Authors  have  developed  a  nonlinear  pipe  element  for  MMG  Co.  Ltd  who  is  owner  of 
several  patents  pending  of  flow  meters  in  order  to  take  into  account  the  Coriolis  forces 
produced by  the flow  of fluids  in  vibrating pipes.  The aim  of the development  is  to improve 
the  construction  of existing  Coriforce  flow  metals  where  the measuring method  is  based 
on  the presence of Coriolis  forces  which  are  in  linear  relation  with  the mass  flow  in  pipes. 
The  paper discusses  the modelling  aspects  and  shows  industrial  example  as  well

Takács, Ferenc

Tóth, György

English

Periodica Polytechnica (Budapest University of Technology and Economics)

PERIODICA POLYTECHNICA SER. TRANSP. ENG. VOL. 22, NO. 3-4, PP. 207-223 (1994) FLOW METER OPTIMIZATION BY USER DEFINED NONLINEAR FINITE ELEMENT Ferenc TAK.A.cs* and Gyorgy TOTH** * Department Transport Engineering lvfechanics Technical University of Budapest H-1521 Budapest, Hungary .* T &T Data Automation Ltd H-1116 Budapest, Hunyadi ?Vl. u. 32. Hungary Abstract Authors have developed a nonlinear pipe element for ~vfMG Co. Ltd who is owner of several patents pending of flow meters in order to take into account the Coriolis forces produced by the flow of fluids in vibrating pipes. The aim of the development is to improve the construction of existing Coriforce flow mete!'s where the measuring method is based on the presence of Coriolis forces which are in linear relation with the mass flow in pipes. The paper discusses the modelling aspects and shows industrial example as well. Keywords: nonlinear finite element, Coriolis force, flow meter. 1. Introduction Direct flow meters based on Coriolis forces \vere developed dynamically during the last decade. These are direct means that we do not have to measure density and volume flow separately and compute the mass flow indirectly. As the mass flow induces Coriolis forces, measurement of these forces lets us know the mass flow in a direct way. The main advantage of Coriolis flow meters is that in a large measuring range (turn-dmvn ratio 1/20) we can measure the mass flow with a 0.2% accuracy independently of the fluid consistency (physical state, viscosity, density). As we do not have moving parts we can achieve high reliability, stability and life expenditure. Therefore Coriolis flmv meters are going to replace indirect flow meters on more and more fields of use. Development of vibrating Coriolis flow meters has produced a big variety of pipe forms and vibrating methods. Most of solutions can be characterized by the following remarks: Every cross section of a pipe constrained at its extremities and excited with its eigenfrequency is moving on an arched path, therefore for every cross section there is a periodic angular velocity. Consequently, the Coriolis forces produced by mass flow in the pipe and the above mentioned angular velocity will be periodic of the same 208 F. TAK.4CS and GY. TOTH frequency which will excite periodic vibration of the pipe. This vibration is due only to mass fiow and it will be superposed on excited vibration. One can always find two points along the length of the pipe where the phase shift of the vibrations shows only the effects of Coriolis forces, depending on mass fiow. The phase shift of the two sinusoidal signals is the output of the fiow meters. The design of optimal fiow meters is to determine the optimal values of numerous parameters Ivhich are function of the others. \Ye mention only the most important parameters here: geometry of vibrating pipe material of pipe vibrating mode position and mass of sensors mounted on pipe - measuring range - error rate of signal processing (error of signal processing divided by the total errors of the fiow measurement) - damping and effect of vibration loss (mechanical coupling with the environment. balancing) effect of outer mechanical noise - effect of static and dynamic mechanical loads (pressure dependency, fatigue) pressure loss \Ye have two solutions for the fim\' meter optimal design: 1. Analysis of measurements of one or more prototypes. This process is highly interactive and requires sequent modification and measurement of prototypes. 