Non-isothermal flow of polymer melts in a curved tube

The results of a numerical study (using finite differences) of heat transfer in polymer melt flow is presented. The rheological behaviour of the melt is described by a temperature-dependent power-law model. The curved tube wall is assumed to be at constant temperature. Convective and viscous dissipation terms are included in the energy equation. Velocity, temperature and viscosity profiles, Nusselt numbers, bulk temperatures, etc. are presented for a variety of flow conditions. Br — Brinkman number - c specific heat, J/kg K - De — Dean number - E dimensionless apparent viscosity, eq. (14d) - G dimensionless shear rate, eq. (19) - k parameter of the power-law model, °C^–1, eq. (7) - mass flow rate, kg/s - m ₀ parameter of the power-law model, Pa · s ⁿ , eq. (7) - n parameter of the power-law model, eq. (7) - Nu 2r _p/ — Nusselt number, eqs. (28,31) - p pressure, Pa - Pe — Péclet number - P —(p/)/r _c — pressure gradient, Pa/m - dissipated energy, W, eq. (29) - total energy, W, eq. (30) - r radial coordinate, m - r _c radius of tube-curvature, m, fig. 1 - r _p radius of tube, m, fig. 1 - r _t variable, m, eq. (6) - R dimensionless radial coordinate, eq. (14a) - R _c dimensionlessr _c, eq. (14a) - R _t dimensionlessr _t, eq. (14a) - t temperature, °C - bulk temperature, °C, eq. (27) - t ₀ inlet temperature of the melt, °C - t _w tube wall temperature, °C - T dimensionless temperature, eq. (14c) - T _w dimensionless tube wall temperature - T dimensionless bulk temperature - u ₁ variable, s^–1, eq. (4) - u ₂ variable, s^–1, eq. (5) - U ₁ dimensionlessu ₁, eq. (18) - U ₂ dimensionlessu ₂, eq. (18) - v velocity in-direction, m/s - average velocity of the melt, m/s - V dimensionlessv, eq. (14b) - dimensionless , eq. (15) - z r _c — centre length of the tube, m - Z dimensionlessz, eq. (14e) - heat transfer coefficient, W/m² K - shear rate, s^–1, eq. (8) - — shear rate, s^–1 - apparent viscosity, Pa · s, eq. (7) - ₀ — apparent viscosity, Pa · s - angular coordinate, rad, fig. 1 - thermal conductivity, W/m K - melt density, kg/m³ - axial coordinate, rad, fig. 1 - rate of strain tensor, s^–1, eq. (8) - (—p) pressure drop, Pa     

Liquid sheet and film atomization: a comparative experimental study

Liquid atomization processes are too complex to allow a purely theoretical study. Therefore experiments are necessary to quantify droplets production. In our problem, the replacement of an original complicated flow by a simpler one, i.e. liquid metal and high gas velocity by water and low air velocity, has led to a relation for the droplet diameter, thanks to dynamical similarity and order of magnitude estimates. Observation of a liquid film disruption development by high speed photography gives some informations about the mechanism of break-up in action. Granulometric measurements by video image analysis have specified the previous dimensionless relation for the mass median diameter. Measurements concern both the film and the sheet atomization, it is shown that the control of the liquid layer thickness is of major importance to control the quality of sprays.List of symbols d droplet diameter (m) - d _mm mass median droplet diameter (m) - g acceleration due to the gravity (ms^–2) - H _g , H _l gas slit width, liquid film thickness (m) - dimensionless parameters - Q ₁ = H ₁ V ₁ liquid flow rate (m²s^–1) - Reynolds number - T time(s) - V _g , V _l gas and liquid velocity (m s^–1) - W _c channel width (m) - Weber number - _g , _l gas and liquid viscosity (kg m^–1 s^–1) - _g , _i gas and liquid density (kg m^–3) - surface tension (kg s^–2) An abridged version of this paper was presented at the 6th ICLASS (Int. Conf. on Liquid Atomization and Spray Systems), Rouen, France, 18–22 July 1994     

Rheological properties of skim milk gels at various temperatures; interrelation between the dynamic moduli and the relaxation modulus

The rheological properties of rennet-induced skim milk gels were determined by two methods, i.e., via stress relaxation and dynamic tests. The stress relaxation modulusG _c (t) was calculated from the dynamic moduliG andG by using a simple approximation formula and by means of a more complex procedure, via calculation of the relaxation spectrum. Either calculation method gave the same results forG _c (t). The magnitude of the relaxation modulus obtained from the stress relaxation experiments was 10% to 20% lower than that calculated from the dynamic tests.Rennet-induced skim milk gels did not show an equilibrium modulus. An increase in temperature in the range from 20° to 35 °C resulted in lower moduli at a given time scale and faster relaxation. Dynamic measurements were also performed on acid-induced skim milk gels at various temperatures andG _c (t) was calculated. The moduli of the acid-induced gels were higher than those of the rennet-induced gels and a kind of permanent network seemed to exist, also at higher temperatures. G storage shear modulus,N·m^–2; - G loss shear modulus,N·m^–2; - G _c calculated storage shear modulus,N·m^–2; - G _c calculated loss shear modulus,N·m^–2; - G _e equilibrium shear modulus,N·m^–2; - G _ec calculated equilibrium shear modulus,N·m^–2; - G(t) relaxation shear modulus,N·m^–2; - G _c (t) calculated relaxation shear modulus,N·m^–2; - G ^*(t) pseudo relaxation shear modulus,N·m^–2; - H relaxation spectrum,N·m^–2; - t time,s; - relaxation time,s; - angular frequency, rad·s^–1. Partly presented at the Conference on Rheology of Food, Pharmaceutical and Biological Materials, Warwick, UK, September 13–15, 1989 [33].     

