Two-dimensional CFD modelling of gasdecompression

University of Wollongong Research Online Faculty of Engineering and Information Sciences - Papers: Part A Faculty of Engineering and Information Sciences 2013 Two-dimensional CFD modelling of gasdecompression behaviour Alhoush Elshahomi University of Wollongong, aes738@uowmail.edu.au Cheng Lu University of Wollongong, chenglu@uow.edu.au Guillaume Michal University of Wollongong, gmichal@uow.edu.au Xiong Liu University of Wollongong, xiong@uow.edu.au Ajit R. Godbole University of Wollongong, agodbole@uow.edu.au See next page for additional authors Publication Details Elshahomi, A., Lu, C., Michal, G., Liu, X., Godbole, A., Botros, K. K., Venton, P. & Colvin, P. (2013). Two-dimensional CFD modelling of gas-decompression behaviour. The 6th International Pipeline Technology Conference (pp. S18-02-1 - S18-02-23). United States: Clarion Technical Conferences. Research Online is the open access institutional repository for the University of Wollongong. For further information contact the UOW Library: research-pubs@uow.edu.au

Two-dimensional CFD modelling of gas-decompression behaviour Abstract ONE OF THE key issues in the design and construction of any gas pipeline is the prevention of material fracture. Since the gas is generally transported under high operating pressures, it must be ensured that the gas pipeline is sufficiently tough to arrest the propagation of any potential fracture. For several decades, control of gas pipeline fracture propagation has been under scrutiny due to economic considerations and ecological and safety hazards related to pressurised pipe failure. The Battelle Two Curve approach has been widely used to determine the minimum material arrest toughness by comparing the gas decompression wave velocity with the fracture velocity, both as functions of the local gas pressure. Sufficient knowledge of the gas decompression behavior following the rupture is therefore crucial in determining running fracture arrest toughness levels. The decompression behaviour is influenced by the operating conditions, fluid composition and the material properties of the pipeline itself. This paper describes a two-dimensional decompression model developed using the commercial Computational Fluid Dynamics (CFD) software ANSYS Fluent. The GERG-2008 Equation of State has been implemented into this model to simulate the rapid decompression of common natural gas mixtures. The evolution of the decompression wave speed and phase changes under arbitrary initial conditions is reported. Comparison with experimental results obtained from shock tube tests showed good agreement between simulation and experiment. Disciplines Engineering Science and Technology Studies Publication Details Elshahomi, A., Lu, C., Michal, G., Liu, X., Godbole, A., Botros, K. K., Venton, P. & Colvin, P. (2013). Twodimensional CFD modelling of gas-decompression behaviour. The 6th International Pipeline Technology Conference (pp. S18-02-1 - S18-02-23). United States: Clarion Technical Conferences. Authors Alhoush Elshahomi, Cheng Lu, Guillaume Michal, Xiong Liu, Ajit R. Godbole, K K. Botros, Phillip Venton, and Philip Colvin This conference paper is available at Research Online: http://ro.uow.edu.au/eispapers/3342

Two-dimensional CFD Modelling of Gas Decompression Behaviour A. Elshahomi 1, C. Lu 1, G. Michal 1, X. Liu 1, A. Godbole 1, K.K. Botros 2, P. Venton 3, P. Colvin 4 1 University of Wollongong, Wollongong, NSW, Australia 2 Nova Research and Technology Centre, Calgary, AB, Canada 3 Venton and Associates Pty Ltd, Bundanoon, NSW, Australia 4 Jemena Ltd, Sydney NSW, Australia ABSTRACT One of the key issues in the design and construction of any gas pipeline is the prevention of material fracture. Since the gas is generally transported under high operating pressures, it must be ensured that the gas pipeline is sufficiently tough to arrest the propagation of any potential fracture. For several decades, control of gas pipeline fracture propagation has been under scrutiny due to economic considerations and ecological and safety hazards related to pressurised pipe failure. The Battelle Two Curve approach has been widely used to determine the minimum material arrest toughness by comparing the gas decompression wave velocity with the fracture velocity, both as functions of the local gas pressure. Sufficient knowledge of the gas decompression behavior following the rupture is therefore crucial in determining running fracture arrest toughness levels. The decompression behaviour is influenced by the operating conditions, fluid composition and the material properties of the pipeline itself. This paper describes a two-dimensional decompression model developed using the commercial Computational Fluid Dynamics (CFD) software ANSYS Fluent. The GERG-2008 Equation of State has been implemented into this model to simulate the rapid decompression of common natural gas mixtures. The evolution of the decompression wave speed and phase changes under arbitrary initial conditions is reported. Comparison with experimental results obtained from shock tube tests showed good agreement between simulation and experiment. INTRODUCTION A running ductile facture is a particularly disastrous kind of pipeline failure. Such a failure includes the rapid axial cracking or tearing of the pipeline, sometimes running for long distances. This may result in considerable damage to the environment, loss of tens of millions of dollars and sometimes loss of life [1-2]. The Battelle Two Curve Method (BTCM) has been widely used to determine the minimum toughness required for ductile fracture

