Der Einfluß der Diffusion auf die Selektivität der Desorption in einer Rieselfilmsäule

Zusammenfassung Bei der Verdunstung eines Zweistoffgemisches in ein inertes Trägergas in einer Rieselfilmsäule hängt der Trenneffekt nicht allein von der relativen Flüchtigkeit, sondern auch vom Verhältnis der Diffusionsgeschwindigkeiten beider Stoffe im Trägergas ab. Bei der Verdunstung von Isopropanol-Wasser-Gemischen in trockene Luft zeigte sich, daß das Verhältnis der gasseitigen Stoffübergangskoeffizienten bei großen Gasgeschwindigkeiten etwa gleich der Wurzel aus dem Verhältnis der Diffusionskoeffizienten war. Da der Alkolhol im Trägergas langsamer diffundiert als das Wasser, konnten flüssige Mischungen durch absatzweise Verdunstung mit Alkohol angereichert werden, obwohl der Alkohol leichterflüchtig war.Bei kleinen Gasgeschwindigkeiten lieferte der Gleichstrom immer höhere Stoffübergangskoeffizienten als der Gegenstrom. Beim Gleichstrom wurde der Einfluß des Diffusionskoeffizienten auf den Stoffübergangskoeffizienten mit abnehmender Geschwindigkeit größer, beim Gegenstrom wurde er schwächer.

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The influence of diffusion on selectivity of desorption in a wetted wall column

The desorption of a binary mixture into a stripping gas flowing through a wetted-wall column is not only governed by the vapour-liquid-equilibrium. Gas-phase diffusivities of the evaporating components have also to be taken into account. Batch wise stripping experiments of Propanol(2)-water-mixtures using dry air as the stripping gas showed, that at high gas rates the mass transfer coefficients were proportional to the square root of the diffusivities. Therefore it was possible to enrich the residual mixture with Propanol(2) because of its lower diffusivity, although Propanol(2) is more volatile.At low gas rates the mass-transfer coefficients were higher for cocurrent flow than for countercurrent flow. Besides at low gas rates the diffusivities had more influence on mass-transfer for cocurrent flow than for countercurrent flow.

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Abbreviations

Formelzeichen A [m²] Oberfläche des Rieselfilms2 r_ph·L - F [m²] freie Strömungsquerschnittfläche für das Gas in der Rieselfilmsäule: r _ph ² - K _g [–] kinetischer Trennfaktor - k _l [–] Kennzahl für den flüssigseitigen Widerstand - L [m] Länge der Rieselfilmsäule - n [mol/m³] molare Dichte - n _l [mol] Behältermolmenge - N _l,0 [mol] Behältermolmenge zu Beginn des Versuchs - n _i [mol/m² s] Molenstromdichte der Komponentei - N _i [mol/s] Molenstrom der Komponentei - N _g [mol/s] Molenstrom des Trägergases - p [Pa] Druck - p _i ⁰ [Pa] Dampfdruck der reinen Komponente - r [m] Radius - r _i [m] Innenradius des Rieselrohres - r ₁ [–] molarer bezogener Verdunstungsstrom, definiert in Gl. (3) - r ₁ [–] molarer bezogener Verdunstungsstrom, definiert in Gl. (9) - S ₁ [–] Selektivität der Desorption - s _l [m] Filmdicke - u [m/s] Geschwindigkeit - t [s] Zeit - V [m³/s] Volumenstrom - x [–] Molenbruch in der Flüssigkeit - y [–] Molenbruch in der Gasphase - z [m] Längenkoordinate Griechische Buchstaben _T [–] thermodynamischer Trennfaktor - [m/s] Stoffübergangskoeffizient - [–] Aktivitätskoeffizient - [m²/s] Diffusionszahl - [°C] Temperatur - v [m²/s] kinematische Viskosität - [–] Absättigung Indices a Austritt - e Eintritt - g gasseitig - i Komponente - l flüssigseitig - Ph Phasengrenze, Gleichgewicht - RFS Rieselfilmsäule - 1 Isopropanol - 2 Wasser dimensionslose Kennzahlen St _g = _g/¯u _g - Gz _g =4/ V _g/ _g·L - Sh _g = _g·2r _ph - Re _g =¯u _g·2r _ph/v_g - Sc _g =v _g/ _g - NTU _g =·A{itdn_g/N _g - Re _l =V _l/2r _i·v _l     

Optimum rib size to enhance forced convection in a horizontal triangular duct with ribbed internal surfaces