A new transient method for analysis of solidification in the continuous casting of metals

Solidification processes involve complex heat and mass transfer phenomena, the modelling of which requires state-of-the art numerical techniques. An efficient and accurate transient numerical method is proposed for the analysis of phase change problems. This method combines both the enthalpy and the enhanced specific heat approaches in incorporating the effects of latent heat released due to phase change. The sensitivity and accuracy of the proposed method to both temporal and spatial discretization is shown together with closed-form solutions and the results from the enhanced specific heat approach. In order to explore the proposed method fully, a non-linear heat release, as is the case for binary alloys, is also examined. The number of operations required for the new transient approach is less than or equal to the enhanced heat capacity method depending on the averaging method adopted. To demonstrate the potential of this new finite-element technique, measurements obtained on operating machines for the casting of zinc, aluminum and steel are compared with the model predictions. The death/birth technique, together with the proper heat-transfer coefficients, were employed in order to model the casting process with minimal error due to the modelling itself.Nomenclature [A] conductance matrix - [B] matrix containing the derivative of the element shape functions - c, C _p specific heat (J kg^–1°C^–1) - effective specific heat (J kg^–1°C^–1) - f(T) local liquid fraction - f thermal load vector - H enthalpy (J kg^–1) - [H] capacitance matrix - h, h _r,h _c heat transfer coefficient (W m^–2°C^–1) - K thermal conductivity (W m^–1°C^–1) - L latent heat of solidification (J kg^–1) - l overall length (m) - N _i shape functions - Q rate of heat generation per unit volume (J m^–3) - q heat flux (W m^–2) - R residual temperature (°C) - T temperature (°C) - T _s solidus temperature (°C) - T _l liquidus temperature (°C) - T _pouring pouring temperature (°C) - T _top temperature at the top of the mould (°C) - T _w temperature of the water spray (°C) - approximated temperature (°C) - T surrounding temperature (°C) - cooling rate (°C/s) - t time (seconds) - x _i,x, y, z spatial variables (m) - t time step (s) - x element size (m) - diffusivity (m²s^–1) - density (kg m^–3) - time marching parameter - _n direction cosines of the unit outward normal to the boundary     

Instability of Buckley-Leverett flow in a heterogeneous medium

We study the simultaneous one-dimensional flow of water and oil in a heterogeneous medium modelled by the Buckley-Leverett equation. It is shown both by analytical solutions and by numerical experiments that this hyperbolic model is unstable in the following sense: Perturbations in physical parameters in a tiny region of the reservoir may lead to a totally different picture of the flow. This means that simulation results obtained by solving the hyperbolic Buckley-Leverett equation may be unreliable.Symbols and Notation f fractional flow function varying withs andx - value off outsideI - value off insideI - local approximation off around¯x - f ^–,f ⁺ values of - f _j ⁿ value off atS _j ⁿ andx _j - g acceleration due to gravity [ms^–2] - I interval containing a low permeable rock - k dimensionless absolute permeability - k ^* absolute permeability [m²] - k _c ^* characteristic absolute permeability [m²] - k _ro relative oil permeability - k _rw relative water permeability - L ^* characteristic length [m] - L ₁ the space of absolutely integrable functions - L the space of bounded functions - P _c dimensionless capillary pressure function - P _c ^* capillary pressure function [Pa] - P _c ^* characteristic pressure [Pa] - S similarity solution - S _j ⁿ numerical approximation tos(x_j, t_n) - S ₁, S₂,S ₃ constant values ofs - s water saturation - value ofs at - s ^L left state ofs (wrt. ) - s ^R right state ofs (wrt. ) - s s for a fixed value of in Section 3 - T value oft - t dimensionless time coordinate - t ^* time coordinate [s] - t _c ^* characteristic time [s] - t _n temporal grid point,t _n=n t - v ^* total filtration (Darcy) velocity [ms^–1] - W, , v dimensionless numbers defined by Equations (4), (5) and (6) - x dimensionless spatial coordinate [m] - x ^* spatial coordinate [m] - x _j spatial grid piont,x _j=j x - discontinuity curve in (x, t) space - right limiting value of¯x - left limiting value of¯x - angle between flow direction and horizontal direction - t temporal grid spacing - x spatial grid spacing - length ofI - parameter measuring the capillary effects - argument ofS - _o dimensionless dynamic oil viscosity - _w dimensionless dynamic water viscosity - _c ^* characteristic viscosity [kg m^–1s^–1] - _o ^* dynamic oil viscosity [kg m^–1s^–1] - _w ^* dynamic water viscosity [k gm^–1s^–1] - _o dimensionless density of oil - _w dimensionless density of water - _c ^* characteristic density [kgm^–3] - _o ^* density of oil [kgm^–3] - _w ^* density of water [kgm^–3] - porosity - dimensionless diffusion function varying withs andx - ^* dimensionless function varying with s andx ^* [kg^–1m³s] - _j ⁿ value of atS _j ⁿ andx _j This research has been supported by VISTA, a research cooperation between the Norwegian Academy of Science and Letters and Den norske stats oljeselskap a.s. (Statoil).     

