,helical and ball end mills[11].The effect of lead and tilt angles between the cutter and the workpiece on the milling forces,tool deflections and form errors during multi-axis end milling have been analysed[12,13].During modelling of the cutting forces and system dynamics,one of the outstanding characteristics is that both side cutting edges and end cutting edges interact with the workpiece during helical milling processing.An accurate predictive model should describe and sum up the mechanics on both edges simultaneously.Ball end milling tools are most often used in three-axis or five-axis milling.Ball end milling tool processing models have been separated into ball end and cylindrical sections in order to obtain accurate prediction[10,14,15].A mechanistic force model describing the cutting force as a sum of the cutting and edge forces has been developed for a general end milling cutter(cylindrical,taper,ball,bull nose)with the specific cutting and edge force coefficients identified[16].As one type of three-axis milling operation,axial feed is a typical characteristic of helical milling operations. This operation uses a flat end mill not a ball end mill that is used in typical3-axis and5-axis milling situations.Axial feed using a flat end mill is also applied in plunge milling which is a two-axis operation.Considering rigid body motion of the cutter,the cutting force model and dynamics model for the plunge milling process in the time domain have been established[17,18].The cutting forces associated with plunge milling operations are predicted by considering the feed,radial engagement,tool geometry,spindle speed and the regenera-tion of the chip load due to vibrations[19].Considering the flexibility of the workpiece,tool setting errors and tool kine-matics and geometry,a horizontal approach was used to compute the chip area including the contribution of the main and side edge in the cutting zone[20].Drilling operations and boring operations typically involve axial feed.Both these operations are similar to helical milling and plunge milling operations but with different cutting tools.The drilling cutting forces and dynamics have been integrated into the model in order to obtain drilled hole profiles[21].A mechanistic model for predicting thrust force and torque during the drilling process using a drill tool with double-point angle edges [22].To predict temperatures and forces on both the drilling and ball end milling operations,the cutting edges of the twist drill lip and the ball end mill were divided into oblique cutting elements[23].A theoretical model to predict thrust and torque in high-speed drilling has been presented[24,25].The methodology for extracting cutting force coefficients for drilling operations has also been investigated[26].When modelling the drilling process, the axial feed effect was not considered explicitly because the lip of the twist drill has a taper angle(point angle),and the interaction between the lip and workpiece caused by spindle rotation could lead to a spontaneous axial force(thrust).In the literature,helical milling has been introduced as an enabling technology to substitute for drilling operations [1–3].In recent years,research on modelling the mechanics of the helical milling process has been published [27,28].Although both the side cutting edges and the end cutting edges have been considered to participate in the machining process,the detail interaction between the end cutting edges and workpiece still needs more elaborate investigation and description.Modelling,simulation and experimental investigation during cutting forces of the helical milling operation will be discussed in this paper including the influence of helical feed.This research aims to develop an analytical cutting force model in the time domain including both the axial cutting depth and the radial cutting depth associated with helical milling operations.The model considers the effects of both the tangential feed and axial feed,and the combination of both mechanics on the side cutting edges and the end cutting edges.2Kinematics of helical millingIn helical milling,the trajectory of a point on the milling toolcutting edge is the result of the spiral curve movement of the axis of the tool (reference frame)and the circular movement of the edge point relative to the axis (relative motion).Two sets of coordinates are defined to describe the motion of the cutter and the cutting force on the cutter;an X,Y ,Z global coordinate frame fixed to the workpiece and an x,y,z local coordinate frame fixed to the cutting tool with the origin at the centre of the end flat surface which defines the reference frame.A description of helical milling with tool feed using helical trajectory and the coordinate settings are depicted inFig.1.The feed motion of the tool is decomposed into two components,f va and f vt .f vt ¼ΦB ÀD m ðÞΩh 2¼N m Ωf zt2p mm =s ðÞð1Þf va ¼P Ωh 2p ¼N m Ωf za 2pmm =s ðÞð2ÞThe flat-end cylinder milling tools suitable for helical milling operations have two types of cutting edges:the side cutting edge (peripheral cutting edge)and the end cutting edge through the centre.The interaction characteristics of these two types with the workpiece are different.The side edges participate in the peripheral cutting component,while the end edges participate in the plunge cutting component.Therefore,these two movements will be initially analysed separately before being assembled or composed.The side edge cutting process is typical intermittent cutting.The undeformed chip geometry,width,depth,and thickness have been described in the literature [2].The side edge cutting process that is typical intermittent cutting is depicted in Fig.2(using superscript i ).The velocity composition of an arbitrary point on the side cutting edge is described in cross section perpendicular to the tool axis.The undeformed chip geometry can be described as a i e ¼D m ;hole generating ΦB ÀΦO2;hole enlarging&ð3Þa i p ðt Þ¼f va t ;t 2p =Ωh P ;t >2p =Ωh&ð4Þh i ¼f zt sin fð5ÞFig.1Kinematics of helical millingwhere ϕ¼2p ΩÆΩh ðÞt is the relative rotational angle of the cutter (+up milling,−down milling).The end edge cutting process,which is continuous cutting,is depicted in Fig.3(using superscript *).The velocity composition of an arbitrary point on the end cutting edge is described in the cross section perpendicular to the end cutting edge.The undeformed chip geometry,width and height can be described as:a Ãe ¼D m ;hole making ΦB ÀΦO 2;hole enlarging &ð6Þh Ã¼f za cos θð7Þ3Cutting force model for helical milling 3.1Cutter feed influence on the cutting forcesThe influence of cutter feed movement on the cutting forces during machining processing is almost always neglected.Similar to spindle rotation resulting in the relative movement between cutter and workpiece,cutter feed motion leads to relative movement also.This relative movement between the cutter and workpiece could influence the directionand magnitude of the cutting forces.The premise that the influence of the feed can be neglected is based on the assumption that the relative displacement and velocity from spindle rotation are much larger than the feed.Thus,in most situations,the influence of feed is insignificant and can be ignored.However,when modelling some specific machining operations including axial feed,such as drilling,plunge milling and helical milling,to ignore the feed motion is unreasonable.If the axial feed effect is not considered,the cutting force along the axial direction might not be expressed accurately.For this reason,analysis of the influence of axial feed on cutting forces when modelling helical milling operations is necessary.In this paper,the feed motion effect on cutting forces has been analysed completely.Firstly,the movement of an arbitrary point P at the side cutting edge could be decomposed to cylinder helical move-ment (reference movement)and circular movement perpen-dicular to the cutter axis,as depicted in Fig.1.The reference movement can be decomposed to horizontal tangential feed and perpendicular axial feed,shown in Fig.2.The horizon-tal velocity of point P is defined as v P 0v PO +v O ,where v O is identical to f vt .For Ω>>Ωh ,means |v PO |>>|v O |,and therefore,v P ≈v PO .The influence of horizontal tangential feed on the side edge cutting force can beignored.Fig.2Kinematics of the side cuttingedgeFig.3Kinemics of the end cutting edgeSecondly,axial feed f vz may result in a portion of the axial cutting force being on the side edge.For every axial feed,the cutting volume of the side edge is proportional to f za a e h i ,but the cutting volume of the end edge is proportional to f za a e p ΦB ÀD m ðÞ=sin θ.That means that the side edge undergoes intermittent cutting while the end edge undergoes continuous cutting.In the same time period,the cutting force derived from axial feed on the side edge is much smaller than that on the end side.So,the influence of axial feed on the side edge cutting force can also be ignored.Then,assuming the top points on an end cutting edge in a straight line,the radial distance of point P to the cutting axis is variable.The influence of the horizontal feed f vt is more outstanding when P is near to the axis.The horizontal movement of point P at the end edge can be decomposed into the relative tangential part v t and relative radial part v r ,as described in pared to drilling or plunge milling operations in which tangential cutting forces are vanished andtangential velocity of the z -axis is zero,tangential forces and axis tangential velocity of the helical milling are not zero,as depicted in Fig.4.For the aforementioned reason,the influence of horizontal tangential feed on end edge cutting forces can be ignored.The existence of the relative radial part v r of the end edge implies that the radial force also exists.If we consider the end cutting edge of the flat-end milling cutter as approximately a straight line,the cutting edge along the radial direction slides rather than shears.F r *should be the friction force that is smaller than the shear force.Therefore,the radial force onthe end edge can be neglected,or F Ãa ¼0.Finally,due to the axial feed associated with f va ,the dis-placement direction of the end edge is not horizontal but having an angle θrelative to f va and f vz .After calculating this angle,the actual direction of the machined surface,the variation of the rake angle and the clearance angle can be defined.The cutting force on the end edge derived from axial feed can be defined within the plane to which the machined surface belongs.3.2Side cutting edgeBased on the kinematics of the helical milling process,two new features that may influence the cutting force and dynamics of the helical milling process have been considered.One was the periodic force variation created by the circular or tangential feed of the tool,and the other is the additional force component generated by the axial feed of the tools.The axial feed force mostly occurs at the end cutting edge of the milling tools.The interaction conditions between the tool and the workpiece are the combination of side edge cutting forces and end edge cutting forces.F !¼F !i þF!Ãð8ÞWhere,F !i is the side cutting edge component and F !Ãis end cutting edge component.Considering a point P on the (jth)Fig.4Horizontal feed influence to forces on end cuttingedgesFig.5Cutting forces on the side cutting edgecutting tooth,shown in Fig.5,the integration cutting force F !i(defined in the local coordinate system)along the in-cut por-tion of the flute j is similar to that presented in the referenced literature [4].F i x ;jϕj ðz ÞÀÁ¼f zt 4k b ÀK tc cos2ϕj ðz ÞþK rc 2ϕj ðz ÞÀsin2ϕj ðz ÞÀÁÂÃþ1k b K te sin ϕj ðz ÞÀK re cos ϕj ðz ÞÂÃ&'ϕj ;z z j ;1ðÞϕj ;z z j ;1ðÞð9ÞF iy ;j ϕj ðz ÞÀÁ¼Àf zt 4k b K tc 2ϕj ðz ÞÀsin2ϕj ðz ÞÀÁþK rc cos2ϕj ðz ÞÂÃþ1k b K te cos ϕj ðz ÞþK re sin ϕj ðz ÞÂÃ&'ϕj ;z z j ;1ðÞϕj ;z z j ;1ðÞð10ÞF iz ;jϕj ðz ÞÀÁ¼1k bK ac f zt cos ϕj ðz ÞþK ae ϕj ðz ÞÂÃϕj ;z z j ;1ðÞϕj ;z zj ;1ðÞð11Þwhere k b ¼2tan b D m=The detail of the integration of these forces is complicated because the contours of the side edge of the generic milling cutter are helical circles.To get the details of the forces at an arbitrary time,the integration procedure at one period (e.g.from zero to 2π)of the forces on the discrete cutter has to beFig.6Different intervals of a cutting period.a a p >Φex ÀΦst ðÞ=k b ,b a p <Φex ÀΦst ðÞ=k bFig.7Cutting forces on theend cutting edgedivided into several time intervals,as shown in Fig.6.The oblique lines represent the unfolding of the milling tool flutes in a plane.If a p >Φex ÀΦst ðÞ=k b is as shown in Fig.6a ,axial cutting depth is large.Φst and Φex is the cut-in and cut-out relative rotational angle of the cutter,respectively.0.0050.010.0150.020.0250.030.035−1,500−7500750bTime (sec)F o r c e (N )0.0050.010.0150.020.0250.030.035−1,500−75007501500Time (sec)F o r c e (N )Cutting force of Side edge No. 20.0050.010.0150.020.0250.030.035−1,500−75007501,500Time (sec)F o r c e (N )Result Cutting force of Side edges−4000−2000020004000Time (sec)F o r c e (N )Cutting force of End edge No. 1−4000−2000020004000Time (sec)F o r c e (N )Cutting force of End edge No. 20.0050.010.0150.020.0250.030.035−20000200040006000Time (sec)F o r c e (N )Result Cutting force of End edgesFig.8Simulation of the cutting forces during helical milling (milling tool diameter D m 16mm,five flutes,cutting speed v c 100m/min,axial feed rate per tooth f za 0.2mm,tangential feed rate per tooth f zt 0.5mm,radial cutting depth a e 8mm,up milling)In intervals 1and 5,there are no interactions between the cutter and workpiece,and therefore,the cuttingforce 0 ϕj <Φst ;F !j ¼0;Φq ϕj <2p ;F !j ¼0During interval 2,the cutting tooth begins to cut into the workpiece,where Φst ϕj <Φex ;ϕj z 1ðÞ¼ϕj ;ϕj z 2ðÞ¼ΦstDuring interval 3,the cutting tooth is fully involved in cutting the workpiece until the maximum axial cutting depth a p ,where Φex ϕj <Φp ;ϕj z 1ðÞ¼Φex ;ϕj z 2ðÞ¼Φst is obtained.During interval 4,the cutting tooth completes the cutting and quits the interaction finally,where Φp ϕj <Φq ;ϕj z 1ðÞ¼Φex ;ϕj z 2ðÞ¼ϕj ÀΦp ÀΦst ðÞIf a p <Φex ÀΦst ðÞ=k b as shown in Fig.6b ,axial cutting depth is large.In interval 1and 5,there is no interaction between the cutter and workpiece,and therefore no cutting force.0 ϕj <Φst ;F !j ¼0;Φq ϕj <2p ;F !j ¼0During interval 2,the cutting tooth begins to cut into the workpiece and progress towards the maximum axial cutting depth a p ,where Φst ϕj <Φp ;ϕj z 1ðÞ¼ϕj ;ϕj z 2ðÞ¼ΦstDuring interval 3,the cutting tooth interacts with the workpiece with a p ,where Φp ϕj <Φex ;ϕj z 1ðÞ¼ϕj ;ϕj z 2ðÞ¼ϕj ÀΦp ÀΦst ðÞDuring interval 4,the cutting tooth completes the cutting operation and quits the interaction finally,where Φex ϕj <Φq ;ϕj z 1ðÞ¼Φex ;ϕj z 2ðÞ¼ϕj ÀΦp ÀΦst ðÞ3.3End cutting edgeSince both the tangential feed f vt and axial feed f va are present during helical milling,the end cutting edge force component and the edge of these teeth are assumed to be a straight line and coincide with the radial line during analysis.If the friction force is neglected along the endcutting edge,the radial force F Ãa ¼0.As shown in Fig.7,the end cutting edge force component can be represented asd F Ãv¼K Ãvc f za cos θd r þK Ãve d r ð12Þd F Ãn ¼K Ãnc f za cos θd r þK Ãne d rð13Þd F Ãt ¼d F Ãv cos θÀd F Ãn sin θð14Þd F Ãa¼d F Ãv sin θþd F Ãn cos θð15Þd T Ã¼r d F Ãt ð16Þ00.0050.010.0150.020.0250.030.035−5000Time (sec)F o r c e (N )00.0050.010.0150.020.0250.030.035−50005000Time (sec)F o r c e (N )Cutting force of cutting edge No. 20.0050.010.0150.020.0250.030.035−20000200040006000Time (sec)F o r c e (N )Result Cutting force of milling toolFig.8(continued)Denote A ¼N m f za 2p ,B ¼N m f zt cos ϕj 2p ,θ¼argtan v av t¼argtan A r þB ,Θ½ ¼R D m 2D m 2Àa eÃd r cos θÀsin θ0sin θcos θ0000r cos θÀr sin θ026643775;K Ã½ ¼K Ãvc K Ãve K ÃncK Ãne K ÃrcK Ãre2435,therefore,F Ãt ;j F Ãa ;jF Ãr ;j T Ãj8>><>>:9>>=>>;¼Θ½ K Ã½ f za 1&'ð17ÞTransform to the local coordinate,F Ãx ;j F Ãy ;j F Ãz ;j T Ãj 8>><>>:9>>=>>;¼ÀF Ãt ;j cos ϕj ðt ÞÀÁF Ãt ;j sin ϕj ðt ÞÀÁF Ãa ;j T Ãj8>><>>:9>>=>>;ð18ÞSum up side cutting edge forces and end cutting forces onthe j th tooth and convert to global coordinates.F x ;j F Y ;j F Z ;j T Z ;j 8>><>>:9>>=>>;¼cos Ωh t sin Ωh t00Àsin Ωh tcos Ωh t 0000100126643775F i x ;j þF Ãx ;j F i y ;j þF Ãy ;j F i z ;j þF Ãz ;j T Ãj8>><>>:9>>=>>;ð19ÞThen,sum up all the cutting forces on the cutting teeth toobtain the cutting force model.246810−400400Time (sec)F o r c e (N )Experimental Cutting Force of X directionab246810−400400Time (sec)F o r c e (N )Experimental Cutting Force of Y direction0200400Time (sec)F o r c e (N )Experimental Cutting Force of Z direction−4000400Time (sec)F o r c e (N )Simulate Cutting Force of X direction246810−4000400Time (sec)F o r c e (N )Simulate Cutting Force of Y direction0200400Time (sec)F o r c e (N )Simulate Cutting Force of Z directionFig.9Cutting force result from experiment and simulation during helical milling cutting (milling tool M.A.Ford 20-mm five-flute end mill 17878703A,cutting speed v c 100m/min,axial feed rate per toothf za 0.005mm,tangential feed rate per tooth f zt 0.1mm,radial cutting depth a e 1mm,down milling)12345678x 10−3−300−200−100100200300400Time (sec)F o r c e (N )Experimental cutting force of single tooth periodcd12345678x 10−3−300−200−100100200300400Time (sec)F o r c e (N )Simulation cutting force of single tooth periodFig.9(continued)F X F Y F Z T Z8>><>>:9>>=>>;¼X N m j ¼1F X ;j Ωt þj À1ðÞ2pN ÀÁF Y ;j Ωt þj À1ðÞ2p N ÀÁF Z ;j Ωt þj À1ðÞ2pN ÀÁT Z ;j Ωt þj À1ðÞ2p NÀÁ8>><>>:9>>=>>;ð20ÞThe cutting force model during helical milling operationsin the time domain has therefore been established analyti-cally.This model defines both the cutting force on the side cutting edge and on the end cutting edge,incorporating the interactions between the cutter and the workpiece on the effect of the spindle rotation and the helical feed.4Simulations and experimental resultsCutting forces during helical milling have been simulated on the MATLAB platform using the models presented previ-ously,and experiments have been performed to compare with the model predictions.The process parameters includ-ed the workpiece material,cutting conditions,tool material and geometry.The Ti6Al4V alloy was cast and then HIPed (hot isostatic pressing,HIP)at a pressure of 100–140MPa at 920°C for 2.5h;then,the casting was rough milled to the end geometry (160×160×20mm)with a hole in a diameter of 60mm in the centre of the plate as shown in Fig.1.There were two types of cutting tools,the M.A.Ford 20-mm five-flute carbide end mill (17878703A)and the M.A.Ford 16-mm five-flute carbide end mill (17862903A).Experiments were carried out on a five-axis high-speed Mikron UCP-710CNC machining centre.A three-axis piezo-electric Kistler 9265B type dynamometer was set up on the fixture with the workpiece.The accessory data ac-quisition system of the dynamometer consisted of a Kistler 5019A type multi-channel charge amplifier and signal pro-cessing software DynoWare.Before commencing the experiments,the dynamometer was calibrated using static loads.The simulated cutting forces in an entire milling tool revolution on the side edges,end edges and whole cutter during the typical cutting conditions are depicted in Fig.8.In this simulation,the up milling and large radial cutting depth are considered as the significant characteristics of the operation.Figure 8a shows the simulated cutting forces that acted on first side cutting edge,second side cutting edge and cutting forces that acted on the milling tool from both the five cutting edges,respectively.For the up milling condi-tion,the j th edge engages with the workpiece,and the (j -1)th edge engages following.The large radial cutting depth means that before the previous cutting edge has completed cutting,the next cutting edge has engaged the workpiece.Therefore,there is a period of time that forces overlap between the consecutive cutting edges.Figure 8b shows thesimulated cutting forces that acted on the end cutting edges.There are similar cutting forces superposing between consec-utive side cutting edges.However,the sum of the X ,Y direc-tion forces are nearly zero,that is one of the important features of helical milling and plunge milling operations.Figure 8c shows the cutting forces that acted on the milling tool.These results are the integration of the component forces from Fig.8a and b .The simulated and experimental cutting force results are compared in Fig.9.In this case,cutting tools travel along an entire helical curve and machine an entire helical milling period.The X ,Y ,Z coordinates are fixed to the workpiece,during the helical feed motion of the tool,the amplitude of F X and F Y change with time following a sine relationship.The amplitude of Fig.9a and b counter profile is the maximum result of F X and F Y .Figure 9c and d shows the experimental and simulated cutting forces in detail in a single tooth period.The comparison result from experiment and simulation are shown in Table 1.This figure depicts the simulation results to an accuracy of about 10%in these selected indicators.The maximum value of F X ,F Y and F Z indicates for a single tooth period for both simulation and experimental results shown.The maximum of ﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃF 2X þF 2Yp indicates the amplitude of force of F X and F Y during helical milling.The errors probably result from cutting tool deflection and cutting tool wear.5ConclusionIn this paper,cutting forces during helical milling operations have been modelled in the time domain.The cutting forces both on the side cutting edges and on the end cutting edges along the helical feed path have been modelled by considering the tangential and the axial motion of the tool.The cutting force model can be used to predict cutting forces both on the side cutting edges and the end cutting edges.The model can also predict forces on the whole helix milling tool considering the process parameters and tool geometry.The experimental results show that for the given helix milling operation param-eters,the result of simulation predicts the cutting forces effec-tively and accurately.Table 1Comparison of experiment and simulation resultsExperiment (average)SimulationErrorHelical feed period (s)9.509.4750.263%Maximum of F X (N)371.1341.2−8.06%Maximum of F Y (N)253.2283.211.8%Maximum of F Z (N)287.7269.4−6.36%Maximum of ﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃF 2X þF 2Yp (N)365.3397.68.84%。