The optimum rib size to enhance heat transfer had been proposed through an experimental investigation on the forced convection of a fully developed turbulent flow in an air-cooled horizontal equilateral triangular duct fabricated on its internal surfaces with uniformly spaced square ribs. Five different rib sizes (B) of 5 mm, 6 mm, 7 mm, 7.9 mm and 9 mm, respectively, were used in the present investigation, while the separation (S) between the center lines of two adjacent ribs was kept at a constant of 57 mm. The experimental triangular ducts were of the same axial length (L) of 1050 mm and the same hydraulic diameter (D) of 44 mm. Both the ducts and the ribs were fabricated with duralumin. For every experimental set-up, the entire inner wall of the duct was heated uniformly while the outer wall was thermally insulated. From the experimental results, a maximum average Nusselt number of the triangular duct was observed at the rib size of 7.9 mm (i.e. relative rib size ). Considering the pressure drop along the triangular duct, it was found to increase almost linearly with the rib size. Non-dimensional expressions had been developed for the determination of the average Nusselt number and the average friction factor of the equilateral triangular ducts with ribbed internal surfaces. The developed equations were valid for a wide range of Reynolds numbers of 4,000 < Re _D < 23,000 and relative rib sizes of under steady-state condition. A Inner surface area of the triangular duct [m²] - A _C Cross-sectional area of the triangular duct [m²] - B Side length of the square rib [mm] - C _P Specific heat at constant pressure [kJ·kg^–1·K^–1] - C ₁, C ₂, C ₃ Constant coefficients in Equations (10), (12) and (13), respectively - D Hydraulic diameter of the triangular duct [mm] - Electric power supplied to heat the triangular duct [W] - f Average friction factor - F View factor for thermal radiation from the duct ends to its surroundings - h Average convection heat transfer coefficient at the air/duct interface [W·m^–2 ·K^–1] - k Thermal conductivity of the air [W·m^–1 ·K^–1] - L Axial length of the triangular duct [mm] - Mass flow rate [kg·s^–1] - n ₁, n ₂, n ₃ Power indices in Equations (10), (12) and (13), respectively - Nu _D Average Nusselt number based on hydraulic diameter - P Fluid pressure [Pa] - Pr Prandtl number of the airflow - _c Steady-state forced convection from the triangular duct to the airflow [W] - _l Heat loss from external surfaces of the triangular duct assembly to the surroundings [W] - _r Radiation heat loss from both ends of the triangular duct to the surroundings [W] - Re _D Reynolds number of the airflow based on hydraulic diameter - S Uniform separation between the centre lines of two consecutive ribs [mm] - T Fluid temperature [K] - T _a Mean temperature of the airflow [K] - T _ai Inlet mean temperature of the airflow [K] - T _ao Outlet mean temperature of the airflow [K] - T _s Mean surface temperature of the triangular duct [K] - T Ambient temperature [K] - U Mean air velocity in the triangular duct [m·s^–1] - _r Mean surface-emissivity with respect to thermal radiation - Dynamic viscosity of the fluid [kg·m^–1·s^–1] - Kinematic viscosity of the airflow [m²·s^–1] - Density of the airflow [kg·m^–3] - Stefan-Boltzmann constant [W·m^–2·K^–4]     

Experimental investigations of the stability limit of the helical flow of pseudoplastic liquids

The stability of the laminar helical flow of pseudoplastic liquids has been investigated with an indirect method consisting in the measurement of the rate of mass transfer at the surface of the inner rotating cylinder. The experiments have been carried out for different values of the geometric parameter = R ₁/R ₂ (the radius ratio) in the range of small values of the Reynolds number,Re < 200. Water solutions of CMC and MC have been used as pseudoplastic liquids obeying the power law model. The results have been correlated with the Taylor and Reynolds numbers defined with the aid of the mean viscosity value. The stability limit of the Couette flow is described by a functional dependence of the modified critical Taylor number (including geometric factor) on the flow indexn. This dependence, general for pseudoplastic liquids obeying the power law model, is close to the previous theoretical predictions and displays destabilizing influence of pseudoplasticity on the rotational motion. Beyond the initial range of the Reynolds numbers values (Re>20), the stability of the helical flow is not affected considerably by the pseudoplastic properties of liquids. In the range of the monotonic stabilization of the helical flow the stability limit is described by a general dependence of the modified Taylor number on the Reynolds number. The dependence is general for pseudoplastic as well as Newtonian liquids.Nomenclature C _i concentration of reaction ions, kmol/m³ - d = R ₂ –R ₁ gap width, m - F _M () Meksyn's geometric factor (Eq. (1)) - F ₀ Faraday constant, C/kmol - i _l density of limit current, A/m³ - k _c mass transfer coefficient, m/s - n flow index - R ₁,R ₂ inner, outer radius of the gap, m - Re = V _m ·2d·/µ _m Reynolds number - Ta _c = _c ·d^3/2·R ₁ ^1/2 ·/µ _m Taylor number - Z _i number of electrons involved in electrochemical reaction - = R ₁/R ₂ radius ratio - µ apparent viscosity (local), Ns/m² - µ _m mean apparent viscosity value (Eq. (3)), Ns/m² - µ _i apparent viscosity value at a surface of the inner cylinder, Ns/m² - density, kg/m³ - _c angular velocity of the inner cylinder (critical value), 1/s     

The effects of span position of winglet vortex generator on local heat/mass transfer over a three-row flat tube bank fin

The naphthalene sublimation method was used to study the effects of span position of vortex generators (VGs) on local heat transfer on three-row flat tube bank fin. A dimensionless factor of the larger the better characteristics, JF, is used to screen the optimum span position of VGs. In order to get JF, the local heat transfer coefficient obtained in experiments and numerical method are used to obtain the heat transferred from the fin. A new parameter, named as staggered ratio, is introduced to consider the interactions of vortices generated by partial or full periodically staggered arrangement of VGs. The present results reveal that: VGs should be mounted as near as possible to the tube wall; the vortices generated by the upstream VGs converge at wake region of flat tube; the interactions of vortices with counter rotating direction do not effect Nusselt number (Nu) greatly on fin surface mounted with VGs, but reduce Nu greatly on the other fin surface; the real staggered ratio should include the effect of flow convergence; with increasing real staggered ratio, these interactions are intensified, and heat transfer performance decreases; for average Nu and friction factor (f), the effects of interactions of vortices are not significant, f has slightly smaller value when real staggered ratio is about 0.6 than that when VGs are in no staggered arrangement. A cross section area of flow passage [m²] - A _mim minimum cross section area of flow passage [m²] - a width of flat tube [m] - b length of flat tube [m] - B _pT lateral pitch of flat tube: B _pT = S ₁/T _p - d _h hydraulic diameter of flow channel [m] - D _naph diffusion of naphthalene [m²/s] - f friction factor: f = pd _h/(Lu ² _max/2) - h mass transfer coefficient [m/s] - H height of winglet type vortex generators [m] - j Colburn factor [–] - JF a dimensionless ratio, defined in Eq. (23) [–] - L streamwise length of fin [m] - L _PVG longitudinal pitch of vortex generators divided by fin spacing: L _pVG = l _VG/T _p - l _VG pitch of in-line vortex generators [m] - m mass [kg] - m mass sublimation rate of naphthalene [kg/m²·s] - Nu Nusselt number: Nu = d _h/ - P pressure of naphthalene vapor [Pa] - p non-dimensional pitch of in-line vortex generators: p = l _VG/S ₂ - Pr Prandtl number [–] - Q heat transfer rate [W] - R universal gas constant [m²/s²·K] - Re Reynolds number: Re = ·u _max·d _h/ - S ₁ transversal pitch between flat tubes [m] - S ₂ longitudinal pitch between flat tubes [m] - Sc Schmidt number [–] - Sh Sherwood number [–]: Sh = hd _h/D _naph - Sr staggered ratio [–]: Sr = (2Hsin – C)/(2Hsin) - T _p fin spacing [m] - T temperature [K] - u _max maximum velocity [m/s] - u average velocity of air [m/s] - V volume flow rate of air [m³/s] - x,y,z coordinates [m] - z sublimation depth[m] - heat transfer coefficient [W/m²·K] - heat conductivity [W/m·K] - viscosity [kg/m²·s] - density [kg/m³] - attack angle of vortex generator [°] - time interval for naphthalene sublimation [s] - fin thickness, distance between two VGs around the tube [m] - small interval - C distance between the stream direction centerlines of VGs - p pressure drop [Pa] - 0 without VG enhancement - 1, 2, I, II fin surface I, fin surface II, respectively - atm atmosphere - f fluid - fin fin - local local value - m average - naph naphthalene - n,b naphthalene at bulk flow - n,w naphthalene at wall - VG with VG enhancement - w wall or fin surface     