Diffusion of moisture in deformable porous media. Anisotropic fibrous materials

This paper presents a study on the deformation of anisotropic fibrous porous media subjected to moistening by water in the liquid phase. The deformation of the medium is studied by applying the concept of effective stress. Given the structure of the medium, the displacement of the solid matrix is not taken into account with respect to the displacement of the liquid phase. The transport equations are derived from the model proposed by Narasimhan. The transport coefficients and the relation between the variation in apparent density and effective stress are obtained by test measurements. A numerical model has been established and applied for studying drip moistening of mineral wool samples capable or incapable of deformation.Nomenclature D mass diffusion coefficient [L²t^–1] - e void fraction - g gravity acceleration [Lt^–2] - J mass transfer density [ML^–2t^–1] - K hydraulic conductivity [Lt^–1] - K _s hydraulic conductivity of the solid phase [Lt^–1] - K ^* hydraulic conductivity of the deformable porous medium [Lt^–1] - P pressure of moistening liquid [ML^–1 t^–2] - S degree of saturation - t time [t] - V speed [Lt^–1] - X horizontal coordinate [L] - Z vertical coordinate measured from the bottom of porous medium [L] - z z-coordinate [L] Greek Letters porosity - ₁ total hydric potential [L] - _g gas density [ML^–3] - ₁ liquid density [ML^–3] - ₀ apparent density [ML^–3] - _s density of the solid phase [ML^–3] - density of the moist porous medium [ML^–3] - external load [ML^–1t^–2] - effective stress [ML^–1t^–2] - bishop's parameter - matrix potential or capillary suction [L] Indices g gas - 1 moistening liquid - p direction perpendicular to fiber planes - s solid matrix - t direction parallel to fiber planes - v pore Exponent * movement of solid particles taken into account     

A three-parameter model describing the behaviour of a viscoelastic liquid in a tangential annular flow

A three-parameter model describing the shear rate-shear stress relation of viscoelastic liquids and in which each parameter has a physical significance, is applied to a tangential annular flow in order to calculate the velocity profile and the shear rate distribution. Experiments were carried out with a 5000 wppm aqueous solution of polyacrylamide and different types of rheometers. In a shear-rate range of seven decades (5 10^–3 s^–1 < < 1.2 10⁵ s^–1) a good agreement is obtained between apparent viscosities calculated with our model and those measured with three different types of rheometers, i.e. Couette rheometers, a cone-and-plate rheogoniometer and a capillary tube rheometer. a physical quantity defined by:a = {1 – ( / ₀)}/ ₀ (Pa^–1) - C constant of integration (1) - r distancer from the center (m) - r ₁,r ₂ radius of the inner and outer cylinder (m) - v _r local tangential velocity at a distancer from the center (v _r = _r r) (m s^–1) - v ₂ local tangential velocity at a distancer ₂ from the center (m s^–1) - shear rate (s^–1) - local shear rate (s^–1) - ₁ wall shear rate at the inner cylinder (s^–1) - dynamic viscosity (Pa s) - _a apparent viscosity (_a = / ) (Pa s) - _a1 apparent viscosity at the inner cylinder (Pa s) - ₀ zero-shear viscosity (Pa s) - infinite-shear viscosity (Pa s) - shear stress (Pa) - _r local shear stress at a distancer from the center (Pa) - ₀ yield stress (Pa) - ₁, ₂ wall shear-stress at the inner and outer cylinder (Pa) - _r local angular velocity (s^–1) - ₂ angular velocity of the outer cylinder (s^–1)     

An approach to the estimation of polymer melt elasticity

An effective method has been proposed to estimate the primary normalstress difference versus shear rate curves at temperatures relevant to the processing conditions only from the knowledge of the melt flow index, the molecular-weight distribution and the glass transition temperature of the polymer. The method involves the use of a unified curve obtained by coalescing the elastic response curves of various grades in terms of the modified normal-stress coefficient ₁ (MFI)² and a modified shear rate . Unified curves have been reported for low density polyethylene, high density polyethylene, polypropylene and nylon.Nomenclature C ₁ constant in eq. (4) - J _e steady state compliance (cm²/dyne) - proportionality constant in eq. (6) - L load (kg) - L ₁ load (kg) at ASTM test conditions - L ₂ load (kg) at required conditions - MFI melt flow index (gm/10 min) - number average molecular weight - weight average molecular weight - z-average molecular weight - (z+1)-average molecular weight - n slope of the shear stress vs. shear curve on a log-log scale - N ₁ primary normal-stress difference (dynes/cm²) - Q molecular weight distribution expressed as - T ₁is> temperature (K) at condition 1 - T ₂is> temperature (K) at condition 2 - T _g glass transition temperature (K) - T _s standard reference temperature equal toT _g + 50 K - shear rate (s^–1) - ₀ zero-shear viscosity (poise) - apparent viscosity (poise) - density (g/cm³) - ₁₂ shear stress (dynes/cm²) - ₁₁– ₂₂ primary normal-stress difference (dynes/cm²) - _1,0 zero shear rate primary normal-stress coefficient (dynes/cm² · sec²) - ₁ primary normal-stress coefficient (dynes/cm² · sec²) - ₂ secondary normal-stress coefficient (dynes/cm² · sec²) NCL-Communication No. 3106     

Effect of solids concentration in dense suspensions of silica-iron(III) oxide and water

The rheological properties of dense suspensions, of silica, iron (III) oxide and water, were studied over a range of solids concentrations using a viscometer, which was modified so as to prevent settling of the solid components. Over the conditions studied, the material behaved according to power—law flow relationships. As the concentrations of silica and iron(III) oxide were increased, an entropy term in the flow equation was identified which had a silica dependent and an iron (III) oxide dependent component. This was attributed to a tendency to order into some form of structural regularity. A, A, B, C pre-exponential functions (K Pa^–n s^–1) - C _ox volume fraction iron (III) oxide - Q activation energy (kJ mol^–1) - R gas constant (kJ mol^–1 K^–1) - R _v silica/water volume ratio - T temperature (K) - n power-law index - H enthalpy (kJ mol^–1) - S entropy change (kJ mol^–1 K^–1) - shear strain rate (s^–1) - shear stress (Pa)     

Shear stress and normal stress measurements of aqueous polymer solutions in a cone-and-plate rheogoniometer

Summary The rheological behaviour of aqueous solutions of Separan AP-30 and Polyox WSR-301 in a concentration range of 10–10000 wppm is investigated by means of a cone-and-plate rheogoniometer. The relation between the shear stress and the shear rate is for lower shear rates characterized by a timet ₀, which is concentration dependent. Both polymers show for 4000 s^–1 < < 10000 s^–1 a behaviour similar to that of a Bingham material, characterized by a dynamic viscosity ₀ and an apparent yield stress ₀, which also depend on the concentration. The inertial forces are measured for water and some other Newtonian liquids. An explanation is given why the theoretical model developed for these forces does not match the experimental values; the shape of the liquid surface is shear rate dependent. To obtain the first normal stress difference, we have to correct for these inertial forces, the surface tension and the buoyancy. The normal forces, measured for Separan AP-30, appear to be a linear function of the shear rate for 350 s^–1 < < 3300 s^–1.