2. State a numerical model of the physical problem and find optimal ·values of the model parameters. \1:\1G Co. Ltd has made Economic efficiency estimations for both methods. These estimations haw made \1::\IG Co. Ltd. purchase Systus finite element soft,vare as a tooi for the second method. Tv,'Q major argu-ments have infiuenced this decision: First: Construction of large prototypes is time consuming and expensive. Second: ,.i.s eyen largest finite element sofn\'ares do 'lOt include Cori-olis pipe elements authors have developed a ne·w elemellt IQ "le linked ":ith S~'stus shared librar:,. UUf paper Sh0'\'5 the most imponant steps of cie-,:l1ent derivation and compares Its beha\'iour \vith theoretical test and in-dustrial measurements. FLOW .IfETER OPTIMIZATIOi, Nomenclature A. cross section [B] damping matrix E Young modulus G shear modulus I area moment of inertia k Timoshenko shear coefficient [K] stiffness matrix L length of pipe [M] mass matrix N shape function vector q nodal variable vector S Hamilton functional t time T kinetic energy T integration time step 6.t time shift of zero crossing U potential energy Vo mean velocity of fluid x distance along pipe axis ex shear strain e rotation due to flexural deformation w angular velocity cp circumferential angle Subscripts f fluid quantities p pipe quantities Superscripts (e) elemental quantities T transpose 2. Derivation of CorioUs Pipe Element 209 The Coriolis forces effect only the transverse vibration of the fluid--conveying pipe therefore we show only the derivation of finite element for-mulation of these equations of motion by Hamilton's principle. For Hamil-ton's principle we have the functional t S = J (T - U)dt . o 210 F. TAK.4CS and GY. TOTH VVe introduce 8 as a rotation due to flexural deformations and a as shear strain. VVe assume that the slope of the pipe neutral axis oulox can be written in the form OU -=8+0'. ox This is a usual form for Timoshenko beams. VVith the above assumptions the potential energy of the pipe is the following - 1 ;L [ (08) 2 ?] Up = "2 Elp ox + kGA_pO'- dx. o Similarly the kinetic energy can be expressed in terms of transitional and rotational inertia The energy contribution of the fluid ,,,as formulated from Timoshenko per-spective. VVe neglected the effect of internal pressure on potential energy and assume that the fluid contributes only kinetic energy. This contribu-tion was formulated as follows Transverse fluid velocity OU f lot is related to the transverse velocity of the pipe oulot, through the- material derivative. The material derivative relates the Eulerian description of the fluid to the Lagrangian description of the pipe by the follmving: OUj ou ou --= vo-+-ot ox at The axial fluid velocity Vo was assumed to be independent of x and to be constant across the cross section of the pipe. In fact the fluid is handled as a solid travelling through the pipe at constant velocity. Similar assumptions ,vere made by other researchers [1 J. If \ye use the finite element technics for the formulation of the \veak from of the equations of motion derivated from the above assumptions with proper shape functions [2J we can ,Hite the displacement as follows FLOW ,lfETER OPTlMIZATJOX 211 as functions of time dependent nodal q values. The integral expressions for the element matrices are as fo11O\vs Taking the variation of S equal to zero we shall have the expression t ( , d2 d \ 8S = 8qT J [M] dti + [B] d~ + [K]q) dt = 0 ' o ' As 8q can be arbitrary \ve get the matrix equation _ d2 q dq I _ [M] dt2 + [B]di T [K]q - 0 . In the absence of