propagation arrest [3-4]. This method involves the superposition of two curves: the gas decompression wave speed characteristic and the ductile fracture propagation speed characteristic, each as a function of the local gas pressure [4-5]. According to this approach, the gas decompression wave velocity needs to be faster than the fracture propagation velocity of the pipe wall at any pressure above the arrest pressure to prevent long running fractures of gas pipelines [6]. Since the 1970s a number of full-scale burst tests for gas and liquid pipelines were conducted to investigate gas decompression behaviour [7-15]. However, full-scale burst tests are expensive and are economically feasible only for major projects. Phillips and Robinson reported measurements of decompression wave speed using an NPS 6 shock tube [16-17]. In recent years, Botros et al. [4, 18-22] have experimentally investigated the decompression wave speed using a small diameter NPS 2 shock tube. In both the shock tube test and fullscale burst test, the decompression wave speed can be determined from pressure-time traces measured by transducers mounted at different locations along the pipe length [18, 21]. MODELLING A number of models have been developed for predicting the gas decompression wave speed [16]. Many of the models assume a one-dimensional (along the pipe axis) and isentropic flow[23]. They also rely on the fundamental thermodynamic expansion of the gas at the crack point independent of the pipe diameter [4]. The Finite Difference Method (FDM) [24] and the Method of Characteristics (MOC) [25-26] are used in many models. One of the most widely used decompression model, implemented in GASDECOM [27], was developed at the Battelle Memorial Institute in the 1970s [16, 18, 22, 28]. GASDECOM uses the analytical solution of the propagation of an infinitesimal decompression front to determine the decompression wave speed without solving the fluid transport equations explicitly. Onedimensional, frictionless, isentropic and homogeneous fluid flow are assumed in the model. GASDECOM uses the Benedict-Webb-Rubin-Starling (BWRS) Equation of State (EOS) with modified constants to estimate the thermodynamic parameters during the isentropic decompression. GASDECOM has been validated against measured results of full-scale fracture propagation tests and shock tube tests [4]. It has been largely accepted as a benchmark model to predict the decompression curve of lean natural gas mixtures. GASDECOM calculates the gas decompression wave speed (W) as: (1)

where C is the speed of sound behind the decompression wave, and u is the outflow velocity behind the decompression wave. Pressure or density drops are used to calculate C and u iteratively. The outflow velocity u at any given pressure is the sum of the incremental velocities Δu determined from: where (2) (3) and is the density, and the subscript s indicates a value on the isentrope. Several models have followed the assumptions of GASDECOM for predicting the gas decompression wave speed [16]. The differences between these models are mostly the use of different EOS. Groves et al. [12] formulated a computer program to calculate the gas decompression velocity using an approach similar to shock tube theory. This model uses the Soave-Redlich-Kwong (SRK) EOS to determine the required thermodynamic properties. This model assumes that the pipe is of a large diameter and that the heat transfer within the boundary layer has a negligible effect on the gas decompression behaviour [12]. The DECAY model was established by Jones and Gough [29] to model single-phase decompression in a pipe undergoing fracture propagation. This model uses assumptions similar to GASDECOM, and the Peng-Robinson (PR) EOS [16]. The Advantica model developed in the late 1990s by Cleaver and Cumber [30] is also based on the theory of one-dimensional shock tube test. This model uses the cubic London Research Station (LRS) EOS, which is similar to the Redlich-Kwong-Soave (RKS) EOS. The results in the Advantica model are adjusted by a scaling factor applied to the speed of sound to match the initial conditions of sound speed predicted using the PR EOS [16]. Makino et al. [31] proposed a one dimensional decompression model at the University of Tokyo to simulate the decompression accompanied by phase change in natural gas pipelines. This model uses the theoretical approach of British Gas that has been implemented in the DECAY model along with the BWRS EOS. Qiu et al. [32] developed a one-dimensional decompression model called DECOMWAVE using four EOS(s): PR, SRK, BWRS-PR, and BWRS-SRK. For calculating the speed of sound in the two phase region, the united equation developed by [33] was adopted. Cosham et al. [28] established a one-dimensional, isentropic and homogeneous-equilibrium model named