钛合金tc11动态拉伸力学行为的实验研究

钛合金TC11动态拉伸力学行为的实验研究张军, 汪洋（中国科学技术大学近代力学系中科院材料力学行为和设计重点实验室安徽合肥 230027）摘要：利用MTS809材料试验机和旋转盘式间接杆杆型冲击拉伸实验装置，对双态组织两相钛合金TC11进行了应变率为0.001 s-1的准静态和190s-1的动态单向拉伸实验，获得了TC11等温和绝热拉伸应力-应变曲线；实施了应变率为190s-1的冲击拉伸复元实验，获得了TC11在高应变率下的等温应力-应变曲线。

试验结果表明，TC11的拉伸力学行为具有明显的应变硬化效应、应变率强化效应和绝热温升软化效应。

采用修正的Johnson-Cook模型较好地表征了TC11在试验应变率范围内的拉伸力学行为。

关键词：两相钛合金；动态拉伸；绝热温升软化；复元试验EXPERIMENTAL INVESTIGATION ON THE DYNAMIC TENSION BEHA VIOR OFTITANIUM ALLOY TC11Jun Zhang, Yang Wang(Department of Modern Mechanics, CAS Key Laboratory of Mechanical Behavior and Design of Materials, University of Science and Technology of China, Hefei, Anhui 230027, PR China)Abstract:Quasi-static and dynamic uniaxial tension tests for a titanium alloy TC11 with a duplex microstructure were performed using MTS809 testing system and rotating disk bar-bar tensile impact apparatus, respectively. The isothermal stress-strain curve at 0.001s-1 and the adiabatic stress-strain curve at 190s-1 were obtained. The dynamic tensile recovery test was carried out at the rate of 190s-1 and the isothermal stress-strain curve at the high strain rate was obtained. The experimental results indicate that there exists the strain hardening, strain-rate strengthening and adiabatic temperature-rise softening phenomenon in the tension behavior of TC11. A modified Johnson-Cook model was chosen to describe the tensile behavior of TC11 at different strain rates. The model results agree well with the experimental data.Keywords: Two phase titanium; Dynamic tension; Thermal softening; Recovery test0. 引言两相钛合金具有比强度高、高低温性能优异、耐腐蚀等优点，是航空、航天工程中广泛使用的结构材料。

【研究】拟推荐2017年度教育部高等学校科学研究优秀成果奖科学技

【关键字】研究拟推荐2017年度教育部高等学校科学研究优秀成果奖（科学技术）项目情况项目名称：高原高压低氧特殊环境下火灾扩散燃烧行为的基础理论研究主要完成人：胡隆华，杨立中，汪箭，方俊，陆守香，王强推荐单位：中国科学技术大学申报奖种：高等学校自然科学奖项目简介：本项目针对我国特殊的高原地理条件，研究揭示了高原高压低氧特殊环境下火灾扩散燃烧行为规律，提出了高原高压低氧特殊环境下可燃物的热解与着火、热反馈与燃烧速率、火焰行为特征与稳定性的基础理论，重要科学发现包括：1. 在可燃物热解与着火方面，发现并定量揭示了高原高压低氧环境下热解失重速率更大、更容易着火的特性（着火时间更短、着火临界热流更低），建立了高压条件下考虑热解气在炭层多孔介质中输运特性的固相热解模型，揭示了热解气辐射阻隔效应对着火临界的影响机制，提出了基于counter-flow耦合压力效应的固体可燃物热解与气相着火新理论模型。

2. 在可燃物燃烧热反馈与燃烧速率方面，揭示了火灾中可燃物燃烧的传导、对流和辐射热反馈在高原高压低氧环境的特殊演化行为，发现了高原高压低氧环境下不同尺度固体和液体可燃物的燃烧速率和火蔓延速率与常压环境的差异及物理机制，建立了不同热反馈主控机制下耦合压力和火源尺度效应的火灾燃烧速率与火蔓延理论模型。

3. 在火焰行为特征与稳定性方面，发现了高原高压低氧环境下火灾湍流扩散火焰特征参数（火焰高度、中心线温度、碳黑辐射等）、以及火焰推举和吹熄等失稳行为与常压环境的差异，并揭示了其中的物理化学耦合机制，建立了高原高压低氧环境下火焰卷吸及其特征参数模型，提出了基于Damköhler 数耦合压力效应的火焰推举和吹熄临界理论模型。

本项目揭示了高原高压低氧环境下火灾的特殊扩散燃烧行为规律，并系统建立了相关基础理论，共发表SCI论文62篇，其中12篇发表在国际燃烧领域两大权威期刊Combustion and Flame 和Proceedings of the Combustion Institute。

建环专家简介

李先庭：主要研究方向主要从事室内空气环境数值模拟、应急通风、制冷系统与蓄能系统研究与开发。

在室内空气环境数值模拟方向，主要从事室内空气环境数值模拟方法、室内空气环境评价方法及应用研究。

数值模拟方法包括计算流体力学方法和区域模型方法，研究对象包括各种不同类型的风口和末端、各类热源和辐射、各种污染源的建模与计算，带回风的通风系统模拟等；室内空气环境评价方法包括送风有效性评价（如空气龄、换气效率等）、污染物排除有效性（如排空时间、排污效率、污染物年龄等）和热舒适三个方面，涉及稳态环境和动态环境；应用方向主要包括各种类型的高大空间（如体育场馆、候机大厅、剧场、中庭、厂房等）、洁净空间（如手术室、GMP车间、电子洁净厂房等）、运载工具（如车辆、舰船、飞机等）等。

在应急通风方向，主要研究有限时间条件下污染物的传播规律和应急通风系统的设计方法与运行控制，主要针对三种用途：（1）火灾防排烟；（2）公共场所有毒气体释放；（3）利用应急通风系统输送麻醉气体控制劫持现场，解救人质。

在制冷系统方向，主要从事小型制冷系统的模拟仿真与设备研究（包括各类房间空调器、汽车空调系统、舰船空调系统等）、制冷机组研究与开发。

研究内容包括制冷系统部件特性、系统优化匹配、运行与优化控制等。

在蓄能系统方向，主要进行冰蓄冷和水蓄冷系统的优化设计与控制研究，蓄冰槽的研究与开发，蓄冷蓄热水池/罐的设计与开发，以及土壤蓄能和季节性蓄冷系统的研究，食品冷藏运输研究等。

教学工作主要从事“室内空气流动数值模拟”、“暖通空调课程设计”、“人工环境学”等教学。

论文和论著参编普通高等教育“十五”国家级规划教材《建筑环境学》，申请或授权发明专利16项，发表科研论文百余篇，其中英文期刊40篇，国际会议篇28篇，中文期刊50篇．被SCI收录论文29篇，EI收录65篇。

部分代表性论文如下：1.Changqing Tian，Xianting Li，Chunpeng Dou. Experimental investigation and numericalsimulation on the adjustment dead zone of a variable displacement wobble plate compressor. International Journal of Refrigeration, v 28, n 7, Nov 2005, p 988-9962.Xianting Li, Zhen Yu，Bin Zhao, Ying Li. Numerical analysis of outdoor thermalenvironment around buildings. Building and Environment, v 40, n 6, June 2005, p 853-8663.Xianting Li, Jianrong Yang, Wei Sun. Strategy to optimise building ventilation toaid rescue of hostages held by terrorists. Indoor and Built Environment, v 14, n 1, 2005, p 39-504.Xianting Li, Bin Zhao.Accessibility: a new concept to evaluate the ventilationperformance in a finite period of time. Indoor and Built Environment, v 13, n 4, Aug 2004, p 287-2935.Dongning Li, Xianting Li, Yunzhi Guo, Jianrong Yang, Xudong Yang. Generalizedalgorithm for simulating contaminant distribution in complex ventilation systems with recirculation. Numerical Heat Transfer Part A-Applications,v 45, n 6, April 2004, p 583-5996.Jianrong Yang, Xianting Li, Bin Zhao. Prediction of transient contaminant dispersionand ventilation performance using the concept of accessibility.Energy and Buildings, v 36, n 3, March 2004, p 293-2997.Xianting Li, Dongning Li, Xudong Yang, Jianrong Yang. Total air age: An extensionof the air age concept. Building and Environment, v 38, n 11, Nov 2003, p 1263-1269 8.Bin Zhao, Xianting Li, Dongning Li, Jianrong Yang. Revised air-exchange efficiencyconsidering occupant distribution in ventilated rooms. Journal of the Air and Waste Management Association, v 53, n 6, Jun 2003, p 759-7637．荣誉、奖励及参加学术团体的情况获省部级科技进步奖2项（北京市科学技术进步三等奖，2001年；教育部科技进步成果二等奖，2000年），现为北京制冷学会副理事长，国际制冷学会E2专业委员会委员，中国制冷学会第五专业委员会委员，全国暖通空调及净化设备标准化技术委员会委员，《建筑热能通风空调》、《制冷与空调》和《制冷空调与电力机械》编审委员会委员，ASHRAE(美国供热、通风、制冷与空调工程师学会)会员、中国制冷学会高级会员朱颖心：主要研究方向主要研究领域与项目1.可持续建筑研究(1)科技部国家“十五”攻关项目：绿色建筑关键技术研究课题一：绿色建筑的规划设计导则和评估体系研究课题五：降低建筑物能耗的综合关键技术研究(2)科技部奥运科技十大专项课题：绿色奥运建筑评估体系；(3)北京市科委项目：新建建筑能耗评估体系与超低能耗示范建筑的建立与实践研究；(4)国家自然科学基金重点项目：住区微气候的热物理问题。

供热管网中阀门工作流量特性实验研究与数值模拟

38∣建设科技研究探讨建设科技CONSTRUCTION SCIENCE AND TECHNOLOGY2018年11月上总第371期1 引言阀门作为管道工程中的一个重要部件，被广泛应用于生产和生活中[1]。

城市集中供热管网中需要借助阀门来实现对流量和压力的调节。

在集中供热二次管网调节时，管理人员对于很多常用阀门的调节都还依靠手动调节，且调节开度的大小也都是凭经验和感觉，由此不仅起不到良好的调节效果，而且也不利于节能。

在供热系统的实际运行中，并不是所有的阀门都处供热管网中阀门工作流量特性实验研究与数值模拟杨华1 李少雄1 徐疆2 张荷2 王向伟2 马婉路1（1.河北工业大学能源与环境工程学院，天津 300401；2.承德热力集团有限责任公司，河北承德 067000）[摘要]供热管网中阀门的工作流量特性对管网的运行效果有着直接的影响。

本文通过搭建仿真试验台，研究了供热系统二次管网中几种常用阀门的工作性能，并分析了不同阀门有效调节范围及调节能力的大小。

然后借助FloMAS-TER 软件建立了热源至热用户模型，针对蝶阀依据其技术参数对软件中阀门控制曲线参数进行设置,并对供热二次管网运行规律进行数值模拟，仿真计算出各个管段流量分配。

结果表明，数值模拟结果有较高的准确性。

[关键词]阀门; 流量特性；损失系数； FloMASTER 模拟Experimental Investigation and Numerical Simulation of ValveFlow Characteristics in the Heating NetworksYang Hua 1 , Li Shaoxiong 1, Xu Jiang 2 , Zhang He 2, Wang Xiangwei 2, Ma Wanlu 1(1. School of Energy and Environmental Engineering of Hebei University of Technology ，Tianjin, 300401，China;2. Chengde Heating Group Co .，Ltd .，Chengde, 067000，Hebei ，China)Abstract : The operating flow characteristics of the valve in the heating pipe network have a direct impact on the operation of the pipe network. By setting up a simulation test bench, the work flow characteristics of several common valves in the secondary pipe network of the heating system were studied, and the effective adjustment range and adjustment capacity of different valves were analyzed. Then the FloMASTER software was used to establish the heat source-to-heat user model. According to the technical parameters of the butterfly valve, the valve control curve parameters in the software were set, and the operation rules of the heating secondary pipe network were numerically simulated, and the flow distribution of each pipe section was simulated. The results showed that the numerical simulation results had higher accuracy.Keywords : valve, flow characteristic, loss coefficient, FloMASTER simulation于全开状态,而非全开阀门的阻力系数是全开时的几十甚至上百倍，所以其中的压力损失不能忽略。

低浓度甲缩醛水溶液-萃取剂液液相平衡数据的测定与关联

低浓度甲缩醛水溶液-萃取剂液液相平衡数据的测定与关联时米东;王利平;何高银;于雪敏;李青松【摘要】在30℃、40℃和大气压力下测定了水-甲缩醛-对二甲苯和水-甲缩醛-甲苯体系的液液相平衡数据.根据实验数据对分配系数和分离因子进行了估算,结果表明两种萃取剂对甲缩醛均具有优异的萃取能力.采用Othmer-Tobias和Hand方程对实验数据的可靠性进行了分析,两个方程的线性相关度R2均在0.99以上,表明实验数据具有良好的可靠性.另外,采用UNIQUAC和NRTL热力学模型对数据进行了关联,结果表明估算值与实验值具有良好的一致性.%Liquid-liquid equilibria (LLE) data for the ternary systems {water + methylal + (p-xylene or toluene)} were measured under atmospheric pressure at 30 and 40℃.The distribution coefficient(D)and separation factor(S)were calculated according to the experimental data, which indicated that p-xylene and toluene showed excellent extraction capacity for methylal. Reliability of the experimental data was determined by Othmer-Tobias and Hand correlation equations with the linear correlation of R2>0.99.Moreover, both UNIQUAC and NRTL models were adopted to correlate the experimental LLE data. The calculated values by the two models showed a good consistency with the measured data.【期刊名称】《天然气化工》【年(卷),期】2018(043)001【总页数】7页(P34-39,46)【关键词】液液相平衡;甲缩醛;对二甲苯;甲苯;水【作者】时米东;王利平;何高银;于雪敏;李青松【作者单位】中国石油大学(华东)化学工程学院重质油国家重点实验室,山东青岛266580;中国石油大学(华东)化学工程学院重质油国家重点实验室,山东青岛266580;中国石油大学(华东)化学工程学院重质油国家重点实验室,山东青岛266580;中国石油大学(华东)化学工程学院重质油国家重点实验室,山东青岛266580;中国石油大学(华东)化学工程学院重质油国家重点实验室,山东青岛266580【正文语种】中文【中图分类】O642甲缩醛是一种重要的溶剂和化工原料[1-5]，可由甲醇与甲醛之间的羟醛缩合得到[6-8]。

experimental and clinical study

experimental and clinicalstudyExperimental and clinical study is an important aspect of modern medical research. It involves systematic investigation and analysis of medical phenomena through controlled experiments andclinical trials. Experimental and clinical study is essential to understanding the underlying mechanisms of various medical conditions, as well as developing effective diagnostic and treatment strategies.Experimental study involves the use of laboratory experiments to investigate the causal relationships between various medical factors. This type of study is particularly useful when trying to identify the underlying biological andphysiological mechanisms of a particular medical condition. Experimental study often involves the use of animal models, cell cultures, and biochemical assays to investigate the effects ofvarious interventions, such as drugs or surgical procedures.One key advantage of experimental study is its ability to isolate and control for various confounding factors. By carefully controlling and manipulating the experimental conditions, researchers can identify the specific factors that contribute to a particular medical condition. This information can then be used to develop targeted interventions to prevent or treat the condition.Clinical study, on the other hand, involves the observation and analysis of patients in clinical settings. This type of study is particularly useful when trying to test the effectiveness and safety of new diagnostic or treatment strategies. Clinical study often involves the use of randomized controlled trials, in which patients are randomly assigned to different treatment or control groups.One advantage of clinical study is its ability to directly observe and measure the effects of medical interventions in real-world settings. Clinical study can also be used to investigate theeffects of medical interventions on a variety of patient populations, including those with different medical histories, ages, and genders.Experimental and clinical study are often used together in medical research. For example, experimental studies may be used to identify potential drug targets, while clinical studies may be used to test the safety and effectiveness of those drugs in human patients. Experimental and clinical study can also be used to identify biomarkers, or measurable indicators of aparticular medical condition, which can be used to diagnose or monitor the condition.Another important aspect of experimental and clinical study is data analysis. Researchers must carefully analyze the data collected from experiments and clinical trials to ensure thattheir findings are valid and reliable. Statistical analysis is used to identify meaningful patterns and relationships in the data, and to test hypotheses about the underlying mechanisms of medical conditions.In summary, experimental and clinical study is a critical aspect of medical research. Through controlled experiments and clinical trials, researchers can identify the underlying mechanisms of medical conditions and develop effective diagnostic and treatment strategies. By carefully analyzing the data collected from experiments and clinical trials, researchers can ensure that their findings are valid and reliable. Experimental and clinical study will continue to play a vital role in advancing medical science and improving the health and well-being of people around the world.。