structural of material damping, the damping matrix con-tains only gyroscopic coupling terms produced by Coriolis effects. The stiffness matrix includes terms originating from the bending and shear en-ergy, and the centripetal acceleration of the fluid. The mass matrix is com-posed of terms arising from the transverse and rotary inertia of both the fluid and pipe. The final matrix equation has the general form of a com-plex eigenvalue problem or complex differential equation, in this case there is an external excitation. The contribution of Coriolis damping is an anti-symmetric matrix. Inclusion of the Coriolis terms in the damping matrix differentiates this ,vork from that of previous researchers [3]. The solution of the above equations cannot be done \vith the standard Systus algorithms. \Ve have to use the nonsymmetrical Gauss algorithm for solving this prob-lem. As the presence of velocity dependent Coriolis damping makes our problem nonlinear \ve had to deyelop a subroutine for our user defined pipe element. This FORTRA:'-i subroutine v,'as linked \vith Systus shared ele-ment library and we have got a new element [4]. \Vithout Coriolis coupling (zero fluid velocity) this element behaves exactly like a standard type Sys-tus beam element. In case \ye define cross section area, area moment of in-ertia, density and mean velocity of the fluid in ~lATERIAL PROPERTIES 212 F. TAKi cs and CY. TOTH and use the DAMPING label in transient nonlinear, our element will take into account the [B] matrix programmed in the above mentioned subrou-tine and modify [M] and [K] matrices with Vo dependent terms. This new element was tested in several ways. In the foHo·wing we present a simple test and an industrial example as well. 3. Test Example For testing the reliability of our new element ,ye have defined a simple but demonstrative example where we can theoretically compute the effect of the Coriolis coupling. This example was a ring of pipe elements in x - y plane. All the el-ements were connected by rigid massless beams with the center point of the ring. We have allowed the rotation of the ring around an axis perpen-dicular to its plane and fixed against; displacement and other rotations in the center of the ring. \Ve have applied a constant 1 Nmm z torque on the axis which accelerated the model with a constant angular acceleration. The data of the model are the following: Ring diameter = 1000 mm, PI = 10-6 kgmm-3 , AI = 12.57 mm2, Internal pipe diameter = 4 mm, External pipe diameter = 5 mm. As we could compute the inertial and mass quantities related to the model it was easy to compute w in the 10th time step of integration when assuming a rigid body like motion around the rotational axis. This value was 1.0488 10-3 5-1 . Regarding the construction of the model we could assume that the model is dynamically balanced so we had not any reaction on the axis due to unbalance of inertia and mass. Therefore, if we regard the reactions on the axis we shall get exactly the sum of Coriolis forces acting on pipe elements. We have defined the flow direction of the fluid different on both halves of the model by inverting element axis y. Therefore, the sum of the Coriolis force has to be a -x direction vector. This vector can be computed by the following integral F. ~ -4pj AjRwvu J cO'<P d<p . -"2 If we take a Vo = 1000 mm/ s value thisforce will be Fx = 1.13 mN. We have computed Fx for fluid velocity 0,1, ... ,10 m/so The results are presented FLOW .IfETER OPTIMIZATIOS 213 in Fig. 1 and Fig. 2. vVe have used a T = 1 ms time step in the transient nonlinear method. The computed errors were found minor than 2 % when comparing the results to theoretical values. 