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Zusammenfassung Das rheologische Verhalten wäßriger Polymerlösungen von Separan AP-30 und Polyox WSR-301 wird in einem Konzentrationsgebiet von 10–10000 wppm in einem Kegel-Platte-Rheogoniometer untersucht. Der Zusammenhang zwischen Schubspannung und Schergeschwindigkeit wird für niedrige Schergeschwindigkeiten durch eine konzentrationsabhängige Zeitt ₀ gekennzeichnet. Für Schergeschwindigkeiten 4000 s^–1 < < 10000 s^–1 zeigen beide Polymere ein genähert binghamsches Verhalten, gekennzeichnet durch eine dynamische Viskosität ₀ und eine scheinbare Fließgrenze ₀, welche ebenfalls konzentrationsabhängig sind. Die Trägheitskräfte werden für Wasser und einige newtonsche Öle bestimmt. Die Abweichung der experimentellen Ergebnisse vom theoretischen Modell wird durch die Abhängigkeit der Gestalt der Flüssigkeitsoberfläche von der Schergeschwindigkeit erklärt. Um die Werte der ersten Normalspannungsdifferenz zu erhalten, muß man bezüglich der Trägheitskräfte, der Oberflächenspannung und der Auftriebskräfte korrigieren. Die Normalspannungen für Separan AP-30, gemessen für 350 s^–1 < < 3300 s^–1, zeigen eine lineare Abhängigkeit von der Schergeschwindigkeit.

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c concentration (wppm) - g acceleration of gravity (ms^–2) - K force (N) - K _b buoyant force (N) - K _c force, acting on the cone (N) - K ₀ dimensional constant def. by eq. [24] (N) - K _s force, def. by eq. [22] (N) - M dimensional constant def. by eq. [24] (Ns) - P _s pressure def. by eq. [17] (Nm^–2) - P ₀ average pressure in the liquid atr = 0 (Nm^–2) - P _R average pressure in the liquid atr = R (Nm^–2) - r ₁,r ₂ radii of curved liquid surface (m) - R platen radius (m) - R _w radius of wetted platen area (m) - S _x standard deviation ofx - t ₀ characteristic time def. by eq. [1] (s) - T temperature (°C) - V volume of the submerged part of the cone (m³) - v tangential velocity of liquid (ms^–1) - x distance (m) - angle (rad) - ₀ cone angle (rad) - calibration constant (Nm^–3) - shear rate (s^–1) - dynamic viscosity (mPa · s) - ₀ viscosity def. by eq. [1] (mPa · s) - contact angle (rad) - density (kgm^–3) - static surface tension (Nm^–1) - shear stress (Nm^–2) - ₀ yield stress def. by eq. [1] (Nm^–2) - _c, _p angular velocity (c = cone,p = plate) (s^–1) With 8 figures and 3 tables     

Bestimmung der zweiten Normalspannungsfunktion von Polymerschmelzen

Zusammenfassung In der axialen Druckströmung durch einen konzentrischen Ringspalt läßt sich aus der Druckdifferenz zwischen Innen- und Außenwand die zweite Normalspannungsfunktion quantitativ bestimmen. In dieser Arbeit werden die Bestimmungsgleichungen für die Scherviskositätsfunktion bzw. die zweite Normalspannungsfunktion hergeleitet. Insbesondere wird die in diesen Ableitungen vorausgesetzte Isothermie der Strömung überprüft.Für die experimentelle Ermittlung der Scherviskositätsfunktion bzw. der zweiten Normalspannungsfunktion wurde eine Meßapparatur (Ringspaltwerkzeug) entwickelt.Die Ergebnisse zeigen, daß die hier ermittelten Werte für die Scherviskositätsfunktion gut mit denen übereinstimmen, die in anderen Geometrien gemessen wurden.Die aus der Druckdifferenz zwischen Innen- und Außenwand des Ringspaltes zu berechnende zweite Normalspannungsfunktion wird als Funktion von Schergeschwindigkeit und Temperatur dargestellt. Ähnlich wie bei den Polymerlösungen ergeben sich auch bei Polymerschmelzen negative Werte für die zweite Normalspannungsfunktion.

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Summary In annular flow, the pressure difference between inner and outer wall can be used to determine the second normal-stress coefficient. In this paper, the equations for the shear viscosity and the second normal-stress coefficient are derived. Special consideration is given to the problem of isothermal flow.On the experimental side, an annular die has been designed for the determination of the shear viscosity and the second normal-stress coefficient.The results show, that the measured values of shear viscosity coincide with those, measured by other geometries. The second normal-stress coefficient, determined by the pressure difference between inner and outer wall of the annular die, is presented as a function of shear rate and temperature. As in the case of polymer solutions, negative values for the second normal stress coefficient are obtained for polymer melts.

