Assessment of gas detection strategies for offshore HVAC ducts based on CFD modelling

Transcription

1 Health and Safety Executive Assessment of gas detection strategies for offshore HVAC ducts based on CFD modelling Prepared by the Health and Safety Laboratory for the Health and Safety Executive 2007 RR602 Research Report

2 Health and Safety Executive Assessment of gas detection strategies for offshore HVAC ducts based on CFD modelling Dr C J Lea & Dr M Deevy Health and Safety Laboratory Harpur Hill Buxton SK17 9JN The aim of this study has been to undertake CFD modelling to provide a basis for advice to inspectors and the industry on the effectiveness of flammable gas detection strategies for offshore HVAC ducts. CFD simulations of a high and low pressure gas release have been undertaken for idealised representations of an offshore platform, as well as a high pressure release for a more realistic geometry based on the Brae Alpha platform. In parallel with this modelling work a literature review has been carried out to build on a scoping study by Walsh et al (2005). This report and the work it describes were funded by the Health and Safety Executive (HSE). Its contents, including any opinions and/or conclusions expressed, are those of the author alone and do not necessarily reflect HSE policy. HSE Books

3 Crown copyright 2007 First published 2007 All rights reserved. No part of this publication may be reproduced, stored in a retrieval system, or transmitted in any form or by any means (electronic, mechanical, photocopying, recording or otherwise) without the prior written permission of the copyright owner. Applications for reproduction should be made in writing to: Licensing Division, Her Majesty s Stationery Office, St Clements House, 2-16 Colegate, Norwich NR3 1BQ or by to hmsolicensing@cabinet-office.x.gsi.gov.uk Acknowledgements The authors would like to thank Marathon Oil UK Ltd, Flamgard Engineering Ltd, Integrated Engineering Services Ltd, Groveley Detection Ltd and Honeywell Analytics for providing information and assistance during the course of this project. Dr Peter Walsh, HSL, has also provided helpful guidance. ii

4 CONTENTS 1 INTRODUCTION 1 2 FLOW AND DISPERSION OF GAS IN A DUCT Turbulent flow in pipes and ducts Dispersion in pipes and ducts Summary 8 3 INGESTION OF GAS INTO HVAC INLETS Introduction Scenario Scenario DISTRIBUTION OF GAS INSIDE AN HVAC DUCT Overview Results 33 5 GAS RELEASE AND INGESTION INTO HVAC DUCTS FOR A 37 REALISTIC SCENARIO 5.1 Introduction Model geometry and computational domain Boundary conditions and flow physics Mesh, solution numerics and convergence Results 45 6 DISCUSSION Review and discussion of results Initial recommendations for gas detection strategies in HVAC ducts 56 7 CONCLUSIONS Recommendations for further work 59 8 APPENDIX 1: EXAMPLES OF COEFFICIENT OF VARIATION 60 9 REFERENCES 61 iii

5 iv

6 EXECUTIVE SUMMARY Objectives The aim of this study has been to undertake CFD modelling to provide a basis for advice to inspectors and the industry on the effectiveness of flammable gas detection strategies for offshore HVAC ducts. CFD simulations of a high and low pressure gas release have been undertaken for idealised representations of an offshore platform, as well as a high pressure release for a more realistic geometry based on the Brae Alpha platform. In parallel with this modelling work a literature review has been carried out to build on a scoping study by Walsh et al (2005). Main Findings The most significant finding is that in all of the CFD simulations the distribution of gas at HVAC inlets is non-uniform: large variations in gas concentration are present over the crosssection of the modelled HVAC inlets. This result is consistent with the theoretical behaviour of a high pressure gas release. Neither is there a substantial reason to believe that it is not also consistent with the behaviour of a low pressure gas release. It means that there is the potential for a gas release to be missed by detection systems unless this non-uniformity in gas concentration is anticipated in the selection and siting of gas detectors at HVAC inlets. The CFD results also show that a variation in gas concentration over a duct cross-section only reduces slowly with distance along a straight duct. This finding is consistent with a body of relevant literature stemming from the sampling of gas distributions in the exhaust ducts of nuclear stacks. This literature highlights that purpose-designed mixing elements and bends in a duct can be effective in creating well-mixed conditions in a duct, but at the cost of increased pressure drop. It also suggests that relatively small-scale obstructions such as louvres and fire dampers are unlikely to significantly enhance mixing. This is borne out by CFD modelling of such obstructions in this study. The implications of the modelling work, substantiated by the literature, are that in the absence of purpose-designed mixing elements or a series of bends upstream from gas detectors no significant benefit would be gained from siting detectors a significant distance downstream from an HVAC inlet. Also, no significant benefit can be expected to be gained from siting detectors inside an HVAC duct compared to locating them immediately outside the HVAC inlet. The CFD results have been post-processed to gain further insight into the likely effectiveness of point and infra-red beam or aspirated detector systems for HVAC ducts. The resulting output has been combined with the findings of the scoping study by Walsh et al (2005), and the outcomes of the literature review and CFD modelling as reported above, to provide a basis for the following initial recommendations on flammable gas detection strategies: x Detector alarm levels should be set as low as reasonably practical: 10% LEL or less. x Point catalytic, point infra-red, extended path point infra-red, cross-duct beam infra-red and aspirated point detector systems all have the potential to be effective in detecting non-uniform distributions of flammable gas in and around HVAC ducts provided that their sensitivity is sufficiently high (low detection limit) and that due regard is given to the possibility that gas will be distributed non-uniformly. v