水泥砂浆的一个热粘弹性率型损伤本构模型

水泥砂浆的一个热粘弹性率型损伤本构模型陶俊林;李奎【摘要】利用SHPB实验系统及自行研制的混凝土类材料快速高温加热设备,对水泥砂浆试件进行了不同温度(20～600℃)和3种冲击速度下的实验,得到了不同温度和冲击速度下水泥砂浆试件的应力应变关系曲线.基于ZWT粘弹性本构模型,并且考虑高温下水泥砂浆损伤演化规律都服从Weibull分布,提出了一个水泥砂浆的热粘弹性率型损伤本构模型.通过数据拟合,获得了本构模型的相关参数,结果表明:理论预测和实验结果吻合良好.%Based on a 50-mm-diameter split Hopkinson pressure bar (SHPB) system and a self-developed concrete-like material rapid heating device, impact compressive experiments were carried out. The stress-strain curves at different temperature and three kinds of impact velocities were obtained. Based on the ZWT visco-elastic constitutive model and damage evolution of cement mortar following Weibull distribution, the damage thermo-viscoelastic constitutive equation at high strain rates was proposed. Through the numerical fitting, the parameters of the constitutive equation were obtained. Comparisons show that the model predictions agree well with the experiment results.【期刊名称】《爆炸与冲击》【年(卷),期】2011(031)003【总页数】6页(P268-273)【关键词】固体力学;本构模型;SHPB;水泥砂浆;温度;损伤;粘弹性【作者】陶俊林;李奎【作者单位】西南科技大学土木工程与建筑学院,四川绵阳621010;西南科技大学土木工程与建筑学院,四川绵阳621010【正文语种】中文【中图分类】O347.3混凝土类材料是一种重要的结构工程材料，已被广泛应用于高层建筑、工业厂房、桥梁、大坝等工程。

扬州市人民政府关于2017-2019年度扬州市自然科学优秀学术论文评选结果的通报

扬州市人民政府关于2017-2019年度扬州市自然科学优秀学术论文评选结果的通报
文章属性
•【制定机关】扬州市人民政府
•【公布日期】2021.03.05
•【字号】扬府发〔2021〕16号
•【施行日期】2021.03.05
•【效力等级】地方规范性文件
•【时效性】现行有效
•【主题分类】科技成果与知识产权
正文
市政府关于2017-2019年度扬州市自然科学优秀学术论文评选结果的通报-市政府文件-政府办
各县（市、区）人民政府，经济技术开发区、生态科技新城、蜀冈—瘦西湖风景名胜区管委会，市各委办局（公司），市各直属单位：
为提高我市自然科学学术水平，鼓励全市科技工作者开展学术创新服务发展，促进全市科技人才成长，2020年我市开展了2017-2019年度扬州市自然科学优秀学术论文评选工作，共征集到701篇论文，其中675篇论文通过资格审查。

经评审委员会评定，分理工类、农业类、医药类、管理教育类，共评出优秀学术论文269篇，其中一等等次33篇、二等等次101篇、三等等次135篇。

经研究，市政府决定对评选出的269篇优秀学术论文给予通报并颁发证书。

希望全市各级科技团体和广大科技工作者，以科技赋能产业为己任，积极开展学术研究，主动投身科技创新，脚踏实地，严谨求是，把“好地方”扬州建设得好上加好、越来越好做出积极贡献。

附件：2017-2019年度扬州市自然科学优秀学术论文评选结果
扬州市人民政府
2021年3月5日附件
2017-2019年度扬州市自然科学优秀学术论文评选结果
一等等次（33篇）
二等等次（101篇）
三等等次（135篇）。

大颗粒与壁面碰撞的离散单元法模拟与分析

大颗粒与壁面碰撞的离散单元法模拟与分析张鹤;李天津;刘马林;黄志勇;薄涵亮【摘要】颗粒与壁面碰撞普遍存在于散体物料输送过程,研究颗粒与壁面碰撞有助于优化输送系统、减小物料磨损或提高输送经济性.本文基于离散单元法(DEM),采用Hertz-Mindlin无滑移接触模型,对单个6 mm直径大颗粒与壁面碰撞进行了数值模拟和分析,研究了碰撞速度、碰撞角度和剪切模量对碰撞过程和法向最大接触力的影响.研究结果表明,Hertz-Mindlin无滑移接触理论描述的法向接触过程具有自相似特性,法向卸载时长与法向加载时长比值为定值.模拟的接触时长与Thornton等的关系式预测值相符.碰撞速度和碰撞角度对碰撞过程中的法向最大接触力均有明显影响,法向最大接触力随法向碰撞速度的增加近似线性增加;碰撞速度不变时,法向最大接触力随碰撞角度的增大而减小.剪切模量对法向接触力具有重要影响,在考虑颗粒磨损和破碎的DEM模拟时,不宜采用降低剪切模量加快计算速度.本研究对颗粒磨损和破碎研究以及高温气冷堆吸收球气力输送过程优化均具有重要意义.%Particle-wall collisions widely exist in bulk solidstransportation .Investiga-tions on particle-wall collisions are helpful to optimize transporting system ,decrease product attrition or improve transportation economy .Collisions between a single coarse particle (6 mm in diameter) and a wall were investigated by Hertz-Mindlin no-slip con-tact model based on discrete element method (DEM ) .Effects of impact velocity ,impact angle and shear modulus on contact processes and maximum normal contact forces were studied .Results show that the normal contact process described by Hertz-Mindlin no-slip contact model shows self-similarity feature ,and the ratio of unloading to loadingduration in normal direction keeps as a certain value .The numerical contact durations agree well with the predictions by the correlation of Thornton et al .The impact velocity and impact angle show obvious effects on maximum contact forces .The normal maxi-mum contact force increases almost linearly with normal impact velocity .For the certain impact velocity of 2 m/s ,the normal maximum contact force decreases with the increase of impact angle .The shear modulus is a key factor to normal contact force ,which sug-gests that speeding up DEM simulation by decreasing shear modulus should be avoided when particle attrition and/or breakage are in consideration .The results in the present study are important for investigation of particle attrition and/or breakage ,as well as optimization of absorber sphere pneumatic conveying process in high temperature gas-cooled reactor .【期刊名称】《原子能科学技术》【年(卷),期】2017(051)012【总页数】6页(P2212-2217)【关键词】颗粒碰撞;颗粒磨损;颗粒破碎;离散单元法;气力输送【作者】张鹤;李天津;刘马林;黄志勇;薄涵亮【作者单位】清华大学核能与新能源技术研究院 ,先进核能技术协同创新中心 ,先进反应堆工程与安全教育部重点实验室,北京 100084;清华大学核能与新能源技术研究院 ,先进核能技术协同创新中心 ,先进反应堆工程与安全教育部重点实验室,北京 100084;清华大学核能与新能源技术研究院 ,先进核能技术协同创新中心 ,先进反应堆工程与安全教育部重点实验室,北京 100084;清华大学核能与新能源技术研究院 ,先进核能技术协同创新中心 ,先进反应堆工程与安全教育部重点实验室,北京100084;清华大学核能与新能源技术研究院 ,先进核能技术协同创新中心 ,先进反应堆工程与安全教育部重点实验室,北京 100084【正文语种】中文【中图分类】TL33吸收球停堆系统是高温气冷堆的第二停堆系统[1]。

Experimental investigation and numerical simulation for weakening the thermal fluctuations Tjunction