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a (cm²/kps) Kenngröße aus dem Ansatz nachRabinowitsch (Gl. [4]) - b (cm²/s) Temperaturleitfähigkeit - c (cm⁶/kp³s) Kenngröße aus dem Ansatz nachRabinowitsch (Gl. [4]) - E (—) dimensionslose Zahl in Gl. [14] - F ₁,F ₂ (kps²/cm²) erste bzw. zweite Normalspannungsfunktion - F _2,0 (kps²/cm²) Wert der zweiten Normalspannungsfunktion für kleine Schergeschwindigkeiten - h (cm) Spaltweite - k (kcal/m h °C) Wärmeleitfähigkeit - l (cm) Länge des Ringspaltes - n ₂ (—) Kenngröße aus dem Ansatz nachCarreau (Gl. [5]) - Na (—) Nahme-Zahl - p (kp/cm²) Druckgradient, Definition s. Gl. [7a] - Pe (—) Péclet-Zahl - r, R (cm), (—) laufender Radius - r _a (cm) Außenradius des Ringspaltes - r _i (cm) Innenradius des Ringspaltes - S _ij (kp/cm²) Komponenten des SpannungstensorsS - T (°C) Temperatur - T ₀ (°C) Ausgangstemperaturniveau - (K) mittlere Temperatur - v _z (cm/s) Strömungsgeschwindigkeit - (cm/s) mittlere Strömungsgeschwindigkeit - (cm³/s) Volumendurchsatz - ₂ (s) charakteristische Zeit aus dem Ansatz nachCarreau (Gl. [5]) - (s^–1) Schergeschwindigkeit - (kps/cm²) Scherviskosität - k (—) Radienverhältnis - (—) dimensionslose Koordinate (() = 0) - (kp/cm²) Schubspannung - (—) laufende Koordinate - (–) laufende Koordinate - 0 Bezugszustand - a an der Außenwand - i an der Innenwand - r, z, Koordinatenrichtungen Auszugsweise vorgetragen auf der Jahrestagung der Deutschen Rheologen in Berlin vom 28.–30. April 1975.Mit 8 Abbildungen     

The influence of boundary layer growth on shock tube test times

Summary A theoretical and experimental investigation of the limitation on shock tube test times which is caused by the development of laminar and turbulent boundary layers behind the incident shock is presented. Two theoretical methods of predicting the test time have been developed. In the first a linearised solution of the unsteady one-dimensional conservation equations is obtained which describes the variations in the average flow properties external to the boundary layer. The boundary layer growth behind the shock is related to the actual extent of the hot flow and not, as in previous unsteady analyses, to its ideal extent. This new unsteady analysis is consequently not restricted to regions close to the diaphragm. Shock tube test times are determined from calculations of the perturbed shock and interface trajectories. In the second method a constant velocity shock is assumed and test times are determined by approximately satisfying only the condition of mass continuity between the shock and the interface. A critical comparison is made between this and previous theories which assume a constant velocity shock. Test times predicted by the constant shock speed theory are generally in agreement with those predicted by the unsteady theory, although the latter predicts a transient maximum test time in excess of the final asymptotic value. Shock tube test times have also been measured over a wide range of operating conditions and these measurements, supplemented by those reported elsewhere, are compared with the predictions of the theories; good agreement is generally obtained. Finally, a simple method of estimating shock tube test times is outlined, based on self similar solutions of the constant shock speed analysis.Nomenclature a speed of sound - A, B, C constants defined in section 5.3 - D shock tube diameter - K =/q ^m, boundary layer growth constant, see Appendices A and B - l hot flow length - m constant, =1/2 or 4/5 for laminar or turbulent boundary layers, respectively - M ₀ initial shock Mach number at the diaphragm - M _s shock Mach number at station x _s - M ₂ =(U ₀–u ₂)/a ₂, hot flow Mach number relative to the shock front - N = ₂ a ₂/ ₃ a ₃, the ratio of acoustic impedances across the interface - P pressure - P* =P _e–P ₂, perturbation pressure - q boundary layer growth coordinate defined in § 2 - Q =(W–1+S) K - r radial distance from shock tube axis - S boundary layer integral defined by equation A6 - t time - t =/ , dimensionless ratio of test times - T =l/l , Roshko's dimensionless ratio of hot flow lengths - u axial flow velocity in laboratory coordinate system, see figure 1a - u* =u _e–u₂, perturbation axial flow velocity - U shock velocity - U ₀ initial shock velocity at the diaphragm - U* =U–U ₀, perturbation shock velocity - v axial flow velocity in shock-fixed coordinate system, see figure 1b - w radial flow velocity - W =U ₀/(U ₀–u ₂), density ratio across the shock - x axial distance from shock tube diaphragm - x _s, x _s axial distance between shock wave and diaphragm - t = _I/ , dimensionless ratio of test times - X =l _I/l , Roshko's dimensionless ratio of hot flow lengths - y =(D/2)–r, radial distance from the shock tube wall - ratio of specific heats - boundary layer thickness; undefined - boundary layer displacement thickness - boundary layer thickness defined by equation A2 - characteristic direction defined by dx/dt = (u ₂–a ₂) - =(M ₀ ² +1)/(M ₀ ² –1) - viscosity - characteristic direction defined by dx/dt=(u ₂+a ₂) - density - * = _te–₂, perturbation density - Prandtl number - shock tube test time - =M ₀ ² /(M ₀ ² –1) Suffices 1 conditions in the undisturbed flow ahead of the shock - 2 conditions immediately behind an unattenuated shock - 3 conditions in the expanded driver gas - 4 conditions in the undisturbed driver gas - e conditions between the shock and the interface, averaged across the inviscid core flow - i conditions at the interface - I denotes the predictions of ideal shock tube theory - asymptotic conditions given when x _s and t - s conditions at or immediately behind the shock - w conditions at the shock tube wall - a, b, b ¹, c, d, d ¹, f, f ¹, g, g ¹, j, k, k ¹ conditions at the points indicated in figure 2     

Heat transfer to superheated steam flowing in tubes and annular channels

Heat transfer measurements with superheated steam have been performed in the pressure range of 10.5–33.3 bar, steam temperatures: 229–470 °C and heat fluxes: 2.6·10⁴–5.7·10⁵ W/m². The measured data with steam flowing in an electrically heated tube of 10 mm inner diameter and 1 m length could be correlated with two different expressions, the mean steam temperature resp. the film temperature being used as reference for physical properties. The results of measurements in annular channels, with a heated rod of 14 mm outer diameter and channel diameters of 21 and 28 mm, were 10–15% lower than the proposed tube correlations. This could be explained by the differences in velocity and temperature distributions between tubes and annular channels. A satisfactory agreement was achieved between the tube correlations and transformed annular channel results.