Die freie Oberfläche einer Flüssigkeit über einer rotierenden Scheibe

Zusammenfassung Es wird eine modifizierte Form des Weissenberg-Effekts untersucht, wobei sich die viskoelastische Flüssigkeit in einem kreiszylindrischen Gefäß befindet, an dessen Boden eine Scheibe rotiert. Normalspannungsdifferenzen rufen in der Flüssigkeit eine Strömung hervor, die auf der Drehachse von unten nach oben gerichtet ist, und die freie Oberfläche wölbt sich nahe der Achse nach außen. Unter der Voraussetzung hinreichend langsamer Strömung wird eine Theorie zweiter Ordnung entwickelt. Sie führt auf elliptische Randwertaufgaben zweiter bzw. vierter Ordnung für das Geschwindigkeitsfeld der Primärströmung in Umfangsrichtung und für die Stromfunktion der Sekundärströmung in der Meridianebene. Ihnen werden äquivalente Variationsaufgaben zugeordnet und mit der Methode der Finiten Elemente numerisch gelöst. Die Gestalt der freien Oberfläche setzt sich bei geeigneter Normierung aus drei universellen Formfunktionen zusammen, die für verschiedene Füllhöhen berechnet werden. Im experimentellen Teil wird nachgewiesen, daß durch entsprechende Messungen der Auslenkung des Flüssigkeitsspiegels die unteren Grenzwerte der beiden Normalspannungskoeffizienten bestimmt werden können. Das Rheometer besitzt den Vorzug, daß die Oberflächenspannung der Flüssigkeit die Meßgröße nur unwesentlich beeinflußt.

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Some kind of Weissenberg effect is considered where the viscoelastic fluid, being within a cylindrical vessel, is set in motion by a rotating disc near the tank bottom. Because of normal-stress differences within the fluid a secondary flow arises which is directed upwards near the axis of symmetry, and thus the free surface is deformed. Under the assumption of sufficiently slow flow a second-order theory is developed. It leads to second-order and fourth-order elliptic boundary value problems for the velocity field in azimuthal direction and for the stream function of the secondary flow, respectively. Equivalent variational problems are formulated and solved by the method of finite elements. When normalized appropriately, the shape of the free surface consists of three shape functions, which are independent of any material constants. It is shown by corresponding experiments, that the zero-shear-rate normal-stress coefficients can be determined by measuring the displacement of the free surface. In this rheometer, the surface tension of the fluid causes only insignificant influence on the quantity to be measured.

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Symbole C _H [—] Verhältnis der FormfunktionenF ₂/F₁ - f [—] die Sekundärströmung treibende radiale Volumenkraft, dimensionslos - F ₀, F₁, F₂ [—] universelle Formfunktionen - Fr [—] Froude-Zahl - g [m s^–2] Erdbeschleunigung - h [—] Auslenkung der Oberfläche, aufr ₀ bezogen - H [—] dimensionslose Füllhöhe - K [—] Kennzahl der Kapillarität - r,z [m] Zylinderkoordinaten - r, z [—] dimensionslose Koordinaten - r ₀ [m] Radius des Meßbehälters - Re [—] Reynolds-Zahl - v _r, v, v_z [m s^–1] Geschwindigkeitskomponenten - We ₁, We₂ [—] Weissenberg-Zahlen - [Pa s] Nullviskosität der Flüssigkeit - [°C] Temperatur - [m] Kapillarlänge - v ₁, v₂ [Pa s²] untere Grenzwerte der Normalspannungskoeffizienten - [kg m^–3] Dichte der Flüssigkeit - [N m^–1] Oberflächenspannung - [—] Zylinderkoordinate - [—] Dissipationsfunktion der Sekundärströmung, dimensionslos - [—] Stromfunktion, dimensionslos - [—] örtliche Winkelgeschwindigkeit, dimensionslos - [s^–1] Winkelgeschwindigkeit der Scheibe     