DECOM to model the decompression behaviour of CO 2 pipelines. This model has assumptions identical to that used in GASDECOM but replaced the BWRS EOS by two EOS(s) Span and Wanger [34]and GERG-2004[35], which are built in the NIST Standard Reference Database 23 (REFPROP version 9.0)[36]. Simulation results generated by DECOM were validated against measurements on shock tube tests commissioned by National Grid [28]. More complex models for simulating transient flow in pipelines use the Computational Fluid Dynamics (CFD) approach. The non-isentropic flow condition can be assumed in those models to examine the effects of friction, heat transfer and pipe diameter on the decompression behaviour. Picard and Bishnoi [37-38] introduced a one-dimensional, nonisentropic model to investigate the decompression characteristics in ductile fracture propagation. The thermodynamic properties were determined using the PR EOS. The result shows that the non-isentropic assumption could be relaxed for pipe diameters larger than 508 mm [37-38]. The commercially available OLGA software [39] is a two-phase, transient computational pipeline simulators. The basic model of OLGA contains separate onedimensional mass conservation equations for the gas and continuous liquid phases. It can account for the presence of liquid droplets coupled with inter-phase mass transfer terms [19]. OLGA uses an implicit finite-difference discretisation method with a staggered mesh to numerically solve the conservation equations. This approach is suitable for both rapid and slow transient flows. OLGA often under-predicts the fast decompression wave speed of common hydrocarbon gas mixtures. This issue was linked to the assumption made to calculate the speed of sound [21]. While OLGA was validated for slow transient flows, rapid transients associated with full pipe rupture have not been tested [19]. PipeTech is another transient multi-phase, non-isentropic simulator based on the model of Mahgerefteh et al. [40-41]. PipeTech uses the MOC to solve the conservation equations by following the Mach-line characteristics inside the pipe. Numerical diffusion related to the finite difference approximation of partial derivatives is avoided by this method [25, 42]. PipeTech uses the cubic PR EOS [43] to calculate the thermo-physical properties of multi-component gas mixtures [25, 40-42, 44]. This model has been validated against a number of full bore rupture tests performed by Shell and BP on the Isle of Grain for hydrocarbon pipelines. Simulations using both the Finite Volume Method (FVM) and MOC have been found to produce results that agree remarkably well with some of the measured data [1].

In contrast to the FVM, the MOC needs significantly longer computation runtime and cannot predict non-equilibrium or heterogeneous flows [1, 44]. Jie et al. [44] developed a one-dimensional decompression model named CFDDECOM using the FVM technique based on the Arbitrary Lagrangian Eulerian (ALE) approach and the Homogeneous Equilibrium Method (HEM) [45-46] for simulating multi-phase transient flow following the rupture of pipelines conveying rich gas or pure carbon dioxide (CO 2 ). The Peng-Robinson-Stryjek-Vera EOS, in addition to the Peng-Robinson and Span and Wanger EOS[34], are implemented into this model [44]. CFDDECOM results were compared to the recently published shock tube tests for pure and multi-component CO 2 mixtures commissioned by National Grid [44]. Most of the existing decompression models are one- dimensional flow models. Their accuracy varies greatly with the EOS employed. The use of an EOS depends on the fluid region where the calculation of the thermodynamic properties is required. It also depends on the uncertainty ranges of the EOS in representing the experimental data [35, 47]. Botros et al. [13, 17] compared the predicted densities in the dense phase region by five different equations of State (GERG, AGA-8, BWRS, PR and RKS), with measured values for different hydrocarbon mixtures. They concluded that the GERG EOS outperforms all other equations in the region up to P = 30 MPa and T > -8 0 C. However, the GERG-2008 EOS has not been implemented in the CFD decompression models to date. The GERG-2008 EOS [35, 48] covers the gas phase, the liquid phase, the supercritical region, and vapour-liquid equilibrium states for natural gases and other mixtures consisting of up to 21 components: methane, nitrogen, carbon dioxide, ethane, propane, n-butane, isobutane, n-pentane, isopentane, n-hexane, n-heptane, n-octane, hydrogen, oxygen, carbon monoxide, water, helium, argon, n-nonane, n-decane, and hydrogen sulphide. The normal range of validity of GERG-2008 covers temperatures from 90 K to 450 K and pressures up to 35 MPa. The uncertainty in estimating the gas phase density and speed of sound is less than 0.1% for temperatures ranging from 270 K to 450 K at pressures below 35 MPa. In the liquid phase, the uncertainty in the density amounts for less than 0.1 to 0.5% for many binary and multi-component mixtures. The estimated uncertainty in liquid phase (isobaric) enthalpy differences is less than 1%, and can be as low as 0.5% for some mixtures. The vapour-liquid equilibrium is described with reasonable accuracy. Accurate vapour pressure data for binary and ternary mixtures consisting of the natural gas main components are reproduced by GERG-2008 to within their experimental uncertainty, which is approximately 1 to 3%.