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Zusammenfassung Wärmeübergangsmessungen an überhitztem Wasserdampf wurden im folgenden Bereich durchgeführt: Druck: 10,5–33,3 bar, Dampftemperatur: 229–470 °C, Wärmestromdichte: 2,6 · 10⁴-5,7 · 10⁵ W/m². Die Eesultate der in einem elektrisch beheizten Rohr von 10 mm Innendurchdurchmesser und 1 m Länge durchgeführten Meß-Serie konnten durch zwei verschiedene Beziehungen je nach Bezugstemperatur der thermodynamischen Stoffwerte des Kühlmittels beschtieben werden. Als Bezugsgröße wurde die mittlere Dampf — temperatur bzw. die Filmtemperatur verwendet. Messungen, die in einem Ringspalt mit einem beheizten Stab von 14 mm Außendurchmesser und mit einem Kanaldurchmesser von 21 bzw. 28 mm durchgeführt wurden, ergaben 10–15% niedrigere Wärmeübergangszahlen, als die vorgeschlagenen Gleichungen für Rohrgeometrie. Dies kann durch den Unterschied der Geschwindigkeitsverteilungen in Rohren und Ringspalten erklärt werden. Die auf Rohrgeometrie umgerechneten Resultate der Ringspaltmessungen zeigten gute Übereinstimmung mit den vorgeschlagenen Gleichungen.

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Nomenclature

Symbol Dimension A m² surface area - C constant - c _p J/kg °C heat capacity - D m channel outer wall diameter - d m channel inner wall diameter - F m² flow area - f friction coefficient - h W/°C m² heat transfer coefficient - k W/°C m thermal conductivity - L, l m lenght - m kg/s mass flow - m exponent ofRe-number - n exponent of Pr-number - p N/m², bar pressure - Q W power - q W/m² heat flux - r m radius - t °C temperature - w m/s velocity - dimensionless factor for reference temperature - y exponent of the temperature ratio - z exponent of the temperature ratiot _w/t _s - Nusselt-Number - Prandtl-Number - Reynolds-Number - Stanton-Number - _B m boundary layer thickness - kg/ms dynamic viscosity - W/°C m heat conductivity including the influence of turbulence - =t _w–t _s°C temperature difference - m²/s kinematic viscosity - kg/m³ density - N/m² shear stress Indices B boundary layer - F film - S steam - SS stainless steel - T turbulent core - W wall - a outside - h hydraulic - i inside - x based on reference temperaturet _x - 0 based on zero shear surface - 1 based on inner subchannel of an annulus - 2 based on outer subchannel of an annulus - - (over a symbol): average - * (over a symbol): transformed     

Rheology and morphology of starch/synthetic polymer blends

Corn starch and maleic anhydride functionalized synthetic polymers were melt blended in a Haake twin-screw extruder. The amount of starch in the blends was 60 and 70% by weight. The synthetic polymer used was either styrene maleic anhydride (SMA) or ethylene propylene maleic anhydride copolymer (EPMA). The blends did not exhibit normal thermoplastic behavior; and hence, rheological data was obtained by extrusion feeding the material through a slit die or cylindrical tube viscometer. The starch/SMA blends were extruded through a slit viscometer with a 45% half entry angle, while the starch/EPMA blends were extruded through a cylindrical tube viscometer with a half entry angle of 37.5°. For the blends, data could be obtained at low to moderate shear rates (10< _app<200s^–1). At higher shear rates, blends exhibited slip and/or degradation of starch. The viscosity of the blends exhibited shear-thinning properties. Regrinding and re-extruding through the viscometer a second time showed a significant reduction in shear viscosity for starch/SMA blends. Gel permeation chromatography data indicated that starch macromolecules degraded upon successive extrusion. Extensional viscosity, as estimated from entrance pressure drop method for starch/EPMA blends showed stretch thinning properties. Regrinding and re-extruding showed that the samples were more sensitive to changes in extensional viscosity as observed from the Trouton ratio versus extension rate plot. Optical microscopy showed the presence of starch granules after melt blending, the size of which was related to the torque (or stress) generated during extrusion. The higher the torque, the smaller the size of the starch granules. Successive extrusion runs reduced the number of unmelted granules.Nomenclature A,B Constants associated with power law fluids (Pa s^{m or n}) - e Entrance correction - H Height of slit die (m) - m, n Flow behavior index in shear and extension flow respectively - P _s Shear component of the entrance pressure drop (Pa) - P _e Extensional component of the entrance pressure drop (Pa) - Q volumetric flow rate (m³S^–1) - R _o radius of barrel exit (m) - R ₁ radius of cylindrical die (m) - T _r Trouton ratio - w width of slit die (m) - pressure gradient (Pam^–1) - half die entry angle - P _en Entrance Pressure Drop (Pa) - apparent extension rate (s^–1) - apparent shear rate (s^–1) - _w wall shear stress (Pa) - first normal stress difference in uniaxial extension (Pa)     