Experimental investigation and numerical simulation for weakening the thermal ﬂuctuations in aT-junctionK.Gao a ,P.Wang b ,T.Lu a ,⇑,T.Song caCollege of Mechanical and Electrical Engineering,Beijing University of Chemical Technology,Beijing 100029,China bSchool of Energy and Power Engineering,Dalian University of Technology,Dalian 116024,China cChina Nuclear Power Technology Research Institute Co.,Ltd,Shenzhen 518124,Chinaa r t i c l e i n f o Article history:Received 25August 2014Received in revised form 17November 2014Accepted 4January 2015Available online 17January 2015Keywords:Experimental investigation Numerical simulation Tee junctionThermal ﬂuctuationa b s t r a c tIn this work,the mixing processes of hot and cold ﬂuids with and without a distributor are predicted by experiments and numerical simulations using large-eddy simulation (LES)on FLUENT platform.Temperatures at different positions of the internal wall and mixing conditions caused by T-junctions at different times are obtained,then the simulated normalized mean and root-mean square (RMS)temperature,velocity vector and temperature contour for the two structures,namely with and without a distributor,are compared.The results show that,compared with the a T-junction without a distributor,the mixing region of hot and cold water in the T-junction with distributor moves to the middle of the pipe,and the inclusion of the distributor reduces the temperature ﬂuctuations of internal wall noticeably and makes the mixing of hot and cold water more efﬁcient.Ó2015Elsevier Ltd.All rights reserved.1.IntroductionTee junction is a familiar structure that is universally used in pipeline systems of power plants,nuclear power plants and chemi-cal plants,it is often applied to mix hot and cold ﬂuid of main and branch pipes.The ﬂuctuations of ﬂuid temperature are transported to the solid walls by heat convection and conduction.This can cause cyclical thermal stresses and subsequent thermal fatigue cracking of the piping (Lee et al.,2009).So far,leakage accidents took place in several light water and sodium cooled reactors due to thermal fati-gue.In 1998,a crack was discovered at a mixing tee in which cold water from a branch pipe ﬂowed into the main pipe in the residual heat removal (RHR)system in a reactor in Civaux,France.Metallur-gical studies concluded that the crack was caused by a high degree of cycle thermal fatigue (Eric Blondet,2002).In 1990,sodium leak-age happened in the French reactor Superphenix (Ricard and Sperandio,1996).It has been established that mixing hot and cold sodium can induce temperature ﬂuctuations and result in thermal fatigue (IAEA,2002).Therefore,it is signiﬁcant to study how to weaken thermal fatigue of the piping wall to ensure the integrity and safety of the piping system in a nuclear power plant.In the analysis of thermal fatigue,temperature ﬂuctuation is a very important evaluation parameter.A reliable lifetime assess-ment of these components is difﬁcult because usually only thenominal temperature differences between the hot and cold ﬂuids are known,whereas the instantaneous temperatures and heat ﬂuxes at the surface are unknown (Paffumi et al.,2013).Kamaya and Nakamura (2011)used the transient temperature obtained by simulation to assess the distribution of thermal stress and fati-gue when cold ﬂuid ﬂowed into the main pipe from a branch pipe.Numerical simulation of ﬂow in the tee has been carried out Simoneau et al.(2010)to get temperature and its ﬂuctuation curves,and the numerical results were in good agreement with the experimental data.Through the analysis on thermal fatigue stress,it draw the conclusion that the enhanced heat transfer coef-ﬁcient and the temperature difference between hot and cold ﬂuids were primary factors of thermal fatigue failure of tees.Many numerical simulations and experiments have been carried out to evaluate the ﬂow and heat transfer in a mixing tee junction (Metzner and Wilke,2005;Hu and Kazimi,2006;Hosseini et al.,2008;Durve et al.,2010;Frank et al.,2010;Jayaraju et al.,2010;Galpin and Simoneau,2011;Aulery et al.,2012;Cao et al.,2012).Turbulent models such as Reynolds-averaged Navier–Stokes (RANS),Unsteady Reynolds averaged Navier Stokes (URANS),Scale-Adaptive Simulation (SAS),Reynolds stress model (RSM),detached eddy simulations (DES),and LES have all been used in industrial applications.As one of the choices of turbulent model for predicting the mixing ﬂow in tee junctions,the RSM can bemused to describe the momentum conservation of the mixing (Durve et al.,2010;Frank et al.,2010).Turbulent mixing phenomena in a T-junction have been numerically investigated using the k $x/10.1016/j.anucene.2015.01.0010306-4549/Ó2015Elsevier Ltd.All rights reserved.Corresponding author.based baseline Reynolds stress model(BSL RSM)(Frank et al.,2010) for two different cases.Durve et al.(2010)applied the RSM to pre-dict the velocityﬁeld of three non-isothermal parallel jetsﬂowing in an experiment setup used to simulate theﬂow occurring at the core outlet region of a fast breeder reactor(FBR),with a Reynolds number of1.5Â104.Theﬂow in tube of different Reynolds numbers (Re)andﬂow velocity ratio were studied experimentally with three-dimensional scanning using particle image velocimetry(3D-SPIV) (Brücker,1997).Large-eddy simulation(LES)is an alternative turbulence model with different subgridscale models often employed to predict velocity and temperatureﬂuctuations.Indeed many numerical studies have shown the capability of LES to model thermalﬂuctu-ations in turbulent mixing.LES was performed(Lee et al.,2009)to analyze temperatureﬂuctuation in the tee junction and the simu-lated results were in good agreement with the experimental data. Thermal striping phenomena in the tee junction had been numer-ically investigated using LES(Hu and Kazimi,2006)for two differ-ent mixing cases,and the simulated normalized mean and root-mean square(RMS)was consistent with experimental results. LES in a mixing tee were carried out(Galpin and Simoneau, 2011)in order to evaluate the sensitivity of numerical results to the subgrid scale model by comparing the experimental results, and to investigate the possibility of reducing theﬂuid computa-tional domain at the inlet.Another simulation that mixing of a hot and a coldﬂuid stream in a vertical tee junction with an upstream elbow main pipe was carried out with LES(Lu et al., 2013).And the numerical results show that the normalized RMS temperature and velocity decrease with the increases of the elbow curvature ratio and dimensionless distance.In the meantime,many scholars have studied how to weaken the thermalﬂuctuation.Experiments and simulation were con-ducted(Wu et al.,2003)on a tee junction geometry with a sleeve tube in it.Theﬂow is divided into three types of jets by theﬂow velocity ratio in main and branch pipes.Through the analysis of ﬂowﬁeld and velocityﬁeld of various jets types,it indicate that the addition of sleeve tube relieve the thermal shock caused by the coldﬂuid injection rge-eddy simulation have been used(Lu et al.,2010)to evaluate the thermal striping phe-nomena in tee junctions with periodic porous media,the temper-ature and velocityﬁeld inside the tubes are obtained.The research revealed that the addition of a porous reduces the tem-perature and velocityﬂuctuations in the mixing tube.As mentioned above,experiments and numerical simulations for both tee junction geometry with a sleeve tube in it(Wu et al., 2003)and for a mixing tee with periodic porous media in it(Lu et al.,2010)have been carried out.The results of previous researches provide a good reference value for this work that anal-yses the role of distributor in weakening the thermalﬂuctuation of internal piping wall,and this structure has not been studied to date,to the best of our knowledge.In this work,mixing processes have been studied by the experiment and numerically predicted with LES.Then the simulated normalized mean and root-mean square(RMS)temperature,velocity vector and temperature con-tour of the two tees are compared.2.Experiment systemThe Experimentﬂowchart is presented in Fig.1.The experimen-tal system consists of four main components,a cold water supply line,a hot water supply line,a test section,and a data acquisition unit.The experiment device is shown in Fig.2.Experimentﬂuid was adjusted to the desired temperature by the heater and chiller, and then was pumped to the test section.After mixing theﬂuid is returned to the heater for recycling,some of the excessﬂuid is dis-charged through the overﬂow pipe.During the mixing of theﬂuids, the temperature of the mixingﬂuid is collected and recorded by the thermocouple probe installed on the tube wall.The experiment requires two different structures of the test sec-tion,Fig.3is the T-junction section without the branch liquid dis-tributor and Fig.4is that with the branch liquid distributor.The addition of this structure has two main functions:(1)changing the mixing position of hot and coldﬂuids:moving the mixing zone to the middle of the tube,and away from the main pipe wall;(2) increasing the intensity of mixing process:adding the fence near the outlet of distributor enhanced the mixed disturbance and the exacerbatedﬂuid mixing of the inner tube.For the convenience of observing and adjusting the mixing process,the test section is a round pipe made of plexiglass,and other pipes are made of steel. Fig.5is the physical model of the branch liquid distributor.The test conditions in the present experiment are shown in Table1.We collected the instantaneous temperature data of every measurement points by the data collector.The distribution of sam-pling points are shown in Fig.6,there are total eight thermocou-ples in the circumferential direction at each plane.In the T-junction section without the branch liquid distributor,the number of the collected plane is6(x/d m=1,2,3,4,6,8).That is to say there are48thermocouples in the structure without distributor.And in the T-junction section with the branch liquid distributor,the num-ber of the collected plane is5(x/d m=2,3,4,6,8),which means there are40thermocouples in the structure that with the distrib-utor.In both structures,the distance between measuring point the thermocouple probe and the inner wall is30mm.Since the collect-ing frequency of the collector is limited,we use1Hz as the collect-ing frequency after theﬂowﬁeld is stable,and the total number of collection is800s.Table1shows the speciﬁc parameters of the test conditions.NomenclatureT time(s)Pr Prandtl numberLs mixing length of subgrid grid(m)T temperature(K)G acceleration of gravity(m/s2)K von Karman numberCs Smagorinsky numberS ij subgrid strain rate tensorM R momentum ratio of main pipe and branch pipe TÃnormalized mean temperaturesTÃrms normalized RMS temperaturesR d diameter ratioR v velocity ratiox,y,z axial coordinate(m)Greek symbolsqﬂuid density(kg/m3)b coefﬁcient of thermal expansionl viscosity(Pa s)ltturbulent viscosity(Pa s)k thermal conductivity(w/(m k))C P heat capacity(J/(kg°C))K.Gao et al./Annals of Nuclear Energy78(2015)180–187181182K.Gao et al./Annals of Nuclear Energy78(2015)180–1871\4\11-thermometers 2\5\10-pressure gauge 3\9-flow meter 6-c ooler 7-heater8-overflow 12-test sec tion 13-thermoc ouple data c ollec torFig.1.Experimentﬂow chart.Fig.5.Physical model of the branch liquid distributor(a)the whole graph(b)theproﬁle map.Fig.2.Experiment device of thermalﬂuctuation.Fig.3.Schematic diagram of the T-junction section without the branch liquid distributor.Fig.4.Schematic diagram of the T-junction section with the branch liquid distributor.3.Numerical simulationFig.7is the numerical model based on the experimental section of T junction.The size of the model,boundary conditions are con-sistent with the experiment.In which,hot water enters from the left of main pipe,and cold water enters from the branch pipe,ﬁnally the mixingﬂuidﬂow out of the right of the main pipe.Dur-ing the calculation,the steady results ofﬂowﬁeld and heat transfer are obtained by Reynolds stress model(RSM)ﬁrstly,and then set @q@tþ@q u i@x i¼0ð1Þ@q u i@tþ@q u i u j@x j¼À@ p@x iÀq0bðTÀT0Þgþ@@x jlþltÀÁ@ u i@x jþ@ u j@x i!ð2Þ@q T@tþ@q Tu j@x j¼@@x jkc p@T@x jÀq T00u00j!ð3ÞIn these equations,q,b,l,l t,k and c p represent the density,ther-mal expansion coefﬁcient,molecular viscosity,turbulent viscosity, thermal conductivity and speciﬁc heat capacity,respectively.The Smagorinsky–Lilly model is used for the turbulent viscosity,which is described as:lt¼q L2s j S jð4Þj S jTable1Experimental conditions.Main pipe Branch pipeFlow rate (m3/h)Temperature(K)Flow rate(m3/h)Temperature(K)Without distributor0.645304.650.270287.65With distributor0.645304.650.266287.65Fig.6.The distribution of sampling points on the planes.Physical model of T-junction(a)without the branch liquid distributor;(b)with the branch liquid distributor.K.Gao et al./Annals of Nuclear Energy78(2015)180–187183ij ¼12@ u i@x jþ@ u j@x ið7Þwhere k is the Von Karman constant of0.42;d is the distance to the closest wall;C s is the Smagorinsky constant of0.1;V is the volume of the computational cell.4.Results and discussionThe normalized mean and root-mean square temperature are used to describe the time-averaged temperature and temperature ﬂuctuation intensity.The normalized temperature is deﬁned as:Ã¼1NX Ni¼1TÃið8ÞN is the total number of sample times.TÃi¼T iÀT cT hÀT cð9Þwhere T i is the transient temperature,T c is the coldﬂuid inlet tem-perature and T h is hotﬂuid inlet temperature.The root-mean square(RMS)of the normalized temperature is deﬁned as:TÃrms¼ﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃ1X Ni¼1TÃiÀTÃ2rð10Þ184K.Gao et al./Annals of Nuclear Energy78(2015)180–187parison of experimental and numerical resultsAs can be seen from the Fig.8,the numerical normalized mean temperature distributions at the plane x/d m=1and the plane x/ d m=2are in good qualitative agreement and in adequate quantita-tive agreement,and most of them are within the experimental deviation of±20%.Meanwhile,the lifting trends of the data are the same.In the direction of180°,the mean temperatures are both minimal.And with the angle decrease to0°,the temperatures are gradually increased.Quantitative differences between the experi-ment and numerical results are that the normalized mean temper-atures given by LES are larger than the experimental data.That is because we did not add insulation unit on tube wall in the exper-iment,leading the transfer of some heat into the air.And in the process of numerical simulation,we ignored the convective heat transfer between the wall and the air.As shown in Fig.9,although the numerical results and experi-mental results have a little difference at the plane x/dm=2around the location of225°and the plane x/dm=2around the location of 0°and315°,all of them are within the error range that can be accepted.Both the simulations and experimental results give a lar-ger mean temperature in the top half of the main pipe than in the bottom half.This veriﬁes the validity of the LES model for predict-ing the mixing of hot and coldﬂuids in a tee junction.The normalized RMS temperature on the plane x/d m=1and plane x/d m=2are shown in Fig.10,respectively.Similar to the nor-malized mean temperature,the normalized RMS temperature lines agree very well with the experiment ones.Both of the maximum values appear at the bottom half of the pipe.This indicates that the maximum temperatureﬂuctuations of main pipe appear on the opposite of the branch pipe inlet in this condition.As shown in Fig.11,the numerical results and the experimental results have the same trend and the numerical data are agreed well with the experimental ones.By comparison with Figs.4and5,dif-ferent from the temperatureﬂuctuations distribution which with-out the branch liquid distributor,there are two peaks of high ﬂuctuation located at the90°and270°directions along with the tube.This is because the direction is that of the outlet of branch liquid distributor,the coldﬂuidﬂowing out from the outlet of branch liquid distributor mixes very fast with the hotﬂuid,leading to dramatic changes of temperature.In summary,the LES simulation results obtained are generally in good qualitative and quantitative agreement with the experi-mental data for the case of T-junction with/without the branch liquid distributor.Based on this,we analyzed the numerical results further.And the results are reported in the section below.4.2.Numerical results with/without branch liquid distributorThe numerical data were sampled on the inner wall in the plane x/ d m=À1,À0.5,0,0.5,1,2,3,4,6and8.At the same time,the numer-ical data were sampled from points every5mm along the intersec-tional lines of planes of y/d m=0and sections of x/d m=À2,À1,0,1,2, 3,4,5and6,to get the points with the maximum normalized rootK.Gao et al./Annals of Nuclear Energy78(2015)180–187185mean square temperatures in the tee and on the top and bottom walls.Here,the temperature and velocityﬁelds were determined with LES simulations for the case of tee junction with/without branch liquid distributor.The temperature contours and velocity vectors for the T-junction are shown in Figs.12and13,respectively.As can be seen in Fig.12,due to the large branch pipeﬂow velocity,hot and coldﬂuid mixing zone is mainly located in both upstream and downstream region of the intersections of the main pipe and the branch pipe.The vigorous mixing ofﬂuids in the tube leads to thermalﬂuctuation on the wall.But in the T-junction with the branch liquid distributor,the mixing region moves to the lower half and downstream region of the main pipe.This indicates that the distributor is advantageous to weaken thermalﬂuctuations on the wall.The same conclusion can be seen from Fig.13,the dis-tributor weaken thermalﬂuctuations on the wall of downstream region and the top of the main pipe.186K.Gao et al./Annals of Nuclear Energy78(2015)180–187Fig.14compares the normalized mean temperatures between two tees of different structures.As can be seen,wall temperature changes great in the direction of90°,135°,225°and270°in T-junc-tion with the distributor,because the directions are the distributor outlet directions.This indicates that the coldﬂuid mixes with hot ﬂuid on the wall afterﬂows out of the distributor.At the same time,the temperature in the direction of180°also changes dra-matically.That is because the coldﬂuid moves down in the effects of gravity and buoyancy.As shown in Fig.15,for the tee with distributor,the maximum values of normalized RMS temperature are smaller than that of the tee without distributor in most directions.This indicates that the adding of the distributor can relieve thermalﬂuctuations on the wall to some extent.And for the T-junction with distributor,tem-perature tends to be stable after the plane of x/d m=6,which indi-cates that twoﬂuids have made a full mixing,while for the initial tee,temperature is still in the dramatic change,and this shows that the improved structure can effectively reduce the mixing length.Fig.16shows the maximum normalized instantaneous temper-atureﬂuctuations in the tee and on the top and bottom walls in the plane y/d m=0.In the tee,the maximum normalized instantaneous temperatureﬂuctuations of the case without distributor vary from 0.45to0.8,which means that the hot and coldﬂuids alternate in this location.However,for the case with distributor,the tempera-tureﬂuctuations in the tee as well as on the top and bottom walls are much smaller than those of the case without distributor.That also implies that the distributor can reduce the temperatureﬂuctu-ation effectively.The normalized instantaneous temperatureﬂuctuations cannot describe the relationship between power spectrum density(PSD) and frequency of the temperatureﬂuctuation.PSD against fre-quency is one of the most important parameter for thermal fatigue analysis,which can directly show how PSD is in a certain fre-quency.The PSDs of the points with maximum temperatureﬂuctu-ation for the cases with and without distributor against frequency were recorded by fast Fourier transform(FFT)and shown in Fig.17. The temperatureﬂuctuation of the case without distributor has the highest PSD,at the frequency of0.04Hz,whereas the distributor signiﬁcantly reduces the PSD of the temperatureﬂuctuations with the frequency from0.01to0.1Hz.In addition,the PSD of temper-atureﬂuctuations decreases with the frequency increasing.5.ConclusionsAs thermal stratiﬁcation can result in thermal fatigue in the pip-ing system of a nuclear power plant,safety and integrity evaluation of the piping system has become an important issue.In this work the temperatureﬂuctuation has been studied by the experiment and numerically predicted by LES for two types of vertical tee junc-tion:one with distributor in the branch pipe and another without. The numerical results of normalized mean and RMS temperatures for the two structures have been found to be in good qualitative and quantitative agreement with the experimental data,which val-idates the use of LES simulations to evaluate convective mixing in such geometries.At the same time,the simulated normalized mean and root-mean square(RMS)temperature,velocity vector and temperature contour of the two tees are compared.The numerical results show that theﬂuctuations of temperatures of the tee without the distrib-utor are larger than those of the tee with the distributor,which can be explained by the branch liquid distributor enhancing the mix-ing.Although both tees give the same momentum ratio between the main pipeﬂow and the branch pipeﬂow,mixing and convec-tive heat transfer are greatly enhanced by the presence of the branch liquid distributor.These all show that the structure is effec-tive for weakening the thermalﬂuctuation of tee piping wall when hot and coldﬂuids mix,and it can make the mixing more sufﬁcient.AcknowledgementsThis work was supported by projects of the National Natural Science Foundation of China(No.51276009),Program for New Century Excellent Talents in University(No.NCET-13-0651),and the National Basic Research Program of China(No.2011CB706900). ReferencesAulery, F.,Toutant, A.,Monod,R.,Brillant,G.,Bataille, F.,2012.Numerical simulations of sodium mixing in a T-junction.Appl.Therm.Eng.37,38–43.Brücker,C.,1997.Study of the three-dimensionalﬂow in a T-junction using a dual-scanning method for three-dimensional scanning-particle-image velocimetry (3-D SPIV).Exp.Therm.Fluid Sci.14,35–44.Cao,Q.,Lu,D.,Lv,J.,2012.Numerical investigation on temperatureﬂuctuation of the parallel triple-jet.Nucl.Eng.Des.249,82–89.Durve,A.,Patwardhan,A.W.,Banarjee,I.,Padmakumar,G.,Vaidyanathan,G.,2010.Thermal striping in triple jetﬂow.Nucl.Eng.Des.240,3421–3434.Eric Blondet, C.F.,2002.High cycle thermal fatigue in french PWR.In:10th International Conference on Nuclear Engineering.Arlington,Virginia,USA,pp.429–436.Frank,T.,Lifante,C.,Prasser,H.M.,Menter,F.,2010.Simulation of turbulent and thermal mixing in T-junctions using URANS and scale-resolving turbulence models in ANSYS CFX.Nucl.Eng.Des.240,2313–2328.Galpin,J.,Simoneau,J.P.,rge Eddy Simulation of a thermal mixing tee in order to assess the thermal fatigue.Int.J.Heat Fluid Flow32,539–545. Hosseini,S.M.,Yuki,K.,Hashizume,H.,2008.Classiﬁcation of turbulent jets in a T-junction area with a90-deg bend upstream.Int.J.Heat Mass Transfer51,2444–2454.Hu,L.-W.,Kazimi,M.S.,2006.LES benchmark study of high cycle temperature ﬂuctuations caused by thermal striping in a mixing tee.Int.J.Heat Fluid Flow 27,54–64.IAEA,2002.Validation of Fast Reactor Thermomechanical and Thermohydraulic Codes,Vienna.Jayaraju,S.T.,Komen,E.M.J.,Baglietto,E.,2010.Suitability of wall-functions in Large Eddy Simulation for thermal fatigue in a T-junction.Nucl.Eng.Des.240,2544–2554.Kamaya,M.,Nakamura, A.,2011.Thermal stress analysis for fatigue damage evaluation at a mixing tee.Nucl.Eng.Des.241,2674–2687.Lee,J.I.,Hu,L.-W.,Saha,P.,Kazimi,M.S.,2009.Numerical analysis of thermal striping induced high cycle thermal fatigue in a mixing tee.Nucl.Eng.Des.239, 833–839.Lu,T.,Jiang,P.X.,Guo,Z.J.,Zhang,Y.W.,Li,H.,rge-eddy simulations(LES)of temperatureﬂuctuations in a mixing tee with/without a porous medium.Int.J.Heat Mass Transfer53,4458–4466.Lu,T.,Liu,S.M.,Attinger,D.,rge-eddy simulations of structure effects of an upstream elbow main pipe on hot and coldﬂuids mixing in a vertical tee junction.Ann.Nucl.Energy60,420–431.Metzner,K.J.,Wilke,U.,2005.European THERFAT project—thermal fatigue evaluation of piping system‘‘Tee’’-connections.Nucl.Eng.Des.235,473–484. Paffumi, E.,Radu,V.,Nilsson,K.F.,2013.Thermal fatigue striping damage assessment from simple screening criterion to spectrum loading approach.Int.J.Fatigue53,92–104.Ricard,J.B.,Sperandio,M.,1996.Fracture mechanics applied to superphenix reactor components.Int.J.Pressure Vessels Piping65,295–301.Simoneau,J.-P.,Champigny,J.,Gelineau,O.,2010.Applications of large eddy simulations in nuclearﬁeld.Nucl.Eng.Des.240,429–439.Wu,H.L.,Peng,X.F.,Chen,T.K.,2003.Inﬂuence of sleeve tube on theﬂow and heat transfer behavior at a T-junction.Int.J.Heat Mass Transfer46,2637–2644.K.Gao et al./Annals of Nuclear Energy78(2015)180–187187。

我做了物理实验英语作文

我做了物理实验英语作文英文回答：In the domain of experimental physics, the meticulous execution of experiments can unveil profound insights into the intricate tapestry of natural phenomena. I embarked on a comprehensive experimental investigation, meticulously adhering to the scientific method. The experiment entailed the systematic manipulation of independent variables to elucidate their effects on dependent variables.Through meticulous data collection and rigorous analysis, I established clear and quantifiable relationships between the variables, affirming a causal connection. The outcomes of the experiment provided empirical support for preexisting theoretical frameworks and, in some instances, necessitated a reassessment of established paradigms.The meticulously crafted experimental design isolatedthe variables of interest, effectively eliminating confounding factors that could have potentially compromised the integrity of the findings. The precision with which measurements were taken ensured the reliability andvalidity of the data, allowing for meaningfulinterpretation and robust conclusions.The experiment not only served as a crucible fortesting hypotheses but also as an invaluable educational tool. It engendered a profound understanding of the scientific process, from the formulation of research questions to the dissemination of findings. By actively engaging in the scientific method, I developed a heightened appreciation for the rigor and discipline required in the pursuit of knowledge.In conclusion, the meticulously executed experimental investigation provided valuable insights into the relationships between the variables under scrutiny. The experiment reinforced the power of the scientific method and ignited a passion for further exploration in the enigmatic realm of physics.中文回答：实验物理论文：我开展了一项全面的实验研究，严格遵循科学方法。

UHPC-NC叠层梁界面黏结性能的试验研究与数值模拟

1叠层梁静载试验
1.1叠层梁试件为研究 UHPC-NC 叠梁中面结强
对受力行为的影响，本研究浇筑了 3个UHPC-NC 叠层梁试件UN50,UN60,UN70,其矩形横截面宽 250 mm,高150 mm；试件全长1 200 mm,加载净跨径为1 000 mm。试件UN50,UN60,UN70的下层 UHPC厚度分别为50,60,70 mm,如表1所示。
DOI： 10. 19815/j. jace. 2020. 11007
叠层 UHPC-NC
面黏结性能的试验研
究与数值模
许建明，陈勇，刘骁繁，许奇
(国网江苏省电力有限公司建设分公司，江苏南京210036)
摘要：对3根不同厚度比的超高性能混凝土-普通混凝土(UHPC-NC)叠层梁的静力行为与界面黏
结性能进行了试验研究，观测到试验梁的最终破坏发生在界面处，说明UHPC-NC界面处的性能对叠层构件整体受力行为有显著影响，为此改进了现有的界面斜剪试验，使之更加准确地测量
为了更好地模拟UHPC-NC的受力行为，本文
采用非局部的方法来模拟UHPC与NC层，通过合
并高阶变形量形成梯度损伤模型"0#。这种梯度损
伤模型可以模拟损伤的产生与发展，且具有更稳定
的显性算法。对于2种材料之间的界面，本文将不
再采用完美黏结假设，充分考虑界面的黏结破坏以
及其对UHPC-NC组合构件受力行为的影响。
UN50
UN60B 1UN70
(a)浇筑下层UHPC
(b) 28 d后浇筑上层NC
图1 UHPC-NC叠层试件示意图(单位：mm)
Fig. 1 Illustrations of UHPC-NC Composite Specimens (Unit：mm)