As reported in our previous study [5], the GERG-2008 EOS was implemented in a onedimensional decompression model named EPDECOM developed at the University of Wollongong. The current study aims at developing a two-dimensional decompression model using the commercial CFD package ANSYS Fluent. The built-in EOS(s) in ANSYS-Fluent are not capable of predicting the fluid properties of natural gas mixtures accurately. GERG- 2008 EOS was successfully implemented in ANSYS-Fluent to calculate the thermo-physical behaviour of most common gas mixtures. This imparts flexibility to this model which can simulate the phase change of gas mixtures over a wide range of initial conditions. Since the model is two-dimensional, it is possible to examine effects of parameters such as friction, heat transfer and pipe diameter. CFD DECOMPRESSION MODEL The modelling of gas decompression behaviour requires solving the transient form of the flow governing equations using an accurate model of the fluid properties. The CFD package ANSYS-Fluent can simulate a wide range of compressible, laminar and turbulent, steady-state or transient flows of ideal or real fluids, in multi-dimensional geometries [49]. The decompression of rich and lean natural gas mixtures is not modelled accurately using an ideal gas model. Real gas models are better suited to the application. GERG 2008 is not available in Fluent by default. The equation of state must be implemented by using the User Defined Functions (UDF) available through the User- Defined-Real-Gas Model (UDRGM). The UDRGM implements a library of functions written by the end user in the C programming language. These functions are compiled and grouped in a shared library, which is later linked by ANSYS Fluent at runtime. The calculated fluid properties are passed to Fluent through the UDF as a single and homogeneous fluid. More details on using UDFs can be found in [50]. The procedure of how ANSYS Fluent solver works with the new UDF are illustrated in Figure 1. NUMERICAL SOLUTION In ANSYS-Fluent, the Finite-Volume technique is employed to discretise the governing equations of fluid flow [51]. The basic methodology involves dividing the computational domain into a number of sub-regions called control volumes or cells. The governing equations are integrated over all the control volumes of the computational domain

[51-52]. The resulting integral equations are discretised into a system of algebraic equations that is solved numerically [51]. In this model the coupled-implicit density-based solver was chosen [52]. This solver is designed for high-speed compressible flows and allows the implementation of user-defined real gas models. The governing flow equations of mass, momentum and energy conservation, supplemented by the auxiliary equations (e.g. EOS) are solved simultaneously while the turbulence equations are sequentially treated. The advective terms are treated using the second-order upwind scheme [53]. In the density-based solver, the momentum equations are used to obtain the velocity field, and the continuity equation is used to find the density field while the pressure field is determined from the EOS. The physical flow domain consists of the initially highly pressurised gas in a straight pipeline, which undergoes a full-bore opening. It is beneficial to take advantage of the axial symmetry to construct a two-dimensional computational domain, thereby reducing the computational runtime. The forms of the unsteady, two-dimensional governing differential equations of conservation of mass, momentum and energy required to be solved in this model are: Continuity: (4) where, x is the axial coordinate, r is the radial coordinate, is the axial velocity, is the radial velocity, and is the source term [52]. x-momentum: 2. (5) r-momentum: 2. 2. (6) where F is the external body force, is the swirl velocity and. (7)