Extensional flow of polymer solutions through porous media

The flow behaviour of various polymer solutions of non-hydrolyzed polyacrylamide, hydrolyzed polyacrylamide, polyox and Xanthan was investigated in a plexiglass column having a succession of enlargements and constrictions, and compared with the flow behaviour and mechanical degradation of a solution of non-hydrolyzed polyacrylamide in a packed column of non-consolidated sand. The flow behaviour of this solution was found to be very similar in both the sand pack and plexiglass pore.Apart from the Xanthan solution, all other polymer solutions showed a viscoelastic behaviour in the plexiglass pore. The onset of viscoelastic behaviour, which has previously been defined using the shear rate ( ), stretch rate ( _s ) and Ellis number (E ₁), could be more precisely evaluated using a modified stretch rate (S _G). The pressure losses across the plexiglass pore for different polymer solutions of the same type were found to follow a unique curve provided the suggested group (S _G) was used, a situation which was not achieved with the other rheological parameters.The multipass mechanical degradation of the non-hydrolized polyacrylamide was tested through the sand pack against the suggested group (S _G) and Maerker's group (M _a). It was found that the loss of the solution viscoelasticity due to multipass mechanical degradation was better represented usingS _G thanM _a. A cross-sectional area (cm²) - C ^* critical concentration of polymer (ppm) - d plexiglass pore enlargement diameter - D average sand grain diameter (cm) - e equivalent width for the plexiglass pore - E ₁ Ellis number (a Deborah number) - F _R resistance factor - F _Ri resistance factor at the first pass - h height of the flow path of the plexiglass pore - K power-law constant - K _h,K _w effective permeability to hydrocarbon and water, respectively (10^–8 cm²) - M _a Maerker's group for a given porosity (s^–1) - M _ai value ofM _a at the first pass - N _D Deborah number - n power-law index - Q flow rate (cm³/s) - R capillary radius (cm) - R _g radius of gyration - S _G suggested group of rheological parameters representing a modified maximum stretch rate (s^–1) - S _Gi value ofS _G at the first pass - T _R,t characteristic time for the fluid (s) - t _s residence time (s) - V ₀ superficial velocity (cm/s) - V mean velocity of flow through a porous medium (cm/s) - average axial velocity in the enlargement section of the plexiglass pore (cm/s) - V ₁,V ₂ maximum velocity at a plexiglass enlargement neck and centre - [] intrincis viscosity - viscosity (mPa s) - _r relative viscosity (ratio of the viscosity of the polymer solution to that of the solvent) - shear rate (s^–1) - _s stretch rate (s^–1) - characteristic time for the polymer solution (s)     

Rheological properties of reactive polymer melts.

A new method for describing the rheological properties of reactive polymer melts, which was presented in an earlier paper, is developed in more detail. In particular, a detailed derivation of the equation of a first-order rheometrical flow surface is given and a procedure for determining parameters and functions occurring in this equation is proposed. The experimental verification of the presented approach was carried out using our data for polyamide-6.Notation E Dimensionless reduced viscosity, eq. (34) - E ₀ Newtonian asymptote of the function (36) - E power-law asymptote of the function (36) - E _{= 1} the value ofE at = 1 - k degradation reaction rate constant, s^–1 - k ₁ rate constant of function (t), eq. (26), s^–1 - k ₂ rate constant of function (t), eq. (29), s^–1 - K(t) residence-time-dependent consistency factor, eq. (22) - M _w weight-average molecular weight - M _x x-th moment of the molecular weight distribution - R gas constant - S _x M _x /M _w - t residence time in molten state, s - t _j thej-th value oft, s - T temperature, K - % MathType!MTEF!2!1!+-% feaafiart1ev1aaatCvAUfeBSjuyZL2yd9gzLbvyNv2CaerbuLwBLn% hiov2DGi1BTfMBaeXatLxBI9gBaerbd9wDYLwzYbItLDharqqtubsr% 4rNCHbGeaGqiVu0Je9sqqrpepC0xbbL8F4rqqrFfpeea0xd9vqpe0x% c9q8qqaqFn0dXdir-xcvk9pIe9q8qqaq-xir-f0-yqaqVeLsFr0-vr% 0-vr0db8meaabaqaciGacaGaaeqabaWaaeaaeaaakeaaieGaceWFZo% Gbaiaaaaa!3B4E!\[\dot \gamma \] shear rate, s^–1 - _i thei-th value of , s^–1 - _r =1 the value of at = 1, s^–1 - ^* reduced shear rate, eq. (44), s^–1 - dimensionless reduced shear rate, eq. (35) - viscosity, Pa · s - shear-rate and residence-time dependent viscosity, Pa · s - zero-shear-rate degradation curve - degradation curve at - _{t0 (t)} zero-residence-time flow curve - Newtonian asymptote of the RFS - instantaneous flow curve - power-law asymptote of the RFS - _0,0 zero-shear-rate and zero-residence-time viscosity, Pa · s - _E=1 value of viscosity atE=1, Pa · s - ^* reduced viscosity, eq. (43), Pa · s - zero-residence-time rheological time constant, s - density, kg/m³ - (t),(t) residence time functions     

Instantaneous visualization of surface flows

An experimental system is described for visualizing the surface flow of a wing, using an oil smoke tracer technique. The method leads to the determination of the instantaneous velocity direction at the output of surface injectors. A preliminary investigation is made on a flat plate to optimize the conditions of oil smoke injection. Then, the visualization is performed on the upperside of a sweptback wing in the vicinity of the reattachment of the vortex flow. This visualization technique can be applied to other types of wall flows — separated or not — around various bodies.List of symbols b wing span - c _n normal (to leading edge) chord - c _r streamwise (or root) chord - d diameter of the injectors - distance from the apex along the leading edge - relative distance from the apex along the leading edge ( = /C _d) - sweep angle - e injector geometric parameter (e = d/l) - angle of attack - K injection parameter - l length of the injectors - v kinematic viscosity - P _t, P_s total and static pressure of the flow - P _inj injection pressure - P _r reduced pressure (P _r = (P_inj – P_t)/(P_t – P_s)) - Re flow Reynolds number (Re = V ·c _n/v) - Re _i injector Reynolds number (Re = V ·d/v) - s curvilinear distance along c _d - s relative curvilinear distance along c _d(s = s/c _d) - V infinite upstream flow velocity     