Heat and mass transfer characteristics of simulated high moisture flue gases

The heat transfer process occurring in a condensing heat exchanger where noncondensible gases are dominant in volume is different from the condensation heat transfer of the water vapor containing small amount of noncondensible gases. In the process the mass transfer due to the vapor condensation contributes an important part to the total heat transfer. In this paper, the Colburn-Hougen method is introduced to analyze the heat and mass transfer process when the water vapor entrained in a gas stream condenses into water on the tube wall. The major influential factors of the convective-condensation heat transfer coefficient are found as follows: the partial pressure of the vapor p _v , the temperature of the outer tube wall T _w , the mixture temperature T _g , Re and Pr. A new dimensionless number Ch, which is defined as condensation factor, has been proposed by dimensional analysis. In order to determine the relevant constants and investigate the convection-condensation heat and mass transfer characteristics of the condensing heat exchanger of a gas fired condensing boiler, a single row plain tube heat exchanger is designed, and experiments have been conducted with vapor-air mixture used to simulate flue gases. The experimental results show that the convection-condensation heat transfer coefficient is 1.52 times higher than that of the forced convection without condensation. Based on the experimental data, the normalized formula for convention-condensation heat transfer coefficient is obtained. A heat transfer area m² - Ch condensation factor - c _p specific heat at constant pressure, J/(kg·K) - G mass flux Kg/(m²·s) - h heat transfer coefficient W/(m²·K) - J J-factor - Nu Nusselt number - pa pressure - Pr Prandtl number - Q heat transfer rate - q heat flux W/m² - r latent heat, kJ/kg - Re Reynolds number - Sc Schmidt number - T temperature, C or K - heat conductivity m W/(m·K) - density, kg·m³ - g gas - h moistened hot air - i interface - v vapor - w water     

Die scheinbare Wärmedurchgangszahl von Plattenwärmeaustauschern

Zusammenfassung Um bei Platten-Wärmeaustauschern die übertragene Wärme richtig zu erhalten, muß das Produkt aus der Wärmedurchgangszahl k=1/(1/a ₁ + / + 1/a ₂), der gesamten Heizfläche A und der mittleren logarithmischen Temperatur differenz _m im allgemeinen noch mit einem Korrekturfaktor Z multipliziert werden, der von der Fließweganzahl n und der dimensionslosen Größe =(k A₀)/(M c) abhängt: Q=Z (kA_m) =k_sA_m (k_s=scheinbare Wärmedurchgangszahl). Unter der Voraussetzung, daß Gegenstrom herrscht und das Umwälzverhältnis U=M₁c₁/M₂c₂=1 ist, konnte diese Funktion Z=Z (n, )=k_s/k jetzt im Anschluß an eine frühere Rechnung für die reine Hintereinanderschaltung beliebig vieler Fließwege bestimmt werden. Die gefundenen Formeln weisen für gerade und ungerade Fließweganzahlen kleine Unterschiede auf. Doch verlaufen beide Kurvenscharen so gleichartig, daß sie sich gegenseitig ergänzend sehr gut ineinander fügen. — Eine wichtige Folgerung aus den Rechnungsergebnissen ist, daß die scheinbare Wärmedurchgangszahl k_s bei einem gegebenem Austauscher für jeden Massenstrom einen absoluten Höchstwert hat.

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The apparent overall heat transfer coefficient of plate heat exchangers

In order to get the correct value of the transfered heat Q of plate heat exchangers one must multiply the product of the overall heat transfer coefficient k=(1/a ₁ + / + 1/a ₂), the total heating area A and the logarithmic mean temperature difference _m with a correction factor Z: Q=Z · k · A · _m =k_s · A_m, where k_s means the so called apparent overall heat transfer coefficient. Z is, as was shown in a previous paper, a function of the numbers of flow channels and the dimensionless quantity =(k · A₀)/M · c. In this paper, assuming counter flow and the validity of the relation U=M₁c₁/M₂c₂=1, the correction factor Z is determined for the pure series con nexion of any desired number of flow channels. — An important conclusion drawn from our results is, that for a given heat exanger, k_s has an absolute maximum value for every mass flow rate.

     

[0,k_i]_1~m-FACTORIZATIONS ORTHOGONAL TO A SUBGRAPH

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Effect of hydrodynamic forces on the pressure-difference equation

Because of the influence of hydrodynamic forces, the difference in macroscopic pressure which exists, at static equilibrium, between two immiscible phases located in a porous medium may be different from that which pertains during flow. In this paper, the concept of relative pressure difference, together with a new pressure-difference equation, is used to investigate the impact that the hydrodynamic forces have on the difference in macroscopic pressure which pertains when two immiscible fluids flow simultaneously through a homogeneous, water-wet porous medium. This investigation reveals that, in general, the equation defining the difference in pressure between two flowing phases must include a term which takes proper account of the hydrodynamic effects. Moreover, it is pointed out that, while neglect of the hydrodynamic effects introduces only a small amount of error when the two fluids are flowing cocurrently, such neglect is not permissible during steady-state, countercurrent flow. This is because failure to include the impact of the hydrodynamic effects in the latter case makes it impossible to explain the pressure behaviour observed in steady-state, countercurrent flow. Finally, the results of this investigation are used as a basis for arguing that, during steady-state, countercurrent flow, saturation is uniform, as is the case of steady-state, cocurrent flow.Roman Letters a parameter in Equation (18) - k absolute permeability, m² - k _i effective permeability to phasei;i=1, 2, m² - k _ij generalized effective permeability for phasei;i, j=1, 2, m² - p _d p ₂–p ₁=difference in macroscopic pressure between two flowing phases, N/m² - p _i pressure for phasei;i=1, 2, N/m² - p _h hydrodynamic contribution to difference in macroscopic pressure which exists during flow, N/m² - P _c macroscopic static capillary pressure, N/m² - R ₁₂ function defined by Equation (18) - S _i saturation of phasei;i=1, 2 - S _n normalized saturation of phase 1 - t time, s - u _i flux of phasei;i=1, 2m³/m²/s - x distance in direction of flow, m Greek Letters _R relative pressure difference - _i k _i / _i =mobility of phasei;i=1, 2m²/Pa·s - _ij k _ij / _j =generalized mobility of phasei;i, j=1, 2m²/Pa·s - _i viscosity of phasei;i=1, 2, Pa·s - porosity     