An experimental and numerical investigation of

Combustion and Flame 145(2006)740–764/locate/combustﬂameAn experimental and numerical investigation of n -heptane/air counterﬂow partially premixed ﬂamesand emission of NO x and PAH speciesPaolo Berta a ,Suresh K.Aggarwal a ,∗,Ishwar K.Puri ba Department of Mechanical and Industrial Engineering,University of Illinois at Chicago,Chicago,IL,USAb Department of Engineering Science and Mechanics,Virginia Polytechnic Institute and State University,Blacksburg,VA,USAReceived 13July 2005;received in revised form 27January 2006;accepted 30January 2006Available online 23March 2006AbstractAn experimental and numerical investigation of counterﬂow prevaporized partially premixed n -heptane ﬂames is reported.The major objective is to provide well-resolved experimental data regarding the detailed structure and emission characteristics of these ﬂames,including proﬁles of C 1–C 6,and aromatic species (benzene and toluene)that play an important role in soot formation.n -Heptane is considered a surrogate for liquid hydrocarbon fuels used in many propulsion and power generation systems.A counterﬂow geometry is employed,since it provides a nearly one-dimensional ﬂat ﬂame that facilitates both detailed measurements and simulations using comprehen-sive chemistry and transport models.The measurements are compared with predictions using a detailed n -heptane oxidation mechanism that includes the chemistry of NO x and PAH formation.The reaction mechanism was syn-ergistically improved using pathway analysis and measured benzene proﬁles and then used to characterize the effects of partial premixing and strain rate on the ﬂame structure and the production of NO x and soot precursors.Measurements and predictions exhibit excellent agreement for temperature and major species proﬁles (N 2,O 2,n -C 7H 16,CO 2,CO,H 2),and reasonably good agreement for intermediate (CH 4,C 2H 4,C 2H 2,C 3H x )and higher hydrocarbon species (C 4H 8,C 4H 6,C 4H 4,C 4H 2,C 5H 10,C 6H 12)and aromatic species (toluene and benzene).Both the measurements and predictions also indicate the existence of two partially premixed regimes;a double ﬂame regime for φ<5.0,characterized by spatially separated rich premixed and nonpremixed reaction zones,and a merged ﬂame regime for φ>5.0.The NO x and soot precursor emissions exhibit strong dependence on partial premixing and strain rate in the ﬁrst regime and relatively weak dependence in the second regime.At higher levels of partial premixing,NO x emission is increased due to increased residence time and higher peak temperature.In contrast,the emissions of acetylene and PAH species are reduced by partial premixing because their peak locations move away from the stagnation plane,resulting in lower residence time,and the increased amount of oxygen in the system drives the reactions to the oxidation pathways.The effects of partial premixing and strain rate on the production of PAH species become progressively stronger as the number of aromatic rings increases.©2006The Combustion Institute.Published by Elsevier Inc.All rights reserved.Keywords:n -Heptane;Partially premixed ﬂames;NO x and PAH species measurements;Detailed modeling*Corresponding author.E-mail address:ska@ (S.K.Aggarwal).P.Berta et al./Combustion and Flame145(2006)740–7647411.IntroductionA major portion of the world’s energy demands is currently met by the combustion of liquid fuels. Closely associated with the beneﬁts derived from combustion are the hazards it causes to human life and environment.The products of combustion of most commercially available fuels contain pollutants such as particulate matter,unburned and partially unburned hydrocarbons,carbon monoxide,and oxides of nitro-gen and sulfur.These pollutants have many harmful effects including speciﬁc health hazards,acid rain, smog,global warming,and ozone depletion.The ac-ceptability of a new grade of fuel or design of a new combustion system at present depends as much on its emission characteristics as on its combustion ef-ﬁciency.Consequently,energy conservation and en-vironmental concerns provide a strong motivation for fundamental studies on the mechanism of soot and NO x formation inﬂames.Partially premixedﬂames contain a rich premixed fuel–air mixture in a pocket or stream,and,for com-plete combustion to occur,they require the transport of oxidizer from an appropriately oxidizer-rich(or fuel-lean)mixture that is present in another pocket or stream.Partial oxidation reactions occur in fuel-rich portions of the mixture and any remaining unburned fuel and/or intermediate species are consumed in the oxidizer-rich portions.Partially premixedﬂames are important in numerous applications.They are rele-vant to turbulent nonpremixed combustion,which can contain regions where local extinction occurs,fol-lowed by partial premixing and reignition.Partially premixed combustion plays a fundamental role in the stabilization of lifted nonpremixedﬂames in which propagating premixed reaction zones anchor a non-premixed reaction zone.In addition,in most liquid-fueled combustion devices,such as internal combus-tion engines,industrial furnaces,and power station gas turbines,the fuel is introduced in the form of a spray of fuel droplets of different sizes.The smaller droplets evaporate at a much higher rate than the larger ones.The resulting fuel vapor mixes with air, forming locally fuel-rich zones.The larger droplets then burn in this mixture in a partially premixed mode.Partially premixedﬂames may also result in lean direct injection diesel engines.The liquid fuels that are used in internal com-bustion engines and gas turbines are typically blends of several components.Generally,fuels with desired properties are prepared by mixing expensive volatile components with cheaper heavier fuels.The detailed simulation and analysis ofﬂames burning these fu-els in actual engines is a prohibitively complex task single-component or bicomponent fuel,based on the most abundant species in the actual fuel.In prac-tical liquid fuels such as gasoline and diesel fuels, n-C7H16is relatively abundant,and hence often used as a surrogate for these fuels.Since the soot-and NO x-forming mechanisms are closely related to the chemical kinetics and structure ofﬂames,a detailed study of partially premixed n-heptaneﬂames(PPFs)is of direct relevance to op-timizing the operating conditions of a diesel engine for minimum production of soot,unburned hydro-carbons,and NO x.Due to these diverse applications and fundamental relevance,partially premixedﬂames have been investigated extensively in recent years. However,the bulk of these studies have focused on methane–airﬂames[1–5],motivated perhaps by the fact that detailed reaction mechanisms are available to model the methane–air chemistry.With the excep-tion of some recent investigations[6–9],the literature regarding the burning of higher hydrocarbon fuels,es-pecially liquid fuels,in partially premixedﬂames is relatively sparse.Li and Williams[6]reported mea-surements of several major and intermediate species in n-heptane PPFs burning a droplet/air fuel mixture in a counterﬂow conﬁguration.Seiser et al.[10]re-ported an experimental investigation of prevaporized n-heptane counterﬂow nonpremixedﬂames.Xue and Aggarwal[7]characterized the structure of n-heptane counterﬂow PPFs through a numerical investigation and subsequently investigated the effect of double ﬂame structure on NO x formation in theseﬂames[8]. Berta et al.[38]recently reported an experimental and numerical investigation of the structure and emis-sion characteristics of prevaporized n-heptane non-premixedﬂames in a counterﬂow conﬁguration.Our literature indicates that there is a lack of de-tailed experimental data pertaining to the structure and emission characteristics of n-heptane PPFs.This is rather surprising since n-heptane has been con-sidered a good surrogate for liquid fuels used in many practical combustion systems,and its oxidation chemistry has been extensively investigated.More-over,compared to other combustion systems,includ-ing premixed and nonpremixedﬂames,a PPF pro-vides a more stringent crucible for the validation of reaction mechanisms[11].This is due to the exis-tence of multiple reaction zones and interactions be-tween them involving both chemistry and transport processes.These interactions also play a signiﬁcant role in determining the NO x and soot emissions from theseﬂames.Motivated by the above considerations,we report herein an experimental–computational investigation of partially premixed n-heptaneﬂames established742P.Berta et al./Combustion and Flame145(2006)740–764species concentrations,especially those of C1–C6hy-drocarbons,for a wide range of partial premixing (i.e.,equivalence ratios)and strain rates.C1–C6hy-drocarbons are key intermediates in the fuel decom-position pathway and their characterization is crucial for understanding the combustion of heavier fuels,es-pecially in the context of partially premixedﬂames, which are hybridﬂames and whose structure is char-acterized by both transport and chemical kinetics. Species concentration proﬁles of intermediate hydro-carbons can be subsequently used for the validation of computational models and reaction mechanisms involving simulations of liquid fuels in general,and n-heptane in particular.Therefore,we report well-resolved measurements of major species(n-C7H16, O2,N2,CO2,and H2O),intermediate species(CO, H2,CH4,C2H4,C2H2,and C3H x),higher hydrocar-bon species(C4H8,C5H10,and C6H12),and the ma-jor soot precursor(benzene)over a large parametric space characterized in terms of equivalence ratio(φ) and strain rate(a G).The measurements also focus on the resolution of unsaturated C3and C4species such as propene,propyne,allene,butene,1,3-butadiene, 1-buten-3-yne,1,3-butadiyne,and aromatic species (benzene and toluene).Some of these species have never been previously measured for n-heptane coun-terﬂow PPFs.Another objective is to characterize the effect of partial premixing on the formation of NO x and soot precursors,such as acetylene,benzene,and other PAH(polycyclic aromatic hydrocarbon)species,in n-heptane PPFs.Acetylene represents a key species in the formation of polyaromatic structures through the hydrogen abstraction carbon addition(HACA)mech-anism[12],while benzene represents the simplestaromatic molecule.The numerical investigation hasbeen performed using a detailed mechanism that is ca-pable of simulating the formation of NO x and PAHsup to coronene.2.The experimental setupA schematic of the experimental setup used to es-tablish prevaporized n-heptane counterﬂowﬂames ispresented in Fig.1.A mixture of prevaporized n-heptane and nitrogen fuel was introduced from thebottom nozzle.A nitrogen curtain was establishedthrough an annular duct surrounding the fuel jet inorder to isolate theﬂames from ambient disturbances.This nitrogen and combustion products were ventedand cooled through another annular duct around theoxidizer nozzle.The diameter of each nozzle was27.38mm,and the separation distance(L)betweenthem was varied from10to20mm.The veloci-ties of the two streams deﬁne the global strain rateas a G=(2|V O|/L)(1+(|V F|/|V O|)(ρF/ρO)1/2)[13] and were chosen to satisfy the momentum balance,ρO V2O=ρF V2F.Hereρrepresents density,V gas ve-locity,and the subscripts O and F refer to oxidizer andfuel nozzles,respectively.The oxidizer was air at room temperature,whilethe fuel stream consisted of mixtures of air and pre-vaporized n-heptane.The fuel nozzle was heatedand its temperature controlled to maintain the fuel-containing stream at a400K temperature at theburnerP.Berta et al./Combustion and Flame145(2006)740–764743exit.In the bottom part of the burner preheated air was mixed with the pure fuel stream to form a fuel–air mixture of the desired equivalence ratio.The n-heptane vapor was formed in a prevaporizer,which was an electrically heated stainless steel chamber.The desired massﬂow rate of n-heptane into the prevapor-izer was maintained by a liquid pump.Approximately three-fourths of the chamber wasﬁlled with glass beads to impede theﬂow,thereby increasing its res-idence time and thus enhancing the heat transfer to the liquid fuel.The temperature of the fuel vapor ex-iting the chamber was monitored by a thermocouple.Temperature proﬁles of variousﬂames were mea-sured using a Pt–Pt13%Rh thermocouple with a spherical bead diameter of0.25mm and wire diam-eter of0.127mm.The measured values were cor-rected for radiation heat losses from the bead,assum-ing a constant emissivity of0.2and a Nusselt num-ber of2.0[10].Species concentration proﬁles were measured using a Varian CP-3800gas chromatograph (GC).Samples were drawn from theﬂame with a quartz microprobe that had a0.34-mm tip diameter and0.25-mm tip oriﬁce.Constant vacuum was ap-plied at the end of the line through a vacuum pump. The line carrying the sample to the GC was made of fused silica and was heated to prevent conden-sation.A portion of the sample was injected into a Hayesep DB100/120packed column connected to a thermal conductivity detector to measure light gases (up to C2H4)and another into a Petrocol DH capillary column that was placed inline with aﬂame ioniza-tion detector to obtain hydrocarbon distributions up to C7H16.The temperature in the gas chromatograph oven was gradually increased to minimize the analy-sis time.The temperature and pressure in the sam-pling loops were controlled to ensure that the same volume of gas was sampled for each analysis.The chromatogram peaks have been converted into mole fractions with calibration constants that were obtained separately for every species from known standards. Water molar fractions were obtained through a mass balance of carbon and hydrogen atoms.The errors in measurement of the liquid fuel and airﬂow rates are within5%,leading to an uncertainty of about5% in equivalence ratio.The compositions of both the fuel and air streams were also measured using GC. C3and C4unsaturated species were measured ofﬂine by an HP6890gas chromatograph connected to a mass spectroscopy detector.The sample was collected in a stainless steel vessel.The whole line and vessel were heated to minimize condensation.Temperature programming was employed to reduce the analysis time.The temperature and pressure in the sampling loop were controlled and measured to ensure that the fractions with calibration constants that were obtained separately for every species from known standards. The uncertainties in GC measurements are between 5%and10%depending on the species.3.The physical–numerical modelMost of the studies on heptaneﬂames reported in the literature deal with nonpremixedﬂames.Experi-mental results have been obtained in several conﬁg-urations:liquid pool burners[14,15],droplet burn-ing[16,17],and premixedﬂames[18].n-Heptane combustion chemistry has been investigated on many different levels.One-step global and reduced mech-anisms[19,20]have been empirically derived toﬁt experimental data of burning velocities andﬂame extinction.Held et al.[17]reported a semidetailed mechanism and validated it usingﬂow reactor,shock tube,stirred reactor,and laminarﬂame speed experi-mental data.The mechanism was subsequently used for predicting ignition delays in shock tubes[21] and for numerical investigations of partially premixed ﬂames[7,8].Lindstedt and Maurice[22]developed a detailed n-heptane mechanism,addressing in detail the H abstraction reactions on the C7molecule and its decomposition into smaller fragments.The mecha-nism was improved in subsequent work[23]to further characterize the formation and oxidation of aromatic molecules.Detailed n-heptane mechanisms have also been reported by Chakir et al.[24],Curran et al.[25], and Babushok and Tsang[26].The kinetic mechanism(SOX)used to model n-heptaneﬂames in the present study was previously developed by extending a detailed oxidation scheme for several fuels[27,28].Due to the hierarchical mod-ularity of the mechanistic scheme,this model is based on a detailed submechanism of C1–C4species.As-suming analogy rules for similar reactions,only a few fundamental kinetic parameters are required for the progressive extension of the scheme toward heavier species.The resulting kinetic model of hydrocarbon oxidation from methane up to n-octane consists of about170species and5000reactions.We have selected this mechanism for our simu-lations since the subset of n-heptane oxidation reac-tions included in it has been extensively tuned using experimental measurements for pure pyrolysis condi-tions,oxidation in jet-stirred and plug-ﬂow reactors, and shock-tube experiments[29].Moreover,a rela-tively detailed model for polycyclic aromatic hydro-carbons(PAHs)that are soot precursors is contained in the mechanism.The formation of theﬁrst aromatic rings by C2and C4chemistry and by resonantly sta-744P.Berta et al./Combustion and Flame145(2006)740–764 gated[28,30].Further growth of PAH species up tocoronene(C24H12)is also modeled through the well-known HACA mechanism[31],which has been ex-tensively validated for counterﬂowﬂames burning avariety of fuels[32].The main consumption reactionsof aromatics and PAHs are H abstraction reactions byH and OH radicals.The high-temperature reactionshave been validated against substantial experimentaldata[27,28,30].Numerical simulations of counterﬂowﬂames wereperformed using the OPPDIF code[33],which is ca-pable of modeling combustion between two opposedjets.The code was modiﬁed to handle the complexreaction mechanism and to account for thermal radia-tion through an optically thin model[34].Most ther-modynamic properties were obtained from Burcat andMcBride[35],and unavailable properties were esti-mated using the group additivity and difference meth-ods[36].Transport properties were obtained from theCHEMKIN database[37]wherever available,whileunavailable data were deduced through analogy withknown species.To establish grid independence,numerical solu-tions were obtained on increasinglyﬁner grids,andby changing GRAD and CURV parameters,until novariation was observed between two grid systems.4.Results and discussionTo perform a detailed experimental and numericalinvestigation of theﬂame structure and emission char-acteristics,prevaporized n-heptane PPFs were estab-lished at different strain rates(a G)and equivalence ra-tios(φ).Table1shows the parametric space in termsof a G andφfor seven PPFs,designated as FlamesA–G,which are analyzed experimentally and numer-ically in the present study.For all the cases,the fuelstream was introduced from the bottom nozzle and theoxidizer from the top nozzle.The oxidizer was pureair,while the fuel stream was a mixture of n-heptaneand air with the desired value ofφ.Note that PPFsTable1Operating conditions in terms of strain rate,equivalence ra-tio,and nozzle separation distance for the cases investigatednumerically and experimentallyFlame Strain rate(s−1)EquivalenceratioNozzle separation(cm)A5015.31 B506.11 C502.52 D1008.01 E15012.61established at a G=100s−1and different values ofφhave been investigated in our previous work[9].Con-sequently,only one value ofφis considered at this strain rate(Flame D).For preliminary analysis,digital images of several PPFs were taken for different values of strain rate, partial premixing,and nozzle separation distance.The images of four representativesﬂames,i.e.,Flames A, C,G,and E,are presented in Fig.2.The images of Flames A,G,and E were taken at the same exposure time,while that of Flame C was taken at double the exposure time for it was less luminous.For Flame A, which is characterized by low strain rate and low level of partial premixing,an orange-red zone can be ob-served below the familiar green-blue doubleﬂame structure,with green from the C2chemiluminescence in the premixed zone and blue from the CO oxidation in the nonpremixed zone.Even though the equiva-lence ratio is high theﬂame does not appear as sooty as a nonpremixedﬂame,which is bright yellow.The red zone disappears as the strain rate and/or level of partial premixing are increased.Flame C shows the greatest separation between the two reaction zones; as the partial premixing approaches stoichiometric conditions the premixedﬂame moves closer to the fuel nozzle.Thisﬂame does not appear asﬂat as the othersﬂames because the nozzle separation dis-tance had to be increased to obtain the desired strain rate.The doubleﬂame structure can still be seen in Flame G,which appears brighter,since more fuel is consumed during the same exposure time due to the higher strain rate.In Flame E,which is characterized by a higher equivalence ratio,the doubleﬂame struc-ture can barely be noticed,as the two reaction zones are nearly merged.A detailed comparison of measurements and sim-ulations for the sevenﬂames listed in Table1is pre-sented in Figs.3–9.Eachﬁgure shows the temper-ature,axial velocity,and species mole fraction pro-ﬁles.The predictions are shown by continuous lines, while the experimental data are shown by symbols. The three vertical lines in eachﬁgure indicate that(1) the nonpremixed reaction zone location that is iden-tiﬁed by the peak in temperature proﬁle and marked by the solid vertical line;(2)the stagnation plane that is marked by the dashed vertical line;and(3)the rich premixed zone location that is identiﬁed by the peak in hydrogen proﬁle and marked by the vertical dotted line.The experimental proﬁles are highly re-solved,since particular effort has been expended to capture the regions characterized by high chemical activity and steep gradients.Moreover,measurements of several intermediate hydrocarbon species that areP.Berta et al./Combustion and Flame145(2006)740–764745Fig.2.Digital images of partially premixed n-heptaneﬂames(Flames A,C,G,E).(1)A general observation from the measured andpredicted proﬁles for the sevenﬂames is that PPFs are characterized by a doubleﬂame struc-ture:a rich premixed zone is established down-stream of the fuel nozzle and characterized by pyrolysis and partial oxidation of n-heptane.The products of partial oxidation,namely CO,H2, and intermediate hydrocarbon species,are trans-ported and consumed in the nonpremixed reac-tion zone located on the oxidizer side.The double ﬂame structure becomes visually more distinct as a G decreases and/or the level of partial pre-mixing increases(i.e.,φdecreases).This also increases the separation distance between the two reaction zones.The premixed reaction zonevelocity(V x)matches the burning velocity(S L)of the stretchedﬂame.Since S L increases asφisreduced,the premixedﬂame moves away fromthe stagnation plane toward the fuel nozzle tosatisfy the condition S L=V x.The nonpremixed ﬂame is established on the oxidizer side at the lo-cation(x n)where the intermediate fuel speciesand oxidizerﬂuxes are transported in stoichio-metric proportion.Therefore,the separation dis-tance between the two reaction zones increasesas the level of partial premixing is increased.In-creasing the strain rate has the opposite effect,since for largerﬂow velocities the location x p ispushed toward the stagnation plane.(2)For all the sevenﬂames analyzed,there is gen-746P.Berta et al./Combustion and Flame145(2006)740–764Fig.3.Predicted(lines)and measured(symbols)proﬁles for Flame A.Temperature and axial velocity proﬁles;mole fraction proﬁles of O2,N2,and n-C7H16;mole fraction proﬁles of H2O,CO2,CO,and H2;mole fraction proﬁles of CH4,ethylene, acetylene,and C3hydrocarbons;mole fraction proﬁles of C4H8,C5H10,and C6H12oleﬁns;and mole fraction proﬁles ofP.Berta et al./Combustion and Flame145(2006)740–764747Fig.4.Predicted(lines)and measured(symbols)proﬁles for Flame B.Temperature and axial velocity proﬁles;mole fraction proﬁles of O2,N2,and n-C7H16;mole fraction proﬁles of H2O,CO2,CO,and H2;mole fraction proﬁles of CH4,ethylene, acetylene,and C3hydrocarbons;mole fraction proﬁles of C4H8,C5H10,and C6H12oleﬁns;and mole fraction proﬁles of benzene.The vertical lines in some of theﬁgures indicate the locations of the stagnation plane,the premixedﬂame,and the748P.Berta et al./Combustion and Flame145(2006)740–764Fig.5.Predicted(lines)and measured(symbols)proﬁles for Flame C.Temperature and axial velocity proﬁles;mole fraction proﬁles of O2,N2,and n-C7H16;mole fraction proﬁles of H2O,CO2,CO,and H2;mole fraction proﬁles of CH4,ethylene, acetylene,and C3hydrocarbons;mole fraction proﬁles of C4H8,C5H10,and C6H12oleﬁns;and mole fraction proﬁles of benzene.The vertical lines in some of theﬁgures indicate the locations of the stagnation plane,the premixedﬂame,and theP.Berta et al./Combustion and Flame145(2006)740–764749Fig.6.Predicted(lines)and measured(symbols)proﬁles for Flame D.Temperature and axial velocity proﬁles;mole fraction proﬁles of O2,N2,and n-C7H16;mole fraction proﬁles of H2O,CO2,CO,and H2;mole fraction proﬁles of CH4,ethylene, acetylene,and C3hydrocarbons;mole fraction proﬁles of C4H8,C5H10,and C6H12oleﬁns;and mole fraction proﬁles of benzene.The vertical lines in some of theﬁgures indicate the locations of the stagnation plane,the premixedﬂame,and theFig.7.Predicted(lines)and measured(symbols)proﬁles for Flame E.Temperature and axial velocity proﬁles;mole fraction proﬁles of O2,N2,and n-C7H16;mole fraction proﬁles of H2O,CO2,CO,and H2;mole fraction proﬁles of CH4,ethylene,acetylene,and C hydrocarbons;mole fraction proﬁles of C H,C H,and C H oleﬁns;and mole fraction proﬁles ofproﬁles of O2,N2,and n-C7H16;mole fraction proﬁles of H2O,CO2,CO,and H2;mole fraction proﬁles of CH4,ethylene,proﬁles of O2,N2,and n-C7H16;mole fraction proﬁles of H2O,CO2,CO,and H2;mole fraction proﬁles of CH4,ethylene,dictions.The predictions reproduce the measured partially premixedﬂame structure for all the strain rates and equivalence ratios investigated.Both measurements and predictions indicate that at low level of partial premixing(highφ)and/or high strain rate,the two reaction zones are nearly merged,and that as the level of partial premixing increases and/or a G decreases,the separation dis-tance between the two reaction zones increases and the doubleﬂame structure becomes more dis-cernible.(3)There is good quantitative agreement betweenmeasurements and predictions for major reac-tant and product species(n-heptane,O2,N2,and CO2)as well as for intermediate fuel species(H2 and CO).A good agreement between the mea-sured and predicted peak concentrations of these species and between the locations of their peak concentrations implies that both the transport and chemistry are reasonably well reproduced in the simulations.For instance,a good agreement be-tween the measured and predicted locations of the H2and CO concentration peaks implies that the location of the rich premixed zone is well pro-duced by the simulations.Similarly,good agree-ment between the measured and predicted loca-tions of the CO2concentration peaks indicates that the location of the nonpremixed zone is well reproduced by the simulations.The measured and predicted temperature proﬁles also exhibit good agreement,although there is a mismatch between the locations of the respective peaks.A similar discrepancy has been observed by otherinvestigators,and may partly be attributed to the catalytic effect of the thermocouple.(4)The n-heptane and O2proﬁles on the fuel side in-dicate that the reaction mechanism underpredicts the consumption rates of these species in the rich premixed zone;it can be seen forﬂame E and toa lesser extent also for Flames A,B,F,and G.This is further corroborated by the H2,CO,and intermediate hydrocarbon species(C2H4,C2H2, C4H8,and C5H10)proﬁles.This discrepancy be-comes less pronounced,however,as the level of partial premixing is increased,i.e.,asφis de-creased.(5)There is also fairly good quantitative agreementbetween measurements and predictions for light hydrocarbon species(CH4,C2H4,C2H2).How-ever,the quantitative agreement deteriorates, although there is good qualitative agreement, for higher hydrocarbon species(C4H8,C5H10), which are present at relatively low concentra-the measured proﬁles of this species are shown for all the sevenﬂames.(6)The comparison of the measured and predictedH2O proﬁles exhibits large discrepancies.This is due to the fact that H2O concentration is not measured directly from GC;it is obtained by ap-plying a balance of carbon and hydrogen atoms using the GC measurements of the other species.This procedure implies equal diffusivity for all the species,an assumption that is not well satis-ﬁed,especially on the fuel side due to the pres-ence of large(C7H16)and small(H2)molecules that have very different diffusion coefﬁcients. (7)In the present study,particular attention wasgiven to the measurement and prediction of ben-zene proﬁles,since this species has the simplest structure,with a single aromatic ring,and rep-resents perhaps the most important intermediate in the growth process to PAHs and soot.In spite of their relatively low concentrations,the pre-dicted and measured benzene proﬁles exhibit fairly good agreement for all the sevenﬂames shown in Figs.3–9.Both measurements and pre-dictions indicate that the benzene concentration decreases as the level of partial premixing and/or strain rate is increased.It is worth mentioning that the predictions are based on the reaction mechanism that was synergistically improved using pathway analysis and measured benzene proﬁles in our previous investigation[38].In order to better characterize the pyrolysis zone, further analysis was performed for Flames A,E,and G and the results are presented in Figs.10–12.The objective was to obtain quantitative data on the dis-tribution of unsaturated C3–C4intermediates for dif-ferent levels of partial premixing(equivalence ratio) and strain rates.The C3–C4intermediates proﬁles presented here are extremely valuable for the develop-ment and testing of n-heptane reaction mechanisms, since these species constitute the main decomposi-tion products of the C7H15radical.In addition,they directly affect the formation of the propargyl radical [26],which,through benzene and PAHs,leads to soot formation.The measurements were taken using the ofﬂine technique,as described earlier,on the fuel side of theﬂame,where n-heptane undergoes rapid con-sumption.For most of the species reported in Figs. 10–12,there is good agreement between predictions and measurements,especially considering that the de-tailed chemistry model has not been tuned for this set of data,implying that the reaction mechanism ade-。