Influence of cavitation on blade characteristics

An experimental study of flow around a blade with a modified NACA 4418 profile was conducted in a water tunnel that also enables control of the cavitation conditions within it. Pressure, lift force, drag force and pitching moment acting on the blade were measured for different blade angles and cavitation numbers, respectively. Relationships between these parameters were elaborated and some of them are presented here in dimensionless form. The analysis of results confirmed that cavitation changes the pressure distribution significantly. As a consequence, lift force and pitching moment are reduced, and the drag force is increased. When the cavitation cloud covers one side of the blade and the flow becomes more and more vaporous, the drag force also begins to decrease. The cavity length is increased by increasing the blade angle and by decreasing thé cavitation number.List of symbols A (m²) blade area,B ·L - B (m) blade width - C _D (–) drag coefficient,F _D /(p _d ·A) - C _L (–) lift coefficient,F _L /(P _d ·A) - C _M (–) pitching moment coefficient,M/(P _d ·A ·L) - C _p (–) pressure coefficient, (p-p _r )/p _d - F (N) force - L (m) blade length - M (Nm) pitching moment - p (Pa) local pressure on blade surface - p _d (Pa) dynamic pressure, ·V ²/2 - p _r (Pa) reference wall pressure at blade nose position if there would be no blade in the tunnel - p _v (Pa) vapor pressure - p ₁ (Pa) wall pressure 350 mm in front of thé blade axis - Re (–) Reynolds number,V ·L/v - V (m/s) mean velocity of flow in the tunnel - x (m) Cartesian coordinate along thé blade profile cord - x _c (m) cavity length,x-coordinate of cavity end - (°) blade angle - v (m²/s²) kinematic viscosity - (kg/m³) fluid density - (–) cavitation number, (p _r –p _v )/p _d - (°) angle of tangent to thé blade profile contour     

The shear viscosity dependence on concentration,molecular weight,and shear rate of polystyrene solutions

The solution viscosity of narrow molecular weight distribution polystyrene samples dissolved in toluene and trans-decalin was investigated. The effect of polymer concentration, molecular weight and shear rate on viscosity was determined. The molecular weights lay between 5 10⁴ and 24 10⁶ and the concentrations covered a range of values below and above the critical valuec ^*, at which the macromolecular coils begin to overlap. Flow curves were generated for the solutions studied by plotting log versus log . Different molecular weights were found to have the same viscosity in the non-Newtonian region of the flow curves and follow a straight line with a slope of – 0.83. A plot of log ₀ versus logM _w for 3 wt-% polystyrene in toluene showed a slope of approximately 3.4 in the high molecular weight regime. Increasing the shear rate resulted in a viscosity that was independent of molecular weight. The sloped (log)/d (logM _w ) was found to be zero for molecular weights at which the corresponding viscosities lay on the straight line in the power-law region.On the basis of a relation between _sp and the dimensionless productc · [], simple three-term equations were developed for polystyrene in toluene andt-decalin to correlate the zero-shear viscosity with the concentration and molecular weight. These are valid over a wide concentration range, but they are restricted to molar masses greater than approximately 20000. In the limit of high molecular weights the exponent ofM _w in the dominant term in the equations for both solvents is close to the value 3.4. That is, the correlation between _sp andc · [] results in a sloped(log _sp)/d(logc · []) of approximately 3.4/a at high values ofc · [] wherea is the Mark-Houwink constant. This slope of 3.4/a is also the power ofc in the plot of ₀ versusc at high concentrations. a Mark-Houwink constant - B ₁,B ₂,B _n constants - c concentration (g · cm^–3) - c ^* critical concentration (g · cm^–3) - K, K constants - K _H Huggins constant - M molecular weight - M _c critical molecular weight - M _n number-average molecular weight - M _w weight-average molecular weight - n sloped(log _sp)/d (logc · []) at highc · [] - PS polystyrene - T temperature (K) - shear rate (s^–1) - critical shear rate (s^–1) - viscosity (Pa · s) - ₀ zero-shear viscosity (Pa · s) - _s solvent viscosity (Pa · s) - _sp specific viscosity - [] intrinsic viscosity (cm³ · g^–1) - dynamic viscosity (Pa · s) - | ^*| complex dynamic viscosity (Pa · s) - angular frequency (rad/s) - density of polymer solution (g · cm^–3) - ₁₂ shear stress (Pa) Dedicated to Prof. Dr. J. Schurz on the occasion of his 60^th birthday.Excerpt from the dissertation of Reinhard Kniewske: Bedeutung der molekularen Parameter von Polymeren auf die viskoelastischen Eigenschaften in wäßrigen und nichtwäßrigen Medien, Technische Universität Braunschweig 1983.     

Mathematical model of a single-screw plasticating extruder

A five zone mathematical model of a plasticating extruder is presented. Its application in the design of new and improvement of existing extruders is briefly described. The model is based on theories proposed by Darnell and Mol, Tadmor, Broyer, McKelvey, Klein, Schneider, Fenner, Poon and Jankov. A comparison between experiments and theoretical calculations is included. E energy, W - f melt film thickness, m - f _k friction coefficient - h channel depth, m - l axial screw distance, m - k power-law parameter, °C^–1 - m ₀ power-law parameter, Pa s ⁿ - MI melt index, g/10 min - n power-law parameter - p pressure, Pa - S screw lead, m - t temperature, °C - t _c time, s - T temperature, K - v velocity, m s^–1 - X solid bed width, m - y rectangular coordinate (channel depth direction), m - Z 1/S (turn), m^–1 - shear rate, s^–1 - apparent viscosity, Pa s - feed angle, ° - density, kg m^–3 - shear stress, Pa - a solid - b barrel or bulk - d dissipated - f flight - m melt - s screw - t total - x width channel direction - z length channel direction