Free convection effects on the oscillatory flow of water at 4° C. past an infinite vertical,porous plate with constant suction

An analysis of a two-dimensional flow of water at 4 °C past an infinite vertical, porous plate is presented under the following conditions — i) suction velocity normal to the plate is constant, ii) the free stream oscillates in time about a constant mean, iii) the plate temperature is constant, iv) the difference between the temperature of the plate and the free stream is moderately large causing free convection currents. — Approximate solutions to coupled non-linear equations are derived for the mean velocity, the mean temperature, the mean skin-friction, the mean rate of heat transfer, the transient velocity and the transient temperature, the amplitude and the phase of the skin-friction and the rate of heat transfer. The mean flow of water at 4 °C is compared with that of water at 20 °C in a quantitative manner for both G >0 (cooling of the plate) and G < 0 (heating of the plate). — It is observed that owing to a fall in the temperature of the water from 20 °C to 4 ° C, there is a fall in the mean skin-friction when the plate is being cooled by the free convection currents, and a rise in the mean skin-friction when the plate is being heated by the free convection currents. The amplitude of the skin-friction, for water at 4°C, remains the same for both G > <0 whereas greater cooling of the plate causes a rise in the amplitude of the rate of heat transfer ¦Q ¦ /E and greater heating of the plate causes a fall in ¦ Q ¦ /E.

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Zusammenfassung Die zweidimensionale Stromung von Wasser bei 4 °C an einer unendlichen senkrechten Wand wird unter folgenden Bedingungen untersucht: 1) konstante Absauggeschwindigkeit normal zur Wand, 2) zeitliche Schwankungen der Freistromgeschwindigkeit um einen Mittelwert, 3) konstante Wandtemperatur, 4) mäßige Temperaturdifferenz zwischen Platte und Freistrom zur Erzeugung freier Konvektion. — Näherungslösungen der gekoppelten nichtlinearen Gleichungen sind abgeleitet für die mittlere Geschwindigkeit, die mittlere Temperatur, die mittlere Wandreibung, die mittlere Wärmeübertragung, die nichtstationäre Geschwindigkeit und Temperatur und die Amplitude und Phase der Wandreibung und der Warmeübertragung. Die Strömung von Wasser bei 4°C is quantitativ verglichen mit der bei 20°C für G > 0 (Kühlung der Platte) und G < 0 (Heizung der Platte). — Erniedrigung der Temperatur von 20°C auf 4°C ergibt geringere Wandreibung bei Kühlung und höhere Wandreibung bei Heizung der Platte. Für Wasser von 4°C bleibt die Amplitude der Wandreibung für G < 0 gleich; stärkere Kühlung ergibt einen Anstieg in der Amplitude der Warmeübertragung ¦Q¦/E, starkere Heizung einen Abfall in ¦q¦/E.

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Nomenclature ¦B¦ amplitude of the skin-friction - C_p specific heat at constant pressure - E Eckert numer {U ₀ ² /c_p(T'_w–T')} - g_x acceleration due to gravity - G Grashoff number {vg_x(T'_w–T')/u₀v ₀ ² } - k thermal conductivity - M_r, M_i fluctuating parts of the velocity profile - P Prandtl number,c _p /k - p pressure - q' rate of heat transfer - ¦Q¦ amplitude of the rate of heat transfer - t' time - T' temperature of fluid - T'_w temperature of the plate - T' temperature of the fluid in the free-stream - T_r,T_i fluctuating parts of the temperature profile - u',v' velocity components in the X8,y' directions - U' free stream velocity - U₀ amplitude of free stream fluctuations - u₀ mean velocity - v₀ suction velocity - x', y' coordinate system - ' frequency of free stream oscillations - non-dimensional frequency,'/vsk0/2 - ' skin-friction - ₀ mean tempeature - ₁ amplitude of the temperature fluctuations - phase angle of the skin-friction - ₁ coefficient of volume expansion - ' density of fluid in the boundary layer - ' density of fluid in the free-stream - viscosity     

Injection moulding of plastics

Summary In continuation of a previous investigation a simple analytical expression is derived in closed form for the thickness distribution of the freeze-off layer which is vitrified at the (flat) wall of an oblong rectangular cavity. As has been pointed out previously, this layer is marked for amorphous polymers by the molecular orientation (birefringence pattern) in the moulded sample. One can show that a more detailed study with the aid of the coupled equations of energy and of motion will not furnish essential improvements. Problems of polymer physics like glass transition or crystallization kinetics at extreme rates of cooling and shearing must be solved first.

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Zusammenfassung In Fortsetzung einer früheren Untersuchung wurde ein einfacher analytischer Ausdruck in geschlossener Form für die Dickenverteilung der eingefrorenen Schicht abgeleitet, die an der (flachen) Wand eines langgestreckten rechteckigen Formnestes während des Einspritzvorgangs glasig erstarrt. Wie früher auseinandergesetzt wurde, wird diese Schicht bei amorphen Polymeren durch die Molekülorientierung (Doppelbrechungsmuster) im gespritzten Formteil markiert. Man kann zeigen, daß eine eingehendere Studie mit Hilfe der gekoppelten Energie- und Impulsgleichungen keine essentiellen Verbesserungen bringt. Probleme der Polymerphysik, wie Glasübergang oder Kristallisationskinetik bei extremen Abkühlungs- und Schergeschwindigkeiten, müssen erst gelöst werden.