ANSYS Fluent solves the general energy equation in the following form:... (8) where E is the fluid energy, k eff is the effective conductivity (= k + k t ), where is the turbulent thermal conductivity, h j is the enthalpy per unit mass of species j, and is the diffusion flux of species j. The first three terms on the right-hand side of Equation 8 represent the energy flow due to conduction, species diffusion, and viscous dissipation, respectively. In the present study there are no multiple species. Thus, the term may be considered similar to axial dispersion the effect which wanes as the flow velocity increases. includes enthalpy sources where heat transfer can be quantified [52]. The above four equations (4, 5, 6 and 8) involve five unknowns: velocity components v x, v r, pressure p, temperature T and density. For closure, the fifth equation is the GERG-2008 EOS (invoked using UDF). BOUNDARY CONDITIONS AND ASSUMPTIONS Based on the physical dimensions of the shock tube test facility described in [21], a horizontal pipe with 30 m length and 49.3 mm internal diameter is employed as the flow domain for the simulation. The length of the pipe was reduced from the original 172 m to 30 m to limit the computational runtime. The 30 m length was sufficient to ensure that the decompression wave would not be reflected off the closed far end in the simulated time of the decompression. The scored rupture disc in the shock tube test is modelled using pressureoutlet boundary condition at atmospheric pressure. A boundary condition of axial symmetry was imposed. The no-slip condition was specified at the wall boundaries. The physical flow domain and the computational domain are shown in Figure 2. The fluid, considered at rest, filled the pipe at the initial pressure and temperature conditions before a full bore opening was instantaneously created (time t = 0). Driven by the large pressure drop at the opening, the gas escapes from inside the pipe to the low pressure region. The decompression wave front immediately progresses away from the opening. Ahead of the decompression front, the fluid remains at rest and at the initial conditions. The local Mach number ranges from 0 at the decompression front to 1, corresponding to the choked condition, at the full bore opening end.

Because the flow generated after rupture is considered compressible, viscous and turbulent, the turbulence models available in ANSYS-Fluent need to be applied. The mesh grid was refined in regions of large flow gradients, i.e. near the outlet and along the boundary layer in the vicinity of the wall. The two layer modelling approach offered in ANSYS- Fluent was used in the latter region. This approach divides the flow domain into a viscosityaffected region (near the wall) and a fully-turbulent region (away from the wall). The Realizable k-ε turbulence model was adopted to model the fully-turbulent region while the near wall region was treated using the enhanced wall treatment function. More details about the formulations can be found in [52]. In ANSYS-Fluent, these two regions are defined using the turbulence Reynolds number Re: (9) where,, y,, and k are fluid density, the distance from the wall, dynamic viscosity and turbulent kinetic energy respectively. The flow is considered fully turbulent when Re > 200. The dimensionless wall distance y + for a wall tube-bounded flow is defined as: (10) where, is the distance to wall, and is the friction velocity. The dimensionless distance (theoretical) was used to find out the proper location and the size of the cell nearest to the wall. Since axial symmetry was assumed, the computational grid was generated over the 30 m length of the pipe and 0.025 m pipe diameter. At both wall boundaries 10 cells were generated to cover the boundary layer. The cell adjacent to the wall was set at 0.05 mm from the wall with a mesh-growth factor (from the wall) of 1.2. Following the 10 th cell in both radial and axial directions, the cells dimensions were kept constant at 1 mm up to the pipe axis and the outlet boundary respectively. Overall, 1,020,306 rectangular cells of quadratic mesh distribution were generated for the entire 2-dimensional axi-symmetric pipe. A detail of the mesh distribution of flow domain near wall boundaries is shown in Figure 3. SOLUTION STRATEGY