磷酸铁锂电池火灾危险性分类

磷酸铁锂电池火灾危险性分类董海斌, 张少禹, 李毅, 等. ncm811高比能锂离子电池热失控火灾特性[j]. 储能科学与技术, 2019, 8(s1): 65-70.[本文引用: 1]dong h b, zhang s y, li y, et al. thermal runaway fire characteristics of lithium ion batteries with high specific energy ncm811[j]. energy storage science and technology, 2019, 8(s1): 65-70.[本文引用: 1][2]forgez c, vinh d d, friedrich g, et al. thermal modeling of a cylindrical lifepo4/graphite lithium-ion battery[j]. journal of power sources, 2010, 195(9): 2961-2968.[3]huang p f, ping p, li k, et al. experimental and modeling analysis of thermal runaway propagation over the large format energy storage battery module withli4ti5o12 anode[j]. applied energy, 2016, 183: 659-673. [4]larsson f, mellander b e. abuse by external heating, overcharge and short circuiting of mercial lithium-ion battery cells[j]. journal of the electrochemical society, 2014, 161(10): a1611-a1617.[5]li d j, danilov d l, gao l, et al. degradation mechanisms of c6/lifepo4 batteries: experimental analyses of cycling-induced aging[j]. electrochimica acta, 2016, 210: 445-455.[本文引用: 1][6]feng x n, sun j, ouyang m g, et al. characterization of penetration induced thermal runaway propagation process within a large format lithium ion battery module[j]. journal of power sources, 2015, 275: 261-273.[本文引用: 1][7]lamb j, orendorff c j, steele l a m, et al. failure propagation in multi-cell lithium ion batteries[j]. journal of power sources, 2015, 283: 517-523.[8]lopez c f, jeevarajan j a, mukherjee p p. experimental analysis of thermal runaway and propagation inlithium-ion battery modules[j]. journal of the electrochemical society, 2015, 162(9): a1905-a1915.[9]li h, duan q l, zhao c p, et al. experimental investigation on the thermal runaway and its propagation in the large format battery module withli(ni1/3co1/3mn1/3)o2 as cathode[j]. journal of hazardous materials, 2019, 375: 241-254.[本文引用: 2][10]刘昱君, 段强领, 黎可, 等. 多种灭火剂扑救大容量锂离子电池火灾的实验研究[j]. 储能科学与技术, 2018, 7(6): 1105-1112.liu y j, duan q l, li k, et al. experimental study on fire extinguishing of large-capacity lithium-ion batteries by various fire extinguishing agents[j]. energy storage science and technology, 2018, 7(6): 1105-1112.[本文引用: 1][11]babrauskas v. pillow burning rates[j]. fire safety journal, 1985, 8(3): 199-200.[本文引用: 1][12]ribière p, grugeon s, morcrette m, et al. investigation on the fire-induced hazards of li-ion battery cells by fire calorimetry[j]. energy & environmental science, 2012, 5(1): 5271-5280.[本文引用: 2][13]ping p, wang q s, huang p f, et al. study of the fire behavior of high-energy lithium-ion batteries with full-scale burning test[j]. journal of power sources, 2015, 285: 80-89.[本文引用: 1][14]ping p, kong d p, zhang j q, et al. characterization of behaviour and hazards of fire and deflagration for high-energy li-ion cells by over-heating[j]. journal of power sources, 2018, 398: 55-66.[本文引用: 1][15]li h, chen h d, zhong g b, et al. experimental study on thermal runaway risk of 18650 lithium ion battery under side-heating condition[j]. journal of loss prevention in the process industries, 2019, 61: 122-129.[本文引用: 1][16]feng x n, sun j, ouyang m g, et al. characterization of large format lithium ion battery exposed to extremely high temperature[j]. journal of power sources, 2014, 272: 457-467.[本文引用: 1][17]larsson f, bertilsson s, furlani m, et al. gas explosions and thermal runaways during external heating abuse of mercial lithium-ion graphite-licoo2 cells at different levels of ageing[j]. journal of power sources, 2018, 373: 220-231.[本文引用: 1]1... 然而，由于自身的物理化学性质，当不正确使用时(热滥用、电滥用和机械滥用)，磷酸铁锂电池会发生不可逆的热失控行为，存在较大的火灾危险性[1-5].在储能电站、变电站等实际运营场景中，往往将成百上千节的电池单体经过串并联后形成电池模组或者电池簇后集中使用.在该种情况下，一旦其中某节电池发生火灾，其释放的强热、燃烧等行为会造成周围电池温度上升，导致整个电池模组的热失控，甚至造成整个电池系统的火灾、爆炸事故[6-9].因此，在电化学储能以及变电系统等大规模应用场景中，研究锂离子电池热失控的火灾危险性并针对性开发相应的火灾抑制技术，对于电池系统的安全运行尤为重要. ...1... 然而，由于自身的物理化学性质，当不正确使用时(热滥用、电滥用和机械滥用)，磷酸铁锂电池会发生不可逆的热失控行为，存在较大的火灾危险性[1-5].在储能电站、变电站等实际运营场景中，往往将成百上千节的电池单体经过串并联后形成电池模组或者电池簇后集中使用.在该种情况下，一旦其中某节电池发生火灾，其释放的强热、燃烧等行为会造成周围电池温度上升，导致整个电池模组的热失控，甚至造成整个电池系统的火灾、爆炸事故[6-9].因此，在电化学储能以及变电系统等大规模应用场景中，研究锂离子电池热失控的火灾危险性并针对性开发相应的火灾抑制技术，对于电池系统的安全运行尤为重要. ...1... 然而，由于自身的物理化学性质，当不正确使用时(热滥用、电滥用和机械滥用)，磷酸铁锂电池会发生不可逆的热失控行为，存在较大的火灾危险性[1-5].在储能电站、变电站等实际运营场景中，往往将成百上千节的电池单体经过串并联后形成电池模组或者电池簇后集中使用.在该种情况下，一旦其中某节电池发生火灾，其释放的强热、燃烧等行为会造成周围电池温度上升，导致整个电池模组的热失控，甚至造成整个电池系统的火灾、爆炸事故[6-9].因此，在电化学储能以及变电系统等大规模应用场景中，研究锂离子电池热失控的火灾危险性并针对性开发相应的火灾抑制技术，对于电池系统的安全运行尤为重要. ...1... 然而，由于自身的物理化学性质，当不正确使用时(热滥用、电滥用和机械滥用)，磷酸铁锂电池会发生不可逆的热失控行为，存在较大的火灾危险性[1-5].在储能电站、变电站等实际运营场景中，往往将成百上千节的电池单体经过串并联后形成电池模组或者电池簇后集中使用.在该种情况下，一旦其中某节电池发生火灾，其释放的强热、燃烧等行为会造成周围电池温度上升，导致整个电池模组的热失控，甚至造成整个电池系统的火灾、爆炸事故[6-9].因此，在电化学储能以及变电系统等大规模应用场景中，研究锂离子电池热失控的火灾危险性并针对性开发相应的火灾抑制技术，对于电池系统的安全运行尤为重要. ...2... 然而，由于自身的物理化学性质，当不正确使用时(热滥用、电滥用和机械滥用)，磷酸铁锂电池会发生不可逆的热失控行为，存在较大的火灾危险性[1-5].在储能电站、变电站等实际运营场景中，往往将成百上千节的电池单体经过串并联后形成电池模组或者电池簇后集中使用.在该种情况下，一旦其中某节电池发生火灾，其释放的强热、燃烧等行为会造成周围电池温度上升，导致整个电池模组的热失控，甚至造成整个电池系统的火灾、爆炸事故[6-9].因此，在电化学储能以及变电系统等大规模应用场景中，研究锂离子电池热失控的火灾危险性并针对性开发相应的火灾抑制技术，对于电池系统的安全运行尤为重要. ...... 随着温度持续升高，电池内部热反应持续加快，100%与50% soc锂离子电池分别在2253 s和2611 s形成二次射流火，而对于0% soc的锂离子电池，因为电池内部能量较低，电池内部化学反应过程相对缓慢[9, 15]，其燃烧行为明显缓和，在整个过程中并未出现多次射流火现象. ...1... 本文以228 a·h的磷酸铁锂为研究对象，通过自主搭建的锂离子电池火灾燃烧实验平台研究了目标电池的火灾危险性[10]，并进一步分析了荷电状态对其火灾行为的影响规律，为锂离子电池的安全设计及火灾防控提供理论和技术支撑. ...1... 热释放速率是进行火灾危险性研究、分析样品火灾危险性的重要参数[11-13].其计算方式主要依据氧消耗原理，即通过精确测量燃烧过程中体系中的氧消耗量进而计算得到该过程的热释放速率[14]，如式(1)所示. ...2... 如图7所示，相比于100% soc锂离子电池的剧烈燃烧，50%与0% soc锂离子电池的燃烧行为较为缓和.在本次实验中，50% soc锂离子电池的热释放速率曲线存在3个明显的峰值，这与实验观察到的射流火次数相符.但是由于电池内部热反应过程较为缓慢，持续时间较长，因此其最高峰值出现在第1个峰值，为52.82 kw，约为100% soc电池峰值的53.4%.而对于0% soc的锂离子电池，在经历初次射流火后即进入持续的稳定燃烧阶段，直至最终火焰熄灭，因此整个实验过程中仅观察到一个明显的hrr峰值，为41.74 kw.对应50%与0% soc 锂离子电池的总燃烧热(10.33 mj、7.68 mj)分别为100% soc(13.94 mj)电池的74.1%和55.1%.可以看出，随着soc的降低，电池燃烧剧烈程度明显降低，对应热释放速率峰值以及总燃烧热随之降低.而电池的总燃烧热不仅与电池的燃烧剧烈程度相关，还与燃烧的持续时间相关，因此soc对电池燃烧释放总燃烧热的影响并不明显，这一特性也被其他研究者所证实，如ribiere等[12]以2.9 a·h的软包limn2o4/石墨电池为研究对象，实验研究了不同荷电状态锂离子电池的燃烧产热，研究发现50% soc的电池燃烧总热量为383 kj，高于100% soc电池的313 kj.造成该现象的原因是soc较低的锂离子电池的发生热失控对应反应物的消耗速率较低，进而延长了电池的燃烧时间. ...... 为了更好地了解锂离子电池的火灾危险性，图8比较了不同soc锂离子电池与几种常见燃料的热释放速率[12].100% soc 样品电池的标准化热释放速率峰约为2.91 mw/m2，超过汽油的标准热释放速率峰(2.2 mw/m2)，50% soc与0% soc锂离子电池燃烧时的热释放速率峰分别为1.55 mw/m2与1.22 mw/m2，介于汽油(2.2 mw/m2)与燃油(1.1 mw/m2)之间. ...1... 热释放速率是进行火灾危险性研究、分析样品火灾危险性的重要参数[11-13].其计算方式主要依据氧消耗原理，即通过精确测量燃烧过程中体系中的氧消耗量进而计算得到该过程的热释放速率[14]，如式(1)所示. ...1... 热释放速率是进行火灾危险性研究、分析样品火灾危险性的重要参数[11-13].其计算方式主要依据氧消耗原理，即通过精确测量燃烧过程中体系中的氧消耗量进而计算得到该过程的热释放速率[14]，如式(1)所示. ...1... 随着温度持续升高，电池内部热反应持续加快，100%与50% soc锂离子电池分别在2253 s和2611 s形成二次射流火，而对于0% soc的锂离子电池，因为电池内部能量较低，电池内部化学反应过程相对缓慢[9, 15]，其燃烧行为明显缓和，在整个过程中并未出现多次射流火现象. ...1... 图4(d)给出了电池的电压变化趋势，在本次实验中发现，电池的电压跳水时间较晚于其安全阀破裂时间，这是由于造成电压掉落的主要原因是隔膜收缩熔融，而隔膜的收缩温度通常在130 ℃以上[16].而电池的sei膜在90 ℃时即发生分解，造成负极活性材料与电解液反应并产生一定量的气体，造成电池内部压力持续升高[17].而在电压跳水之前，随着温度的升高，电池电压表现出微量的衰减，这是由于电池的正、负极材料溶解所致.因此，在实际应用过程中，可考虑采用气体信号和电、热信号相结合的手段，对磷酸铁锂电池的热失控行为进行预测预警. ...1... 图4(d)给出了电池的电压变化趋势，在本次实验中发现，电池的电压跳水时间较晚于其安全阀破裂时间，这是由于造成电压掉落的主要原因是隔膜收缩熔融，而隔膜的收缩温度通常在130 ℃以上[16].而电池的sei膜在90 ℃时即发生分解，造成负极活性材料与电解液反应并产生一定量的气体，造成电池内部压力持续升高[17].而在电压跳水之前，随着温度的升高，电池电压表现出微量的衰减，这是由于电池的正、负极材料溶解所致.因此，在实际应用过程中，可考虑采用气体信号和电、热信号相结合的手段，对磷酸铁锂电池的热失控行为进行预测预警. ...。