7 x Extended path point infra-red detector systems currently appear to offer the greatest sensitivity, but multiple detectors should be used and sited so as to anticipate nonuniform mixing. x Cross-duct beam infra-red, extended path or aspirated point detector systems should be based on two approximately orthogonal beams or lines of aspirated point probes. x No significant benefit can be expected to be gained from siting detectors inside an HVAC duct compared to locating them immediately outside the HVAC inlet. x In the absence of purpose-designed mixing elements or a series of bends upstream from gas detectors no significant benefit is to be gained from siting detectors a significant distance downstream from an HVAC inlet. x Mixing elements have the potential to reduce any non-uniformity in the distribution of gas in a duct but their effectiveness should be proven by physical tests. Recommendations It is recommended that large-scale physical tests using commercial detectors are undertaken to validate the findings of this study and to support and refine the initial recommendations above. It is also recommended that the minimum practical alarm levels for a range of commercial detector types, especially cross-duct beam infra-red systems, are clarified. Reference Walsh P, Johnson A, Ivings M (2005) Effectiveness of gas detection in HVAC ducts: Scoping study. Health and Safety Laboratory Report FM/04/11. vi

8 1 INTRODUCTION The accidental release of flammable gas on offshore installations can potentially lead to the build-up of an explosive mixture. Natural or forced ventilation can help to mitigate such incidents and gas detection systems play a key role in reducing the risks from releases by enabling early detection and subsequent interventions. The provision and siting of gas detectors for open areas and gas turbine enclosures has been studied over a number of years and is comparatively well documented (HSE 1993, Kelsey et al 2000, Ivings et al 2002, HSE 2003, Kelsey et al 2005). However there is much less information available on gas detection systems for HVAC ducts supplying air to accommodation modules, temporary refuges or process areas on an installation. The information which is available on offshore HVAC ducts tends to address other aspects such as mechanical design with minimal information on gas detection systems (BS EN ISO 15138:2001, Chin & Lam 2004). Following an incident on the Brae Alpha platform in November 2004 in which there was a delay in confirmed detection and shutdown of the HVAC system, despite gas being ingested into the HVAC inlets, it became clear that there was a need for a study to establish the current state of the art. This led to the initiation of the scoping study by Walsh et al (2005) into the effectiveness of gas detection systems for HVAC ducts on offshore installations. This was carried out in response to concerns raised by HSE offshore safety inspectors on the positioning of flammable gas detectors in or around HVAC inlets. Walsh et al (2005) reviewed national standards, industry guidance, published papers and reports to obtain information on the siting of flammable gas detection systems for offshore HVAC inlets. Some industry guidance was obtained on performance targets for HVAC gas detection systems specific to a particular platform, stating that flammable gas at concentrations above 20% LEL (Lower Explosive Limit) at ventilation intake ducts should lead to an alarm and that concentrations above 60% should lead to isolation of the intake by stopping fans and closing dampers (Micropack, 2000). Other industry guidance advised that either point or beam detectors may be installed inside or outside of HVAC ducts and that for point detectors three should normally be installed to maintain voting logic (Shell UK, 1995). Guidance has also been obtained from a gas detector company advising on the appropriate proximity of point detectors to features such as bends or diffusers in gas turbine ductwork. Walsh et al (2005) surveyed commercial gas detector systems for HVAC applications. They found that a number of differing systems are available: point catalytic, point infra-red, beam infra-red or aspirated systems. In general the operating range of such systems was found to be up to 100% LEL, with a minimum alarm level of 20% LEL. However, Walsh et al noted that recent developments in HVAC gas detection technology have led to the availability of extended closed path infrared point detectors which measure an average concentration over a path length of typically around 1 m. Such systems have increased sensitivity and typically allow for minimum alarm levels of 5% LEL. Recent developments in catalytic point sensors for the industrial gas turbine industry also apparently allow for minimum alarm levels of 5% LEL. However, Walsh et al (2005) found essentially no guidance specific to the siting of gas detectors in and around HVAC inlets or on the effectiveness of differing siting strategies. They recommended that further work be undertaken to remedy this situation. In consultation with the Offshore Safety Division (OSD) in HSE they proposed that this further work should initially be based on Computational Fluid Dynamics (CFD) modelling, recognising that CFD can guide any potential future experimental work by helping identify effective and ineffective gas detection systems. This provides the basis and motivation for the project which is documented in the present report. 1