The implicit density-based solver was used. The default convergence criteria were assumed (limits on the residual values of continuity, x-, r-velocities, turbulent kinetic energy, turbulent dissipation energy and energy equation). A constant time step of 10-5 s was used to capture the transient flow features at the monitor point nearest to the full bore opening. Properties such as pressure, temperature, outflow velocity and speed of sound were monitored as a function of time traces to calculate the decompression wave speed. These were recorded at four different locations from the opening at the identical locations of the pressure transducers P3, P4, P6 and P7 in the shock tube tests [19]. The pipe wall was smooth and adiabatic. CALCULATION OF DECOMPRESSION WAVE SPEED Two different methods can be used to determine the decompression wave speed. The first method involves calculating the local decompression wave speed using equation (1). This can be done by monitoring the change in both the speed of sound and the outflow velocity against time. Afterwards, the outflow velocity is subtracted from the speed of sound for several pressures below the initial pressure. However, experimental tests such as shock tube test and/or full-scale burst test do not provide the local gas decompression wave speed. In these tests, the gas decompression wave speed w is calculated by determining the times at which a certain pressure level reaches several given locations (pressure transducer locations) on the pipe. By plotting these locations against time, the decompression wave speed is obtained by performing a linear regression of each isobar curves. The slope of each regression represents the average decompression wave speed for each isobar. Both methods are used in this model to calculate the gas decompression wave speed. CFD RESULTS AND VALIDATION The following paragraphs present the results of the two-dimensional CFD decompression model using the GERG-2008 EOS. Validation of this model was performed through comparison with the measured results of shock tube tests [19, 21]. The shock tube test was conducted at the TransCanada Pipeline Gas Dynamics Test Facility in Didsbury, Alberta Canada [21-22]. In this facility, the main test section of the shock tube had a total length of 172 meters. All spools were made from NPS2 x5.5 mm WT, seamless tube with an internal diameter of approximately 49.3mm. These spools were designed for a maximum of 22 MPa pressure, with a design factor of 0.8 and location factor of 0.625. The internal

surface of the first spool near to the rupture disc was honed to a roughness better than 1.0 µm. Eight pressure transducers and three temperature probes were mounted along the length of the shock tube. The locations of the pressure transducers and temperature probes are shown in Table 1. The test program consisted of a total of 10 tests performed with various gas mixtures at different initial conditions. In this work, three different gas mixtures undergoing decompression were simulated using the CFD model: pure nitrogen, conventional natural gas and rich natural gas. The objective is to examine whether the current CFD model can predict the decompression behaviour of gas mixtures. Firstly, the decompression process of pure nitrogen was simulated to examine the decompression behaviour of single-phase flow. The second set of simulations was performed for conventional gas mixture (case 2) with ~95% Methane where the flow is expected to cross the two-phase region. The rich natural gas mixture contained ~80% of Methane for the third case. The distinctive feature of this mixture is that the two-phase region appears at high pressure (~7.5 MPa). Table 2 lists the initial conditions and the gas compositions of these cases. The predicted results of gas decompression velocity were compared with the measured data of shock tube test, and also against results from OLGA and GASDECOM. CASE 1: DECOMPRESSION OF PURE NITROGEN Figure 4 shows the predicted pressure-time traces of pure nitrogen at four locations: P3, P4, P6 & P7 along the pipeline. It can be seen that the pressure at these monitored points starts to drop rapidly once the decompression wave front reaches each location in turn. The pressure becomes gradually steady at pressures below 4 MPa. The rate of change in both speed of sound and outflow velocity as functions of time is presented in Figures (5 & 6) respectively. The results of outlet velocity indicates that the flow was at rest (v = 0 m/s) inside the pipe until the outlet boundary was removed. Comparison of the predicted average and local decompression wave speed with the measured data is shown in Figure 7. The results of both GASDECOM and OLGA are also presented. The comparison shows that the current 2D CFD decompression model predicts the decompression wave speed well. A good agreement can be observed between the measured data and the average decompression wave speed calculated using the pressure-time traces (similar way of measured calculation). Note that the measured data of decompression wave speed deviates further to the left forming a plateau approximately at pressure ratio of 0.4. The