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List of Symbols a heat diffusivity of polymer melt (averaged overT) [m²s^–1] - B breadth of mould cavity [m] - Br Brinkman number ( ) - c heat capacity of polymer melt (averaged overT) [J kg^–1 K^–1] - F ₀ Fourier number (at _i/4H ²) - h heat transfer coefficient by melt flow [J K^–1 s^–1 m^–2] - h heat transfer coefficient by layer growth [J K^–1 s^–1 m^–2] - H half height of mould cavity [m] - L length of mould cavity [m] - n exponent in eq. [18] (= 0.417) - Nu Nußelt number (2Hh/) - P pressure gradientdP/dz in mould [N m^–3] - t time [s] - t _i injection time [s] - T _g glass transition temperature of polymer [K] - T _i injection temperature of polymer melt [K] - T _l stagnation temperature [K] - T _m mould wall temperature [K] - speed of flow front during mould filling [m s^–1] - x coordinate perpendicular to mould wall [m] - z coordinate in the injection direction [m] - thickness of stagnant layer (atT _l) [m] - ₀ optically detectable freeze-off thickness [m] - ⁺ apparent layer thickness (atT _i) [m] - dimensionless freeze-off thickness (= ₀/2H) - dimensionless distance from entrance (=z/L) - _m dimensionless coordinate of layer maximum - _g dimensionless temperature (= (T _i –T _l)/(T _g –T _m)) - _i dimensionless temperature (= (T _i –T _l)/(T _i –T _m)) - _l dimensionless temperature (= (T _i –T _l)/(T _l –T _m)) - _i viscosity of polymer atT _i [N s m^–3] - _l viscosity of polymer atT _l [N s m^–3] - heat conductivity of polymer melt (averaged) [J K^–1 s^–1 m^–1] - density of polymer melt (averaged) [kg m^–3] - dimensionless time (eq. [11]) - ⁺ dimensionless parameter (eqs. [19a] and [19b]) - dimensionless layer thickness (eq. [12]) - ⁺ dimensionless parameter (eq. [20a]) - dimensionless parameter (eqs. [11a] and [11b]) Formerly at Delft University of Technology, Delft (The Netherlands).Paper presented at the Conference on Chemical Engineering Rheology, Annual Meeting of the Deutsche Rheologische Gesellschaft in Aachen, March 5–7, 1979.With 3 figures and 1 table     

Instationäre eindimensionale Wärmeleitung beim Erstarren einer Flüssigkeit mit konstanter Außentemperatur

Zusammenfassung Die Temperaturverteilung in der festen und flüssigen Phase einer erstarrenden Flüssigkeit mit eindimensionaler Wärmeleitung und konstanter Außentemperatur der festen Phase wurde mit Hilfe von Laplace-Transformationen abgeleitet und mit der Neumannschen Losung des gleichen Problems verglichen.Die Übereinstimmung zwischen der Losung im vorliegen Beitrag und der Neumannschen Lösung ist recht gut.

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Transient one dimensional heat transfer in a liquid in solidification and with constant outer surface temperature of the solid phase

The temperature distribution in the solid and liquid phase of a liquid in solidification with one dimensional heat transfer and constant outer surface temperature of the solid phase is laid down by using Laplace transforms.The agreement between the present solution and Neumanns solution of the same problem is very good.

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Formelzeichen h Schmelz- oder Erstarrungswärme [J/kg] - k Erstarrungskoeffizient [m/s^0,5] - k₁ Temperaturleitfähigkeit der festen Phase [m²/s] - k₂ Temperaturleitfähigkeit der fl. Phase [m²/s] - T_o Außentemperatur der festen Phase [K] - T_s Schmelz- Oder Erstarrungstemperatur [K] - T Temperatur der fl. Phase zur Zeit t=0[K] - t Zeit [t] - x Entfernung von der Außenfläche der festen Phase [t] - ₁ Wärmeleitzahl der festen Phase [W/m·k] - ₂ Wärmeleitzahl der fl. Phase [W/m·k] - Dicke der festen Phase, Entfernung der Erstarrungsfront von der Außenfläche der festen Phase [m] - Dichte der fl. Phase [Kg/m³]     

The floquet theory for quasi-periodic linear systems

In this paper we establish the Floquet theory for the quasi-perio-dic systemwhere A(u_1,u_2,…,u_m)is an n×n periodic matrix function of u_1,u_2.…,u_mwith period 2π,and it is of C~τ,τ=(N 1)τ_0,τ_0=2(m 1).N=(1/2)n(n 1).Meanwhile,we define the characteristic exponential roots β_1,β_2,…,β_nof(0.1),and assume thatwhere K(ω),K(ω,β)>0.k_μ,j_v.are integers,all the integers k_1,k_2,…,k_m.are not zero,i~2=-1,Then there exists aquasi-periodic linear transformation,which carries(0.1)into a li-near system with constant coefficients.     

Sensitivity of three-segment electrodiffusion probes to eddy shedding

The influence of eddy shedding on the instantaneous readings of a three-segment cylindrical electrodiffusion velocity probe was investigated in an immersed jet with a very low turbulence intensity, = 1.2%. The velocity fluctuations measured by the three-segment probe were smaller than 2.6%, and the maximum error in the flow angle estimation was 2. Vortices with the Strouhal frequency were detected by a simple electrodiffusion probe placed downstream of the three-segment probe, but no peaks with this frequency were found on the frequency spectra of the three-segment probe. From the probe response to a stepwise change of the polarization voltage the characteristic times of the transient process were estimated. List of symbols a parameter in Eq. (1) [A s^b m^-b] - A amplitude gain - b parameter in Eq. (1) - c parameter in Eq. (3) [A s^–1/2] - d probe diameter [m] - f frequency [s^–1] - f _s recording frequency [s^–1] - G power spectrum - I _k relative current through k-th segment, Eq. (2) - i total current [A] - i _k current through k-th segment [A] - N number of data samples - Re Reynolds number, - Sr Strouhal number, - t time [s] - t ₀ characteristic transient time [s] - v jet velocity [m s^-1] - v time mean value of velocity [m s^-1] - v _x, y velocity components measured by probe [m s^-1] - var variance, var - dynamic viscosity [Pa s] - density [kg m^-3] - relative deviation, [%] - flow angle, see Fig. 1 - dimensionless frequency For the financial support of this work we express our thanks to the DFG, Bonn. The assistance of Dr. Ondra Wein and Dr. Pavel Mitschka is greatly appreciated.     