4. Industrial Example The so-called B type Coriforce flow meter is a standard product of MMG Co. Ltd. \Ve have selected one geometry configuration from the exist-ing product scale in order to compare its behaviour to model results. The model was simplified <":omparing to the real structure because we have ne-glected the flexibility of parts ,,,here Coriolis forces ,,,ere not foreseen. The geometry is presented in Fig. 3. At the nodes marked by 13, 14, 15 and 16 we have modelled sensors like lumped masses. The model was fixed in four points. The flow direction is equal in the lmver and upper parts. All the elements are of the new type. In the reality for the measurement - by ex-citing forces the structure is maintained in a mode shape which is charac-terized by Fig. 4 and is called butterfly mode. Practically, during the mea-surement it is enough to compare the z direction displacement functions of node pairs 13-15 or 14-16. The time shift between the zero crossings of these functions is a linear function of the fluid flow. The tube external and internal diameters were 31.75 and 29.67 mm. respectively. The ma-terial of the pipe and the eigenfrequency of the butterfly mode ·were steel and 96.02 Hz. In order to simulate the measuring situation ,vithout intro-ducing the real controlling system into Systus we had to produce an ini-tial state where the model shape and reactions were equal to the normal-ized butterfly mode. These initial conditions ,vere computed by all equiv-alent model constructed from standard type pipe elements. From these initial conditions v:;ith transient nonlinear method we have computed time shift of first zero crossings between the above mentioned node pairs z di-rectional displacement functions. The selected fluid velocity yalues were 0,1,2, ... ,10 m/so vVe have used several integration time steps. The val-ues computed with 5*10- 5 s are to compared to measured ones in Table 1. The errors of the computed values are the function of the selected time step. Fig. 5 illustrates the effect of T value selected. For Vo = 10 m/s Fig. 6 shmvs a 1/4 period of computed displ~cements at 13, 14, 15 and 16 nodes. Fig. 7 is the appropriat€ zoom of Fig. 6 for time shift measurement. Though the butterfly mode frequency of our model was exactly equal to the measured one we did not own measured values concerning the mode shape. Therefore we regard the computed errors very small. \Ve made several geometrical simplifications in the model which might effect mainly 214 SVSTUS HP7-232-v SVSTUS HP7-232-v Lx F. TAh".4CS and GY. TOTH CORIFORCE TEST C=O MIS CZ=1 NMM MOMENT CORIFORCE TEST C=1 MiS CZ=1 NMM MOMENT CAND_T DATA_AUTOMATION REAC 123 CARD 10 1< .887E-07 ft< .176E-06 m < .264E-06 B < .352E-06 < .440E-06 ! < .533E-06 I T_AND_T DATA_AUTOMATION REAC 123 CARD 10 1< .185 B < .370 g< .555 B < .740 < .925 s< 1.12 Fig. 1. Reaction forces for 0 and 1 m/s fluid velocities SVSTUS HP7-232-v Lx SVSTUS HP7-232-v Lx FLOW .lfETER OPTIMIZATIOt: COAIFOACE TEST C=2 M/S CZ=I NMM MOMENT COAIFORCE TEST C=3 M/S CZ.I NMM MOMENT 215 T_AND_T DATA_AUTOMATION AEAC 123 CARD 10 g< .372 1< .743 Q,1.I2 B < 1.49 < 1.66 0< 2.25 T_AND_T DATA_AUTOMATION AEAC 123 CAAD 10 1< .561 D< 1.12 D < 1.68 m < 2.25 < 2.61 B < 3.40 Fig. 2. Reaction forces for 2 and 3 m/s fluid velocities 216 F. TilK.4CS and GY. TOTH FLOW METER OPTIMIZATION Fig . . f. Butterfly mode of 96.02 Ez eigenfrequency Table 1 .Measured and computed time shifts Velocity of fluid Time shift*10-6 Time shift"1O-6 defo:-m1ltTlpi 87$ [m/s] measured seconds computed seconds 6.73 6.75 2 13.5 13.1 3 20.2 18.9 4 26.9 25.9 5 33.6 32.4 6 40.4 39.5 7 47.1 45.4 8 53.8 51.9 9 60.5 59 .. 5 10 67.3 66.5 217 the error magnitude. Probably with the use of a butterfly mode shape tuned with measured one we could achieve even higher accuracy. 218 8 8 Cl ~ Cl 11> '" § ~ . TOTH F. TAK.4 cs and G)  8 8 Cl Cl '" N 8 q 0-'" C;('I") M V C'l If) ('t") (,0 ("') ('I") - 0... - 0..... - 0.. - 0... LI> 0... CIl en CO (f) (f) zC;zC;zC;zC;zC; { I y .' VI ! 1/ ! I J FLOW METER OPTIMIZATION ~ III ".'" '\ j / \ > .