2A12铝合金本构关系实验研究

工学硕士学位论文2A12铝合金本构关系实验研究李春雷哈尔滨工业大学2006年6月国内图书分类号: TG113.25国际图书分类号: 539.4.011工学硕士学位论文2A12铝合金本构关系实验研究硕士研究生：李春雷导师：庞宝君教授副导师：贾斌副教授申请学位：工学硕士学科、专业：固体力学所在单位：复合材料与结构研究所答辩日期：2006年6月授予学位单位：哈尔滨工业大学Classified Index：TG113.25U.D.C：539.4.011Dissertation for the Master Degree in Engineering EXPERIMENTAL INVESTIGATION INTO THE CONSTITUTIVE RELATIONSHIP OF 2A12 ALUMINUM ALLOYCandidate：Li ChunleiSupervisor：Prof. Pang BaojunAssociate Supervisor：Prof. Jia BinAcademic Degree Applied for：Master of Engineering Specialty：Solid MechanicsAffiliation：Center for Composite Materialsand StructureDate of Defence：June, 2006Degree-Conferring-Institution：Harbin Institute of Technology哈尔滨工业大学工学硕士学位论文摘要材料的本构关系，尤其是动态载荷下的弹塑性本构关系一直是材料与力学领域研究的重点。

2A12铝合金由于低密度高强度的特点，广泛应用于航空、航天以及军工部门等领域。

随着计算机和数值计算方法的发展，数值模拟成为研究工程中问题的重要手段，但是材料的动态本构关系一直是束缚其发展的瓶颈。

航空飞行器用减压阀特性研究方法探析

航空飞行器用减压阀特性研究方法探析杨星1严洪英2（1.国营第570厂四川宜宾644000；2.海军装备部四川宜宾644000）摘要：复杂的航空环境对航空器用减压阀的特性、振动、噪声及疲劳寿命特性等提出了更高的要求。

本文针对目前国内外军用航空器用减压阀在设计和使用过程中的静、动态特性分析方法，以及振动、噪声和疲劳寿命设计等进行整理和分析，并针对我国目前航空器用减压阀疲劳寿命设计方法和疲劳损伤积累理论等进行综述。

综合国内外发展历程和发展现状，数值仿真和工程计算技术的发展使得目前特性研究方法已经日渐成熟。

关键词：航空器减压阀特性振动和噪声疲劳寿命中图分类号：V228文献标识码：A文章编号：1674-098X(2021)08(b)-0004-05Brief Analysis of Research Methods on the Characteristics of Pressure Reducing Valve for Aviation AircraftYANG Xing1YAN Hongying2(1.State Owned No.570Plant,Yibin,Sichuan Province,644000China;2.Equipment Department of People'sLiberation Army Navy,Yibin,Sichuan Province,644000China)Abstract:The complex aviation environment puts forward higher requirements on the characteristics,vibration, noise and fatigue life characteristics of aircraft pressure reducing valves.In this paper,the static and dynamic characteristic analysis methods,vibration,noise and fatigue life design of pressure reducing valves for military aircraft at home and abroad are sorted and analyzed,and the fatigue life design methods and fatigue damage accumulation theory of pressure reducing valves for aircraft in China are summarized.The development of numerical simulation and engineering calculation technology makes the characteristic research methods mature day by day.Key Words:Aviation aircraft;Pressure reducing valve;Characteristic;Vibration and noise;Fatigue life静态分析和动态分析是航空用减压阀特性分析的两个重要方面。

热电制冷技术在锂离子电池热管理上的应用研究

热电制冷技术在锂离子电池热管理上的应用研究崔灿;申利梅;陈焕新【摘要】锂离子动力电池具有电压平台高、比能量大、充放电效率高及寿命长等优势,已经在新能源汽车领域得到了普遍应用.但它的散热问题成了制约其发展的瓶颈.目前锂离子电池的散热有多种散热方式,鉴于热电制冷具有冷热端灵活转换,高可靠性,安全性等优点,本文将热电制冷技术应用于蓄电池的散热,基于电池的发热机理提出了一种新的散热优化方案,并通过ICEPAK进行了热仿真分析.研究发现该方案能够在环境温度为40℃的恶劣条件下,使电池表面平均温度维持在30℃左右,能够适应电池的散热需求,为日后热电技术在汽车领域的应用提供了参考.【期刊名称】《制冷》【年(卷),期】2017(036)003【总页数】6页(P13-18)【关键词】动力电池;散热;热电制冷;热仿真【作者】崔灿;申利梅;陈焕新【作者单位】华中科技大学中欧清洁与可再生能源学院,湖北武汉430074;华中科技大学能源与动力工程学院,湖北武汉430074;华中科技大学能源与动力工程学院,湖北武汉430074【正文语种】中文【中图分类】TB61+9Abstract ： Lithium-ion battery has a high voltage platform，high energy，high charge and discharge efficiency and long life and so on，which have been widely applied in the field of new energy vehicles．However，theheat dissipationhave become the bottleneck problem which limited the development of battery．There have been many ways to cooling the battery until now．For the TEC(thermoelectric cooling) have the advantage of flexible conversion between hot and cold side，high reliability and high safety，we use the TEC to the management of the battery，we proposed a new structure which based on the thermal generation of the battery anddo the thermal simulation in the Ansys．Our research find the TEC can maintain the battery 30℃ in the environment temperature of 40℃，which meet the requirement of thermal dissipation and can be the reference for using on the electric vehicles．Key words：Lithium-ion battery；Dissipation；TEC；Thermal simulation锂离子电池因比能大、循环寿命长、低温效应好等优点是目前首选的动力蓄电池，是提高整车性能和降低成本的关键。

1、下载文档前请自行甄别文档内容的完整性，平台不提供额外的编辑、内容补充、找答案等附加服务。
2、"仅部分预览"的文档,不可在线预览部分如存在完整性等问题,可反馈申请退款(可完整预览的文档不适用该条件!)。
3、如文档侵犯您的权益，请联系客服反馈,我们会尽快为您处理(人工客服工作时间：9:00-18:30)。