local decompression wave speed predicted at the four locations is identical to that predicted by GASDECOM yet a constant local decompression wave speed was observed at the latter stages of the decompression. This shows that using the same definition of W, GASDECOM and the CFD model obtained the same results, indicating that GASDECOM performs as good as CFD (at a fraction of the computation time) despite using a different principle and different equation of state. The only difference that GASDECOM calculates the decompression speed independently of time or location while the CFD results were predicted at several locations away from the outlet boundary and inherently functions of the time. CASE 2: DECOMPRESSION OF LEAN NATURAL GAS Figure 8 shows the CFD prediction and the measured pressure-time traces of the lean gas mixture. The predicted pressure starts to decompress at nearly the same time as the measured data at every locations. This indicates that the decompression wave speed was at the correct operating pressure and temperature. The predicted rate of change in pressure was consistent with the experimental results until the appearance of a kink in the P-t curves at pressures slightly below 5 MPa. This kink is due to the change in fluid properties which results from the phase change. After this stage the measured pressure time traces become almost steady while the predicted P-t traces drops further. At the same time, a sharp drop in the speed of sound was noticed as shown in Figure 9. The observed kink also appeared in the curves of outflow velocity and temperature results, as shown respectively in Figures 10 & 11. By referring to the results of pressure and temperature, the kink was observed to occur at P=4.6 MPa and, T=200.24 K. This point was found to be exactly at the phase boundary of the mixture as expected. Comparison of the predicted decompression speed and the measured data is shown in Figure 12. The predicted results of the current CFD model are more consistent with the measured data than that predicted by GASDECOM and OLGA although there is still discrepancy at the end of the decompression. Referring to Figure 12, a significant decrease in the decompression wave speed is observed, forming a pressure plateau at a pressure ratio of ~0.43 when crossing the dew curve. Figure 13 shows the phase envelope of case 2 and the point at which the decompression enters the two-phase region. A significant drop in the speed of sound at a constant pressure occurs during this process.

The current CFD decompression model successfully predicted the plateau found in the measured data. The decompression speed calculated using the pressure time traces best compared with the measured results. The appearance of such plateau for gas mixtures is significant in terms of fracture propagation arrest requirements as it can lead to high minimum arrest toughness. GASDECOM was not able to predict the plateau for this mixture although its results show good agreement at high pressure ratio while OLGA predictions were inaccurate. The issue in OLGA was linked to the definition used to calculate speed of sound [19, 21]. CASE 3: DECOMPRESSION OF RICH NATURAL GAS Figure 14 shows a comparison between the predicted and experimental results of pressure-time traces of case 5 recorded at locations 4, 6 & 7. The predicted pressure-time traces were in good agreement with the measured data. The two-phase flow region was predicted and clearly observed on the P-t curves for this mixture. The discrepancy between the predicted and the measured data occurred at pressure ratios lower than 0.5. The rate of change for speed of sound, outflow velocity and temperature are shown in Figures 15, 16 & 17 respectively. In Figure 15, a sharp drop in the speed of sound occurred once the decompression crossed the two-phase boundary. The sharp decrease in speed of sound was reflected on the result of decompression wave speed where a long plateau was formed as shown in Figure 18. From this figure, it should be seen that the predicted average and local decompression speed is in good agreement with the measured results except at low pressure ratios where the predicted results deviate further to the right. The major observation for this mixture is that the speed of sound increases towards the end of the decompression until it finally levels off at around 250 m/s. To emphasis that this model predicted the decompression behaviour of gas mixtures with taking account for phase change, the decompression process of pressure against temperature was plotted on the phase envelope of the mixture. The procedure was also performed as in lean gas mixture to find out at which point the flow entrance the twophase region. Figure 19 highlights the point where the flow crosses the two-phase boundary. DISCUSSION It is been found that if the shape and temporal location of the predicted pressure matches the experimental results, the predicted value of W, should be in close agreement with

the measured data. As seen from plots in Figures 12 &18, there is a slight difference at the plateau between the measured and the predicted results, however this difference is small as it is ranging between 2.9-4.8% for case 2 and 1.5-1.8% for case 3. While the reason behind this difference might be due to the uncertainties inherent in the numerical method that could possibly be improved. Other factors such as delayed nucleation and/or the rapid dynamics of the phase change (whereas the EOS uses equilibrium conditions at all time), roughness, thermal exchange all can influence the results to various degrees. The predicted results of decompression wave speed based on the local formulation deviate at the latest stage of the decompression process in all cases. W becomes constant with a continuing drop in pressure. This phenomenon begins at different values of pressure ratio and depends on the location. At the location nearest to the full bore opening, W became constant at lower pressure ratio than that further away from the outlet. This indicates that the difference between the speed of sound and the outflow velocity became constant. This trend is likely to be related to the increase of entropy due to the turbulences at the vicinity of the wall. The increase of entropy limits the decrease of the local speed of sound and limits the rate of increase of the flow velocity. The combined effects on the speed of sound and flow velocity indirectly limit the local decompression wave speed. Further work is under way to study the phenomenon. The most notable difference between the numerical and measured data is that the measured data shows a pressure drop downstream of the decompression wave levelling off at the latter stages of the decompression. Experimental measurements show that this levelling off occurs more rapidly at a higher pressure than that predicted by the model. For locations P3 and P4, where results are available for the longest duration of time after the initial arrival of the decompression wave, the difference between the low pressure values is on the order of 0.4 to 0.8 MPa. This is also reflected on the decompression wave speed curves where, in all cases, the measured data deviates further to the left creating an apparent second plateau. This is attributed to the piezoelectric pressure transducers (PCB) used in these particular experiments. The apparent increase in the measured pressure-time traces (Figs. 8 and 14) should be taken with caution at these low pressure parts of the traces. Later tests using Endevco pressure transducers did not exhibit this increase in the pressure time traces [18, 20]. CONCLUSION The following remarks can be inferred from this study.