\ Ij! / I o , ~ , '\ '~\\ \ \ ~ , -\ \ § ... ," \ UJ 1: ;:: ~ , UJ 0 ~ ~ [(j ~ ~ ~ ~ ~ \; \ ~ 219 N '0 DISP .4000E-Dl . .--/---/-'" /' 2000E-O .-~ -....~,-~ -.... , >"" '"', --"~~'" " ~ --", . 3725E-OO, ........... /~ . ./ /' -2000E-, ", , " " .~ ~ -.4000E-Fig. 7. Zoom of Vig. fi for visua.lizillg time shifL /""" ", " --, '-.. -~ 2 -02 N 13 DISP 3 N 14 DISP 3 N 15 DISP 3 N 16 DISP 3 N 5 DISP 3 TIME ""' ""' 0 "' ~ :>: :>-., () 'n o " "-Cl :<: '-l 0, ~ FLOW METER OPTIMIZATION 221 5. Conclusions Flow meter design is a difficult process. \Ve do not have the knowledge how structural modifications of the flow meter will effect the resolution and sensitivity characteristics and how they influence the accuracy of our measurements. In order to reduce the number of expensive prototypes a reliable computer model has to be used. Generally structural optimization is made by finite element softwares. Unfortunately even the world leader softwares cannot handle CorioEs forces. The fact that SYSTUS has a shared element library has made possible that the authors could develop a ne'w type element for MMG Co. Ltd whose behaviour was proved by measurements and theoretical test. There is a reason for standardize this new element in Systus because the presence of Coriolis forces is important for other fluid-conveying problems as well. References 1. HUA;-;G, C. C. (1974): Vibrations of Pipes Containing Flov,7ing Fluids According to Timoshenko Theory. Journal of Applied Mechanics, September, pp. 814-817. 2. WEAVER, W. Jr. JOHNSON, P. R. (1987): Structural Dynamics by Finite Elements. Prentice-Hall, 1987. 3. KOHL!, A. K. - NAKRA, B. C. (1984): Vibration Analysis of Straight and Curve Tubes Conveying Fluid by Means of a Straight Beam Finite Element. Journal of Sound and Vibration, VoL 93, No. 2, pp. 307-311. 4. Systus User Manual (1993): Transient Nonlinear. Framsoft+CSI (Group Framatome), August. Appendix Shape Functions and Elemental Matrix Formulations \'Ye assume the transversal displacement of our beam element has a cubic distribution over the element length, For the expression for the rotation due to flexural deformation over the element length we had the following distribution which assures constant shear strain, 222 F. TA.KACS and Cl'. TOTH Here 9 is a constant determined from the material and geometry of the pipe Elp g=--. kGAp The coefficients ao to a3 can be determined from transversal displacement and rotation at the ends x = 0 and x = L. The lower triangles of elemental mass, stiffness and damping matrices are the following: (ppA.p + P f A. f )L(1680g2 + 294gL 2 + 13L 4) I m11 I 35(12g + V)2 I ((JpIp +pf l f )6L3 T 5(12g + V) (ppA.p + PfA.f)L2(1260l + 231gL 2 + llL4) m21 = - 210(12g + V)2 (ppIp + PfIf)L2 (60g - L2) 10(12g + V)2 (ppAp + PfAf)L 3 (126g 2 + 21gL2 + L4) 7n.·'2 = + ~ 105(12g + V)2 I (pplp + p flf )2L(360g2 + 15gL2 + L4) T 15(12g + V)2 (ppA.p + PfAf)3L(560g 2 + 84gL 2 + 3["!) m31 70(12g + V)2 (pplp + pf I f )6L3 5(12g + gV)2 _ (ppAp + PfA.f)L2(2520g 2 + 378gL2 + 13L4) m31 -- ~~--~~--------~~~--------- 420(12g + V)2 (pplp + PfIf )L2(60g - L2) 10(12g + V)2 (ppAp + PfAf)L 3(158g2 + 28gL2 + L4) m.J') = - ..L .- 140(12g + V)2 I ? 2 J (pplp + Pflf )L(120g- - 60gL - L') + 30(12g + V)2 771 33 = 77111, m·n = 7n.32 , k - 12Elp _ 6pfAf v5 (120 2+20gL2..LL4). 11 - V + 12Lg 5L(V + 12g)2 9 I. FLOV/ .HETER OPTl.\f/ZATIOf'; 2 4 6Elp PfA.fvoL k21 = £2 + 12g - 10(£2 + 12g)2 4Elp(L2 + 3g) _ 2LpfA.fv5 (gOl + 15gL2 + L4) , k22 = £3 + 12Lg 15(£2 + 12g)2 k31 = -k11 , k32 = -k21 , k33 = k11 . k41 = k21 , k43 = -7;:21 . k44 = k22 . PfA.fl'oL(10g + L2) b21 = .S(12g + £2) b31 = -PfA.fl'O , b32 = b21 . b41 = -b21 , P fA. f voL4 b42 = 360g + 30£2 ' b43 = b21 . 223 

FLOW METER OPTIMIZATION BY USER DEFINED NONLINEAR FINITE ELEMENT

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FLOW METER OPTIMIZATION BY USER DEFINED NONLINEAR FINITE ELEMENT

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