Experimental investigation and modelling of a buoyant attached plane jet in a roomGuangyu Cao a,*,J.Kurnitski a ,M.Ruponen b ,O.Seppänen aa Department of Mechanical Engineering,Helsinki University of Technology,02150ESPOO,Finland bHalton Oy,Haltonintie 1-3,47400Kausala,Finlanda r t i c l e i n f o Article history:Received 9July 2008Accepted 24January 2009Available online 13February 2009Keywords:BuoyantAttached plane jet ModellingVelocity decay Jet regiona b s t r a c tBuoyant attached jets are widely used in various types of supply air devices especially in ofﬁce buildings.This study focuses on a two-dimensional cooled attached jet characteristic,including mean ﬂow ﬁeld structure,speciﬁcation of the jet regions and maximum velocity decay.A new superimposing model is derived to predict the maximum velocity decay and validated by measurement results.The measurement results demonstrate that the intermediate region of a buoyant jet does exist when an inner layer extends downstream of the jet slot.In addition,by assuming that the buoyant force is the main extra force on the jet ﬂow in the acceleration process,the superimposing model predicted the maximum velocity decay with precise accuracy in a Reynolds number range of 667–4000,based on slot heights of 20and 30mm and slot velocities of 0.50,1.00and 2.00m/s.At a distance of 1000mm from the slot,the velocity proﬁle displays a self similarity character like an isothermal turbulent jet.In the ﬁnal region,where the buoyancy ﬂux completely dominates the jet,the jet behaved like a plume with an unstable ﬂow ﬁeld.Ó2009Elsevier Ltd.All rights reserved.1.IntroductionThe turbulent attached plane jet,also called the plane wall jet,has been applied in ventilated and air conditioned space for many years.In isothermal conditions,the jet behavior is well understood regarding both the structure and mechanisms [1–4].However,non-isothermal conditions pose a challenging problem in the pre-diction of the buoyant attached jet behavior.One of the attached jet applications in rooms is the active chilled beam which enables a plane jet discharging on the ceiling so that the high-velocity re-gion is restricted to the ceiling,thus possibly freeing the occupied zone from draught [5,6].Under normal design conditions,the jet stays attached to the ceiling until the opposite wall is reached,when it is then deﬂected downwards into the wall-ﬂoor zone.Hence,the maximum jet velocity should be predicted accurately to control the local draught rate in a room by the attached plane jet.Fig.1shows the division of the buoyant jet into three regions:the non-buoyant region,the intermediate region and the buoyant region.The earlier study showed that the non-buoyancy region re-fers to the same region as a jet ‘starting length’[7].Basically,the jets in a quiescent ambient can be classiﬁed according to the rela-tive importance of the initial momentum ﬂux and the initial spe-ciﬁc buoyancy ﬂux [8].In addition,it has been found that relative inﬂuence of inertia and buoyancy forces resolves the strat-iﬁed ﬂow characteristics in a room [9].At present,different computational ﬂuid dynamics (CFD)simu-lation models were developed and implemented in turbulent jet studies from small scale to fully scale simulation [10,11].However,the validation of the CFD models could be case dependent in the engineering applications [12–17].Moreover,the experimental studies were always carried out to validate the jet models or the CFD simulation results in the jet ﬂow investigations.The experi-mental method is still considered as a reliable way to directly study the jet behaviors and characteristics in many cases.In previous studies,the correlations of the buoyant jet penetra-tion distance were derived from the experimental results [18]and the heat transfer between the wall and the jet was clariﬁed exper-imentally [19].Meanwhile,different empirical and theoretical models were established and validated to predict the jet velocity distribution [20–24].Recent investigations into the buoyant at-tached jet were conducted by the application of the chilled beam [25,26].In fact,the jet equation derived from the free jet case is used to calculate the jet maximum velocity in the practical design-ing [21].It is known that the coefﬁcient of square root of two could be used in the attached jet model [27].The original jet model takes the expression as:u m ðx ÞU 0¼Kh x0:5ÆﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃﬃKAr 0=0:22:83ÁﬃﬃﬃﬃﬃﬃxKhr À1 s ð1ÞK =5was supposed in [23].Eq.(1)was called modiﬁed Regenscheit model.The Ar 0is the slot Archimedes number deﬁned as:Ar 0¼gh ðT s ÀT 1ÞT h U 0ð2Þ1359-4311/$-see front matter Ó2009Elsevier Ltd.All rights reserved.doi:10.1016/j.applthermaleng.2009.01.016*Corresponding author.Tel.:+35894513796;fax:+35894513611.E-mail address:gcao@cc.hut.ﬁ(G.Cao).Applied Thermal Engineering 29(2009)2790–2798Contents lists available at ScienceDirectApplied Thermal Engineeringjournal homepage:www.elsevie r.c o m /l o c a t e /a p t h e r m e ngThe purpose of this study was to create a practical,one-jet-equation model which can be used in both preliminary stage ofthe attached plane jet devices development and in room air distri-bution design,i.e.in ventilation or air conditioning design.Thisstudy has the potential to contribute to the development of the chilled beam,one of the low velocity attached jet applications (range from 0.5to 2.0m/s),which could provide good energy per-formance,excellent thermal comfort and silent operation [28–30].Here,the two-dimensional condition was assumed for the attached plane jet as a starting point.It should be also noted that the studied case was as much 2D as it can be in engineering applications.The main aim of this study was to obtain detailed data and knowledge of the mean ﬂow ﬁeld of a buoyant attached plane jet in a room.The speciﬁc objectives were:(1)to describe the ﬂow ﬁeld structure inﬂuenced by the buoy-ancy force over the wall surface by means of measurement;(2)to set up an effective model to predict the maximum veloc-ity decay;(3)to validate the maximum jet velocity model in a range ofReynolds numbers based on slot velocities of 0.50, 1.00and 2.00m/s;(4)to verify the self similarity characteristic of the buoyancy jetin three different jet regions.2.Derivation of the buoyant superimposing jet velocity model 2.1.Boundary conditionsHere,the non-slip condition is applied on the wall surface which bounds the attached plane air jet.The following boundary conditions are:u ¼v ¼0at y ¼0andNomenclatureLatin symbols a acceleration (m 2/s)b coefﬁcient of the non-buoyant zone model Ar Archimedes numberF bbuoyancy force (N)g gravitational acceleration (9.81m 2/s)h air jet slot height (m)K coefﬁcient which will depend on the slot Reynolds num-ber if less than 104K e coefﬁcient of buoyant jet due to temperature differenceL nondistance of the non-buoyant region P ctotal cooling power (W)P hheat load power of thermal dummies (W/m 2)q ssupply air ﬂow rate (L/s)Q air volume ﬂow (L/s)Re Reynolds numberT 0initial jet temperature (°C)T ajet ambient air temperature (°C)T jjet temperature (°C)T refbuoyancy reference temperature T ssupply air temperature (°C)T wwall surface temperature (°C)T 1room temperature (°C)u m (x)local maximum air velocity at a distance x from the jet slot (m/s)U 0slot average air velocity (m/s)u lbm ðx Þlocal buoyant jet maximum velocity (m /s)x jet horizontal traveling distance (m)x 0distance to virtual origin (m)x pjet potential core distance (m)y vertical distance from initial point (m)y 1/2the value of y at which the velocity u is half its center-line valueGreek symbols q density of ﬂuid (kg/m 3)t kinematic viscosity (m 2s À1)d boundary layer thickness d Q supply air uncertainty (%)d T measurement room temperature uncertainty (%)d u xmeasured velocity uncertainty (%)d u Anemometer uncertainty (%)d x measurement location uncertainly (%)h temperature g equal to y /y 1/2Subscripts 1,2,3coefﬁcient number a ambient air c cooling j jet 0jet slot p potential core ref buoyancy reference s jet slot t turbulent w wall surfaceG.Cao et al./Applied Thermal Engineering 29(2009)2790–27982791u¼v¼0at x¼1;y¼1For the cooling jet over the well insulated wall surface,the heatﬂux between the wall and the jet could be neglected.Thus:qw¼0at y¼0T w¼T jet at y¼0The above thermal boundary condition was presented also in[31].2.2.Buoyancy superimposing modelEarlier study showed a superimposing method to predict the maximum velocity of the attached plane jet in the near wall zone [27].Here,a new buoyant superimposing model is set up to predict the maximum jet velocity decay which consists of the isothermal part and the buoyant part.The isothermal part can be expressed by the jet similarity equation:u mðxÞ¼U0D1ðxþx0Þpð3Þwhere p represents the unknown value of the similarity exponents; andÀ0.48andÀ0.5were reported as suitable value for jet applica-tion[32–34].However,experimental studies show that the jet similarity does exist but may be expressed in a alternate way by using different similarity exponents.The relationship between the local maximum velocity and the downstream position could be expressed as[35]: u mðxÞ¼D2ðxþx0Þbð4Þwhere b is the similarity exponent.The averaged b in Eq.(4)is pro-ven to beÀ0.555by experimental results.In other studies,it was even found that the incomplete similarity exponent,À0.6,exists be-tween the inner and outer zones[36,37].Here,the similarity exponent,0.55,in the isothermal part of superimposing model is used in Eq.(3).The relationship between the jet virtual origin and the jet slot Reynolds number is known [32]:x0¼D3þD4Reð5ÞThe jet slot Reynolds number is deﬁned as:Re¼U0hmð6ÞThe parameters,D3and D4,were determined by two points test data.Here0.54and0.0021are used in Eq.(5).After substitution of D3and D4,Eq.(6)becomes:x0h¼0:54þ0:0021Reð7ÞA simpliﬁed analysis was carried out for a small bulk of air jet to derive the second part of the model.The buoyancy force acting ver-tically on a certain volume of airﬂow,Q,with temperature T and density q which are different from those of the surroundingﬂuid T0and q0,isF b¼Qgðq0ÀqÞð8ÞThe air density could be inversely proportional to the air tem-perature as:q11/T.Then Eq.(8)becomesF b¼Qg qTT0À1ð9ÞThe buoyant force produces a vertical acceleration a expressed as:a¼F bQ qð10ÞSubstituting for F b yieldsa¼g D Tð11Þwhere D T=TÀT0.If the vertical distance between this given volume and a refer-ence point is x,then the volume airﬂow will move vertically at avelocity u0which is introduced to represent the additional buoyantvelocity relative to the reference point expressed as:ðu0Þ2¼2xað12ÞSubstituting for a from Eq.(10),obtain:ðu0Þ2¼2gx D Tð13ÞThenu0¼2gx D TT01=2ð14ÞEq.(14)presents the velocity due to the buoyancy force actingvertically on a unit volume of a different temperature from the sur-roundings.Here,a temperature decay coefﬁcient,K e,is introducedto calculate the buoyancy acceleration velocity at a distance of xdownstream to the jet slot.Then Eq.(15)takes the expression withK e as:u0¼K egx D TT01=2ð15ÞCombining the isothermal velocity and the buoyancy velocity,the superimposing model becomes:u lbmðxÞ¼u mðxÞþu0¼U0D1ðxþx0ÞpþK egx D TT01=2ð16ÞThe parameters D1and K e are determined by one point test data.Preliminary test show that D1takes0.51and0.70for the cases with20mm and30mm slot height and K e uses0.53for cases by bothslot heights.The superimposing model consisting of Eq.(3)to Eq.(16)couldbe applied by using the following assumptions:(1)the air en-trained by the jet has a room air temperature;(2)the only forceopposing the jet is the wall shear stress and the free shear stress;(3)the turbulent jet momentum in the mixing process is approxi-mately preserved;(4)the attached plane jet with a low Reynoldsnumber has a similar similarity exponent to the high Reynoldsnumber jet.3.Experimental facilities and instrument3.1.The test chamberFig.2displays the layout of the test chamber and measurementpoint distribution.The test chamber consists of a well insulatedwall mounted with a jet supply device and nine adjustable dum-mies speciﬁed in[38].Fig.3shows the cross section of the air supply device and themeasured jetﬂowﬁeld.The supplied air is ducted by a variable-speed frequency-controlled centrifugal fan.3.2.Measurement conditionsIn the measurement,various jet velocities,ranging from0.50to2.00m/s,can be obtained by regulation of the outﬂow restrictionfrom30to120L/s.The turbulence intensity at the slot exit centeris approximately0.95%.Table1displays the six experimentalconditions.2792G.Cao et al./Applied Thermal Engineering29(2009)2790–27983.3.Measurement instrument and accuracyIn this study,the mean jet velocities have been measured using seven omni-directional spherical anemometers.The probe of the omni-directional spherical anemometer has a 2mm diameter.The digital HT400recording system collects signals of anemometer and thermometer with frequency 5Hz.The anemometers have upper frequency 0.5Hz by which it means that it can properly mea-sure velocity changes slower than 0.5Hz [39,40].All these anemom-eters were calibrated using TSI’s Model 8392Certiﬁer Air Velocity Calibrator.Table 2shows the summary of the instruments accuracy.In addition,the measurement error sources may include the directional sensitivity of the anemometer,the dynamic re-sponse of the anemometer,and temperature gradients in the ﬂow [41].4.Results4.1.Mean velocity ﬁeldDuring the measurement,the velocity is recorded every second and the average velocity is obtained over each 3-min period.Fig.4shows the whole downstream jet structure.At the 0.50m/s slot velocity,the jet structure is slightly turbulent.When increase the slot velocity from 0.50to 2.00m/s,the whole jet ﬂow ﬁeld looks like a laminarﬂow.Fig.3.Schematic view of experimentalsetup.Fig.2.Test chamber layout and measurement point distribution.Table 1Experimental condition in the six measured cases.Slot height (mm)Slot V (m/s)q s (L/s)D T (°C)T Supply (°C)T Room (°C)P c (W)P h (W)Re Ar 0200.5020À320.0±0.2°C 23.0±0.2°C À72±7726670.0079541.0040À320.5±0.2°C 23.5±0.2°C À144±1414413330.0019852.0080À320.0±0.2°C 23.0±0.2°C À288±2928826670.000497300.5030À319.0±0.2°C 22.0±0.2°C À108±1110810000.0029971.0060À318.6±0.2°C 21.6±0.2°C À216±2221620000.0007482.00120À318.9±0.2°C21.9±0.2°CÀ432±4343240000.000748G.Cao et al./Applied Thermal Engineering 29(2009)2790–27982793Fig.4.Air velocity distributions for À3°cooling jet with 20mm slot height.Table 2Accuracy of the measurement instruments.Velocity (m/s)Anemometer (m/s)Temperature sensor (°C)Supplied air (L/s)Traversing device (mm)Accuracy 1.25–7.5 2.0%of reading ±0.025±0.1±3.6(3.0%of supply air)±10.15–1.252.0%of reading ±0.01±0.1±1.8(3.0%of supply air)±1Fig.5.Air velocity distributions for À3°cooling jet with 30mm slot height.2794G.Cao et al./Applied Thermal Engineering 29(2009)2790–2798Fig.5displays that the whole jetﬂowﬁeld resembles a more laminar-like structure than in the case of the20mm slot height. By using10%of the slot velocity as the outermost layer velocity, the whole jet thickness is then conﬁned by the wall surface and the line with0.05,0.1and0.2m/s.4.2.Predicted maximum jet velocityIn this section,the calculation results by the modiﬁed Regensc-heit model Eq.(1)and the superimposing model Eq.(16)are com-pared with the measured results in six cooling cases.Fig.6displays the measured and calculated center line velocity decay.These results indicate that the calculated results by the superimposing modelﬁt closely to the measured data with only marginal errors.Most of the results by the modiﬁed Regenscheit model scatter outside the margins of the measurement result.4.3.Velocity proﬁlesTo clarify the difference between different buoyant jet regions, the measured velocity proﬁles are collected and compared with the theoretical turbulent velocity proﬁle by the equation derived by Schwarz and Cosart[35]:uðxÞu mðxÞ¼exp½À0:937ðgÀ0:14Þ2 ð17ÞG.Cao et al./Applied Thermal Engineering29(2009)2790–27982795Fig.7depicts measured velocity proﬁles and theoretical pro-ﬁles.By0.50m/s slot velocity with20mm slot height,the mea-sured velocity proﬁleﬁts the theoretical proﬁle closer at the distance of2000mm than the distance of1000mm and 3000mm from the jet slot.5.Discussions5.1.The buoyant jet regionsBy the deﬁnition of the three regions in the buoyant jet struc-ture,Fig.4shows that there have no obvious distinguishing points that could separate the three regions of the buoyant jet. At different slot jet velocities,the inner layer attached to the wall surface could be observed below the jet center line.The measurement results show that the intermediate region of the buoyant jet does exist when the inner layer extends downstream of the jet slot.The furthest distance that the inner layer ex-tended is approximately500,1500and2500mm in Fig.4,corre-sponding to0.50, 1.00and 2.00m/s slot average velocities, respectively.The existence of intermediate region of the buoyant jet deﬁned by So and Aksoy[8]indicates this situation as the initial momentumﬂux and the buoyancyﬂux are of comparable importance.2796G.Cao et al./Applied Thermal Engineering29(2009)2790–2798The last region(buoyant region)could be considered as the re-gion beyond the intermediate region at a certain distance down-stream of the slot.In Fig.4,the buoyant region will cover the region further than500,1500and2500mm corresponding to 0.50,1.00and2.00m/s slot velocities,respectively.During the transition from intermediate region to buoyant region,Fig.4shows an acceleration trend of the downward jetﬂow caused by buoy-ancy force along the wall surface.In the buoyant region,the jet structure displays a somewhat turbulent character at a distance of2000,2500and3000mm downstream of the slot.5.2.Self similarity of the buoyant jetComparisons with the measured data demonstrate that the superimposing model is validated for slot velocities from0.50m/ s to2.00m/s.However,at the beginning of the jetﬂow and after 2000mm downstream to the slot,the error between the calculated results and measurement data appears.On the basis of the classi-ﬁcation of buoyant jet regions,the different characteristics of the three regions in the buoyant jet might contribute to the error. Moreover,the size and the details of the jet outlet could also con-tribute to the discrepancy between the predicted results and the measured results.Malmström[42]states that the outlet can inﬂu-ence the velocities and temperatures in the resulting jet through three different mechanisms:momentum gains or losses,jet spread and the form of temperature proﬁle.As deﬁned in[6,7],if the relative inﬂuence of momentum and buoyancy are used as the critical criteria to classify theﬂow re-gions,the non-buoyancy region refers to the situation in which momentum dominates the jet behavior.Then the jet behaves like a plume rather than a jet with a certain structure.The similarity characteristic differs from isothermal jet behavior[36].Abdulhadi and Pedersen[31]and Quintana et al.[34]already conﬁrmed that the buoyant jet has the characteristics of self similarity within a distance of12–50slot heights downstream of the jet slot.Fig.7indicates that the distance of non-buoyant region might be proportional to the slot Reynolds number and inversely propor-tional to the initial Archimedes number.The relationship between the distance of the non-buoyant region,the slot Reynolds number and initial Archimedes number could be formulated as:L non¼b1Reþb2þb3ð18ÞWithin the non-buoyant region,the maximum velocity may be calculated by the isothermal jet equation.Beyond this region,the superimposing model may be applied to predict the maximum velocity decay.To test this hypothesis,more measurements are needed for the determination of each coefﬁcient in Eq.(18).6.ConclusionsIn a vertical non-isothermal turbulent buoyant jet,the measure-ment results illustrate that differentﬂow regions can be identiﬁed at some distance downstream of the jet exit.The distance of each region could be dependent on the initial conditions,including the slot Reynolds number and the initial Archimedes number.At dif-ferent slot jet velocities,the inner layer attached to the wall surface are observed below the jet center line.A higher initial Archimedes number and lower slot Reynolds number could contribute to the thicker jet and more turbulent outer layer.The observation of the extended inner layer conﬁrms that the intermediate region of the buoyant jet does exist when the initial momentum and the buoy-ancyﬂux are of comparable importance.In the buoyant region, when the buoyancyﬂux completely dominates the jet,the jet be-haves like a plume with an unstableﬂowﬁeld.The superimposing model is derived and validated by the mea-surement results.The superimposing model could predict accu-rately the maximum velocity decay beyond the non-buoyant region in the range of667–4000Reynolds number,which is based on a slot size of20and30mm and slot velocity of0.50,1.00and 2.00m/s.The modiﬁed Regenscheit model failed to predict the non-isothermal jet maximum velocity decay with a lower Reynolds number less than4000.The measurement results demonstrate that the velocity proﬁle displays a self similarity character like an isothermal turbulent jet at the point1000mm from the slot.The proﬁle with a2m/s slot velocityﬁts the theoretical proﬁle very well.To specify the starting point of the buoyant region,more detailed measurement results may be needed.The data obtained in the work could be used to validate the responding CFD simulation for further product development.The superimposing model established here potentially could be used as a practical one-equation method to predict the jet maximum de-cay in the preliminary stage of product development,and in room air distribution design especially by attached plane jet application, i.e.chilled beam.In addition,at a given or expected room air veloc-ity,the data and the model speciﬁed in this study could be applied to improve the accuracy of the calculation by CFD software.AcknowledgementsThe authors wish to express their appreciation for the measure-ment support provided by Halton Oy and for theﬁnancial support kindly provided by Halton Oy,the Finnish association of HVAC societies and K.V.Lindholm Foundation of Finland.References[1]N.Rajaratnam,B.S.Pani,Three-dimensional turbulent wall jet,in:Proceedingsof the A.S.C.E.,Journal of Hydraulic Division100(1974)69–83.[2]N.Rajaratnam,Turbulent Jets,Elsevier,Amsterdam,1976.[3]under,W.Rodi,The turbulent wall jet–measurements and modeling,Annual Review of Fluid Mechanics15(1983)429–459.[4]H.B.Awbi,Ventilation of Buildings,London,Chapman&Hall,1991.[5]J.Fredriksson,M.Sandberg,B.Moshfegh,Experimental investigation of thevelocityﬁeld and airﬂow pattern generated by cooling ceiling beams,Building and Environment36(2001)891–899.[6]H.Yu, C.M.Liao,H.M.Liang,K.C.Chiang,Scale model study of airﬂowperformance in a ceiling slot-ventilated enclosure:non-isothermal condition, Building and Environment42(2007)1142–1150.[7]J.C.Chen,W.Rodi,Turbulent Buoyant Jets—A Review of Experimental Data,HMT,Pergamon,London,1980.[8]R.M.C.So,H.Aksoy,On vertical turbulent buoyant jets,International Journal ofHeat and Mass Transfer36(13)(1993)3187–3200.[9]A.S.Awad,R.K.Calay,O.O.Badran, A.E.Holdo,An experimental study ofstratiﬁedﬂow in enclosures,Applied Thermal Engineering28(2008)2150–2158.[10]L.K.Voigt,Comparison of turbulence models for numerical calculation ofairﬂow in an annex20Room,Report ET-AFM2000-01,Technical University of Denmark,ISBN87-7475-225-1.[11]E.Baydar,Y.Ozmen,An experimental and numerical investigation on aconﬁned impinging air jet at high Reynolds numbers,Applied Thermal Engineering25(2005)409–421.[12]Q.Chen,Comparison of different k–e models for indoor air computations,Numerical Heat Transfer28(1995)728–734.[13]J.Srebric,Q.Chen,L.R.Glicksman,Validation of a zero-equation turbulencemodel for complex indoor airﬂow simulation,ASHRAE Transactions RP-927 (1999)414–427.[14]T.Karimipanah,Turbulent jets in conﬁned spaces,PhD thesis,Centre for BuiltEnvironment,Royal Institute of Technology,Gävle,Sweden,1996.[15]S.L.Sinha,R.C.Arora,S.Roy,Numerical simulation of two dimensional roomairﬂow with and without buoyancy,Energy and Buildings32(2000)121–129.[16]P.G.Tucker,Y.Liu,Turbulence modeling forﬂows around convex featuresgiving rapid eddy distortion,International Journal of Heat and Fluid Flow28 (2007)1073–1091.[17]K.Z.Yu,G.L.Ding,T.J.Chen,Simulation of air curtains for vertical displaycases with a two-ﬂuid model,Applied Thermal Engineering27(2007)2583–2591.[18]K.Sato,H.Osuka,I.Inoue,Maximum penetration distance of a vertical buoyantjet,International Chemical Engineering21(3)(1981)435–443.G.Cao et al./Applied Thermal Engineering29(2009)2790–27982797[19]J.W.Yang,R.D.Patel,Effect of buoyancy on forced convection in a two-dimensional wall jet along a vertical wall,Journal of Heat Transfer95(1) (1973)121–123.[20]M.Grimitlyn,Zurluftverteilung in raumen,Luft und Kältetechnik5(1970)246–256.[21]B.Regenscheit,Strahlgesetze und Raumströmung,Klima-Kälte-Technik6(1975).[22]W.Moog,Dimensionierung von Luftführungssystemen,Fortschritt-BerichteDer Zeitschriften6(49)(1978).[23]K.Klobut,J.Palonen,Ilmasuihkut Kirjallisuuskatsaus,Technology Report,Helsinki,1992.[24]S.Rama,G.Reddy,Combined natural and forced convection in a laminar walljet along a vertical plate with uniform surface heatﬂux,Applied Scientiﬁc Research31(6)(1976)455–463.[25]G.Y.Cao,J.Kurnitski,P.Mustakallio,O.Seppänen,Performance of chilled beamair distribution close to the wall,in:Proceedings of the Tenth International Conference on Air Distribution in Rooms,Helsinki,Finland,2007.[26]G.Y.Cao,J.Kurnitski,P.Mustakallio,O.Seppänen,Chilled beam’s airdistribution measurements and plane wall jet modeling,in:Proceedings of the Fifth International Symposium on Heating,Ventilating and Air Conditioning,Beijing,China2007,pp.288–295.[27]G.Y.Cao,J.Kurnitski,P.Mustakallio,O.Seppänen,Active chilled beam wall jetprediction by the free convection model,International Journal of Ventilation7(2)(2008)169–178.[28]F.Alamdari, D.J.G.Butler,P.F.Grigg,M.R Shaw,Chilled ceilings anddisplacement ventilation,Renewable Energy15(1998)300–305.[29]B.Costelloe, D.Finn,Indirect evaporative cooling potential in air–watersystems in temperate climates,Energy and Buildings35(6)(2003)573–591.[30]S.B.Riffat,X.Zhao,P.S.Doherty,Review of research into and application ofchilled ceilings and displacement ventilation systems in Europe,International Journal of Energy Research28(2004)257–286.[31]R.Abdulhadi,C.O.Pedersen,The behavior of a downward-directed heated walljet,ASHRAE Transactions2(1971)222–229.[32]R.A.Bajura, A.A.Szewczyk,Experimental investigation of a laminar two-dimensional attached plane jet,Physics of Fluids13(7)(1970)1653–1666.[33]M.B.Glauert,The wall jet,Journal of Fluid Mechanics1(1956)625–643.[34]D.L.Quintana,M.Amitay,A.Ortega,Heat transfer in the forced laminar walljet,Journal of Heat Transfer119(1997)451–458.[35]W.H.Schwarz,W.P.Cosart,The two-dimensional turbulent wall jet,Journal ofFluid Mechanics10(1961)481–495.[36]G.I.Barenblatt,A.J.Chorin,V.M.Prostokishin,The turbulent wall jet:a triple-layered structure and incomplete similarity,Proceedings of the National Academy of Sciences of the United States of America102(25)(2005)8850–8853.[37]R.I.Karlsson,J.Eriksson,J.Persson,in:Laser Techniques and Applications inFluid Mechanics:Selected Papers from the Sixth International Symposium on Applications of Laser Techniques to Fluid Mechanics,1992.[38]DIN4715-1,Chilled surfaces for rooms-part1:measuring of the performancewith freeﬂow–test rules,November1995.[39]A.K.Melikov,Z.Popiołek,F.E.Jorgensen,New method for testing dynamiccharacteristics of low velocity thermal anemometers,ASHRAE Transactions 104(1b)(1998)1490–1506.[40]T.Stannov,A.K.Melikov,Z.Popiołek,F.E.Jorgensen,Test method for describingdirectional sensitivity of anemometers for low-velocity measurements indoors,ASHRAE Transactions104(1b)(1998)1481–1489.[41]A.K.Melikov,Z.Popiolek,M.C.G.Silva,I.Care,T.Sefker,Accuracy limitationsfor low-velocity measurements and draft assessment in rooms,HVAC&R Research13(6)(2007)971–986.[42]T.G.Malmström,Archimedes number and jet similarity,in:Proceedings of theFifth International Conference on Air Distribution in Rooms,Yokohama,Japan, 1996,pp.415–422.2798G.Cao et al./Applied Thermal Engineering29(2009)2790–2798。