Lp-estimates for the nonlinear spatially homogeneous Boltzmann equation

This paper studies L^p-estimates for solutions of the nonlinear, spatially homogeneous Boltzmann equation. The molecular forces considered include inverse k^th-power forces with k > 5 and angular cut-off.The main conclusions are the following. Let f be the unique solution of the Boltzmann equation with f(v,t)(1 + ¦v²¦)^(s ₁ ^{+ /p)/2} L¹, when the initial value f ₀ satisfies f ₀(v) 0, f ₀(v) (1 + ¦v¦²)^(s ₁ ^{+ /p)/2} L¹, for some s₁ 2 + /p, and f ₀(v) (1 + ¦v¦²)^s/2 L^p. If s 2/p and 1 < p < , then f(v, t)(1 + ¦v¦²)^{(s s} ₁)^/2 L^p, t > 0. If s >2 and 3/(1+ ) < p < , thenf(v,t) (1 + ¦v¦²)^(s(s ₁ ^{+ 3/p))/2} L^p, t > 0. If s >2 + 2C₀/C₁ and 3/(l + ) < p < , then f(v,t)(1 + ¦v¦²)^s/2 L^p, t > 0. Here 1/p + 1/p = 1, x y = min (x, y), and C₀, C₁, 0 < 1, are positive constants related to the molecular forces under consideration; = (k – 5)/ (k – 1) for k^th-power forces.Some weaker conclusions follow when 1 < p 3/ (1 + ).In the proofs some previously known L-estimates are extended. The results for L^p, 1 < p < , are based on these L-estimates coupled with nonlinear interpolation.     

Two-phase geothermal flows with conduction and the connection with Buckley-Leverett theory

Two-phase flows of boiling water and steam in geothermal reservoirs satisfy a pair of conservation equations for mass and energy which can be combined to yield a hyperbolic wave equation for liquid saturation changes. Recent work has established that in the absence of conduction, the geothermal saturation equation is, under certain conditions, asymptotically identical with the Buckley-Leverett equation of oil recovery theory. Here we summarise this work and show that it may be extended to include conduction. In addition we show that the geothermal saturation wave speed is under all conditions formally identical with the Buckley-Leverett wave speed when the latter is written as the saturation derivative of a volumetric flow.Roman Letters C(P, S,q) geothermal saturation wave speed [ms^–1] (14) - c _t (P, S) two-phase compressibility [Pa^–1] (10) - D(P, S) diffusivity [m s^–2] (8) - E(P, S) energy density accumulation [J m^–3] (3) - g gravitational acceleration (positive downwards) [ms^–2] - h _w (P),h _w (P) specific enthalpies [J kg^–1] - J _M (P, S,P) mass flow [kg m^–2 s^–1] (5) - J _E (P, S,P) energy flow [J m^–2s^–1] (5) - k absolute permeability (constant) [m²] - k _w (S),k _s (S) relative permeabilities of liquid and vapour phases - K formation thermal conductivity (constant) [Wm^–1 K^–1] - L lower sheetC<0 in flow plane - m, c gradient and intercept - M(P, S) mass density accumulation [kg m^–3] (3) - O flow plane origin - P(x,t) pressure (primary dependent variable) [Pa] - q volume flow [ms^–1] (6) - S(x, t) liquid saturation (primary dependent variable) - S ^*(x,t) normalised saturation (Appendix) - t time (primary independent variable) [s] - T temperature (degrees Kelvin) [K] - T _sat(P) saturation line temperature [K] - TdT _sat/dP saturation line temperature derivative [K Pa^–1] (4) - T _c ,T _D convective and diffusive time constants [s] - u _w (P),u _s (P),u _r (P) specific internal energies [J kg^–1] - U upper sheetC > 0 in flow plane - U(x,t) shock velocity [m s^–1] - x spatial position (primary independent variable) [m] - X representative length - x, y flow plane coordinates - z depth variable (+z vertically downwards) [m] Greek Letters _P , _S remainder terms [Pa s^–1], [s^–1] - double-valued saturation region in the flow plane - h =h _s –h _w latent heat [J kg^–1] - = _w – _s density difference [kg m^–3] - line envelope - =D _K /D ₀ diffusivity ratio - porosity (constant) - _w (P), _s (P), _t (P, S) dynamic viscosities [Pa s] - v _w (P),v _s (P) kinematic viscosities [m²s^–1] - v ₀ =kh/KT kinematic viscosity constant [m² s^–1] - ₀ =v ₀ dynamic viscosity constant [m² s^–1] - _w (P), _s (P) density [kg m^–3] Suffixes r rock matrix - s steam (vapour) - w water (liquid) - t total - av average - 0 without conduction - K with conduction     

Study of thermal flows from two-dimensional,upward-facing isothermal surfaces using a laser speckle photography technique

A laser specklegram or speckle photography technique allows a direct measurement of surface temperature gradients and provides a full field interrogation with an extremely high resolution from a single data taking. The specklegram technique has been successfully applied to investigate the natural convection heat transfer from an upward-facing isothermal plate. For a plate with a large aspect ratio of 15, both local and global Nusselt numbers have been determined from the direct measurement of local temperature gradients. The Rayleigh number, based on the length scale equivalent to the ratio of the surface area to the perimeter, has been varied from 9.0 × 10³ to 4.0 × 10⁴. The present result for the global heat transfer has shown that a 1/5-power law, i.e., Nu = C₁ Ra ^1/5, correlates the data more properly whilst previously published results showed a large scatter in the exponent, ranging from 1/8-power to 1/4-power. The proportional constant, C₁ has been determined to be 0.56 which shows a fairly good agreement with previously published theoretical results. The laser specklegram technique has shown a strong potential as a powerful and convenient method for an experimental assessment of natural convection heat transfer problems. The specklegram technique at the same time has eliminated the deficiencies of both the mass transfer analogy technique and the classical heat transfer measurement technique.List of symbols a characteristic length scale defined as a = A/P where A is the surface area and P is the perimeter of the plate edge [mm] - AR aspect ratio [L/H] - c defocusing distance [mm] - d image distance of Young's fringes from speckle negative - h thermal convection coefficient [W/m² · K] - average thermal convection coefficient [W/m² · °C] - H width of the test section measured perpendicular to the optic axis [mm] - k thermal conductivity [W/m · K] - L length of the test section measured parallel to the optical axis [mm] - n index of refraction - Nu local Nusselt number [ha/k] - global Nusselt number - Pr Prandtl number [v/] - q heat flux per unit area [W/m² · s] - Ra Rayleigh number - s fringe spacing [mm] - Sc Schmidt number [v/D] - T temperature [K] Greek symbols thermal diffusivity [m²/s] - volumetric coefficient of expansion (1/T) - v kinematic viscosity of air [m²/s] - wavelength of helium-neon laser [632.8 nm] - amount of speckle dislocation     