(a) A two-dimensional CFD model based on the finite volume technique was developed using ANSYS-Fluent to simulate the decompression behaviour of gas pipelines. (b) The GERG-2008 EOS was successfully implemented into this model to provide an accurate calculation of the thermodynamic properties for gas mixtures containing 21 different components. (c) The results predicted by the current CFD model are in good agreement with the experimental results obtained from shock tube tests. The current work demonstrates that the CFD technique can be used to predict rapid and severe gas decompression by solving the governing flow equations, in conjunction with the wide-range GERG-2008 EOS. This is an effective tool for determining the decompression wave speeds for different gas mixtures. This tool is valuable to develop fracturedecompression coupled models and deliver a better understanding of the interaction between the decompressing fluid and the fracturing pipeline. It is applicable in two or three dimensional problems and for large pipe sizes where experimental tests are impractical. Future work is also needed to examine the capability of this model in simulating the decompression behaviour of other fluids such as CO 2 mixtures. ACKNOWLEDGMENTS This work was funded by the Energy Pipelines CRC, supported through the Australian Government s Cooperative Research Centres Program. The funding and in-kind support from the APIA RSC is gratefully acknowledged. REFRENCES [1] S. F. Brown, "CFD Modelling of Outflow and Ductile Fracture Propagation in Pressurised Pipelines," Doctor of Philosophy, Department of Chemical Engineering, University College London, London, 2011. [2] H. Mahgerefteh, G. Denton, and Y. Rykov, "A hybrid multiphase flow model," AIChE Journal, vol. 54, pp. 2261-2268, 2008. [3] W. A. Maxey, J. F. Kiefner, and R. J. Eiber, Ductile fracture arrest in gas pipelines, 1976. [4] K. K. Botros, J. Geerligs, L. Fletcher, B. Rothwell, P. Venton, and L. Carlson, "Effects of Pipe Internal Surface Roughness on Decompression Wave Speed in Natural Gas Mixtures," ASME Conference Proceedings, vol. 2010, pp. 907-922, 2010d. [5] C. Lu, G. Michal, A. Elshahomi, A. Godbole, P. Venton, K. K. Botros, L. Fletcher, and B. Rothwell, "INVESTIGATION OF THE EFFECTS OF PIPE WALL ROUGHNESS AND PIPE DIAMETER ON THE DECOMPRESSION WAVE SPEED IN NATURAL GAS PIPELINES," presented at the 9th International Pipeline Conference 2012, Calgary, Alberta, Canada, 2012.

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Table 1: The location of monitoring points Node P2 P3 P4 P6 P7 Location from Rupture (mm) 59 240 440 840 1240 Table 2: Shock tube test conditions and gas compositions Case 1 Case 2 Case 3 P i (MPa) 10.059 10.58 9.94 T i (K) 260.44 247.65 273.21 C 1 (mole %) 0 95.4741 79.3089 C 2 (mole %) 0 2.9363 14.1967 C 3 (mole %) 0 0.1902 5.2556 ic 4 (mole %) 0 0.0156 0.0114 nc 4 (mole %) 0 0.0253 0.0164 ic 5 (mole %) 0 0.0041 0.0029 nc 5 (mole %) 0 0.003 0.002 C 6+ (mole %) 0 0.0013 0.0009 N 2 (mole %) 100 0.5689 0.5513 CO 2 (mole %) 0 0.7812 0.6539

Figure 1: ANSYS Fluent density based solver integrated with GERG-08 EOS Figure 2: Flow domain and computational domain schematic -

Figure 3: Two-dimensional computational grid Figure 4: Pressure-time traces at four locations (case 1)

Figure 5: Speed of sound-time traces at four locations (case 1) Figure 6: Outflow velocity-time traces at four locations (case 1)