Symmetry and secondary bifurcation of elastic structures

We consider non-linear bifurcation problems for elastic structures modeled by the operator equation F[w;α]=0 where F:X×R^k→Y,X,Y are Banach spaces and XY. We focus attention on problems whose bifurcation equations are of the form

f_i(α₁,α₂;λ,μ)=(a_iμ+b_iλ)α_i+p_iα_i³+q_iα_i∑_j=1,j≠i^kα_j+1²+α_ih_i(λ,μ;α₁,α₂,…α_k) i=1,2,…k

which emanates from bifurcation problems for which the linearization of F is Fredholm operators of index 0. Under the assumption of F being odd we prove an important theorem of existence of secondary bifurcation. Under this same assumption we prove a symmetry condition for the reduced equations and consequently we got an existence result for secondary bifurcation. We also include a stability analysis of the bifurcating solutions.     

Messung der Wärmeleitfähigkeit von Stickstoff und Kohlenmonoxid bei hohen Temperaturen im Stoßwellenrohr

Zusammenfassung Die Meßergebnisse für die Wärmeleitfähigkeit von Stickstoff bei Temperaturen zwischen 1230 und 6000 K und Drückenzwischen 1 und 10 bar und von Kohlenmonoxid zwischen 1150 und 5000 K bei 1 bar werden mitgeteilt. Diese mit dem Stoßwellenrohr gemessenen Werte werden mit jenen verglichen, die sich aus der strengen kinetischen Gastheorie ergeben. Auch verfügbare Daten anderer Autoren werden zum Vergleich herangezogen.

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Measurement of thermal conductivity of nitrogen and carbon monoxide at high temperatures in a shock tube

The paper presents results of shock-tube measurements of thermal conductivity of nitrogen at temperatures between 1230 and 6000 K and at pressures between 1 and 10 bar and of carbon monoxide at temperatures between 1150 and 5000 K at 1 bar. Experimental results are compared with several variants of theoretical values, computed from rigorous kinetic theory, and with available data of other authors.

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Bezeichnungen (Einheiten in Klammern) a [m² s^–1] Temperaturleitzahl - C _p[J mol^–1 K^–1] molare Wärmekapazität - k [J K^–1] Boltzmann-Konstante - M [kg mol^–1] molare Masse - p bar Gesamtdruck - R [J mol^–1 K^–1] Gaskonstante - T [K] thermodynamische Temperatur - t [s] Zeit - U [J mol^–1] innere Energie - w [m s^–1] Geschwindigkeit - x [m] Ortskoordinate - x _i [1] Molanteil der Komponentei im Gasgemisch - [Wm^–1 K^–1] Wärmeleitfähigkeit - [mol m^–3] molare Konzentration Indizes i die Komponentei im Gasgemisch - g bezieht sich auf das (kalte) Gas bei der Wandtemperatur - w bezieht sich auf die feste Wand - p bei konstantem Druck Dieser Beitrag wurde auf dem Thermodynamik-Kolloquium des VDI im Oktober 1969 in Zürich vorgetragen.     

Numerical analysis of simulation accuracy for hypersonic heat transfer in subsonic jets of dissociated nitrogen

One-dimensional problems of the flow in a boundary layer of finite thickness on the end face of a model and in a thin viscous shock layer on a sphere are solved numerically for three regimes of subsonic flow past a model with a flat blunt face exposed to subsonic jets of pure dissociated nitrogen in an induction plasmatron [1] (for stagnation pressures of (10⁴–3·10⁴) N/m² and an enthalpy of 2.1·10⁷ m²/sec²) and three regimes of hypersonic flow past spheres with parameters related by the local heat transfer simulation conditions [2, 3]. It is established that given equality of the stagnation pressures, enthalpies and velocity gradients on the outer edges of the boundary layers at the stagnation points on the sphere and the model, for a model of radius R_m=1.5·10^–2 m in a subsonic jet the accuracy of reproduction of the heat transfer to the highly catalytic surface of a sphere in a uniform hypersonic flow is about 3%. For surfaces with a low level of catalytic activity the accuracy of simulation of the nonequilibrium heat transfer is determined by the deviations of the temperatures at the outer edges of the boundary layers on the body and the model. For this case the simulation conditions have the form: dU_e/dx=idem, p₀=idem, T_e=idem. At stagnation pressuresP ₀2·10⁴ N/m² irrespective of the catalycity of the surface the heat flux at the stagnation point and the structure of the boundary layer near the axis of symmetry of models with a flat blunt face of radius R_m1.5·10^–2 m exposed to subsonic nitrogen jets in a plasmatron with a discharge channel radius R_c=3·10^–2 m correspond closely to the case of spheres in hypersonic flows with parameters determined by the simulation conditions [2, 3].Translated from Izvestiya Akademii Nauk SSSR, Mekhanika Zhidkosti i Gaza, No. 2, pp. 135–143, March–April, 1990.