9 The overall aim of this project is therefore to apply CFD modelling to provide a basis for advice to inspectors and the industry on the effectiveness of flammable gas detection strategies for offshore HVAC ducts. It builds upon the scoping study of Walsh et al (2005). The primary focus of the project is on siting strategies for gas detectors located in or around HVAC inlets for circumstances in which gas is not ingested uniformly over the entire inlet. The present project and the earlier scoping study have both been funded by OSD. In Section 2 of this report the behaviour of flow and dispersion of gas in a duct is discussed. This discussion is supported by a literature review that encompasses sources of information which compliment that from the field of gas detection in offshore HVAC ducts. Section 3 addresses circumstances by which a non-uniform distribution of gas could be present immediately outside or inside of an HVAC inlet. This is of particular relevance to the present project. Thus, if a well-mixed and uniform distribution of gas and air is present then the positioning of detectors should have little impact on the performance of a system in detecting gas; instead the performance will be governed by other factors such as the sensitivity, speed of response and reliability of the detector. Such information is available from gas detector companies supported by experience of use. Therefore in this section the outcome of CFD modelling of gas releases is presented for two differing idealised scenarios showing how a nonuniform distribution of gas and air may be present in and around an HVAC inlet. Section 4 reports the mixing and distribution of gas inside an HVAC duct as simulated using CFD. The modelling extends that presented in Section 3 by focusing on the detail of the geometry and associated flow features likely to be found in an actual HVAC inlet. Section 5 presents CFD results for a more realistic scenario based on the release from the Brae Alpha platform in November Although the most significant elements of this incident are represented, the scenario is not a detailed replication of the incident; this would be very timeconsuming to undertake and may not be practical. The value of these results instead lies in the more general lessons which they provide on the behaviour of a release in a complex geometrical environment and which is subsequently ingested into HVAC ducts of varying size. The results from the CFD modelling are discussed in Section 6. Their implications for gas detection strategies for HVAC ducts are highlighted. Initial recommendations on the effective siting of flammable gas detectors in and around HVAC inlets are also provided. Conclusions and recommendations for further experimentally-based work are given in Section 7. Appendix 1 provides illustrative values of a parameter which can be used to characterise the degree of mixing in a duct. References are provided in Section 9. All of the CFD modelling has been undertaken using ANSYS CFX 10, a commercial software package. 2

10 2 FLOW AND DISPERSION OF GAS IN A DUCT 2.1 TURBULENT FLOW IN PIPES AND DUCTS An introduction to the flow regime in HVAC ducts is provided by Walsh et al (2005). They show that for offshore HVAC ducts the Reynolds number will be sufficiently high that the flow will be fully turbulent. This is based on a typical duct velocity of 5 m/s as outlined in BS EN ISO and a range of duct hydraulic diameters from 0.5 to 5 m, giving Reynolds numbers 6 in the range of approximately 10 5 to 10. Turbulence is often assumed to be synonymous with good mixing. Although this is generally the case, turbulent mixing does not occur instantaneously. It would be wrong to assume that because the flow in HVAC ducts is turbulent any non-uniformity in a distribution of gas at an HVAC inlet will very quickly be dispersed to give well-mixed uniform conditions. This has a significant impact on the siting of gas detectors in HVAC ducts. An indication that this is the case is given by experimental data on turbulent developing flow in a pipe. Sufficiently far downstream from the entrance to a long straight pipe the flow eventually reaches a condition in which it no longer changes. It is then said to be fully-developed. For turbulent flow in a smooth pipe the fully-developed mean velocity profile approximately follows a 1/n power law (Hinze, 1975): 1 / n U 2x U d ¹ max where U is the velocity at distance x from the wall of a pipe of diameter d and U max is the peak 5 velocity on the axis of the pipe. The constant, n, is 7 for a Reynolds number of 10 and approximately 9 for a Reynolds number of For rough pipes, n increases to between 4 and 5. The turbulent mean velocity profile is therefore rather flat across much of the cross-section of a pipe, dropping away steeply at the walls. Turbulent flow in a pipe is therefore often characterised as consisting of a relatively thin near-wall region and a central core. The distance before fully-developed conditions are established in a pipe is known as the development length. The development length depends on a number of factors: the Reynolds number, geometry of the inlet and flow conditions upstream of the inlet. For flow from quiescent surroundings into a straight pipe with rounded edges at its inlet the development length for fully-developed turbulent flow is given by White (1987): L e d 4.4Re 1/ 6 where L e is the development length and d is the pipe diameter. The development length is 6 approximately 30 d to 40 d for Reynolds numbers in the range of interest, i.e to 10. For laminar flow (Reynolds number below about 2000 to 4000) the development lengths are much greater. The above expression applies to idealised conditions in which the flow is from quiescent surroundings and the pipe is straight and rounded at its inlet. In these circumstances the flow which enters the pipe is initially uniform and laminar. A laminar boundary layer grows on the surface of the pipe and then rapidly becomes turbulent. The turbulent boundary layer grows in 3

11 thickness until eventually it extends to the axis of the pipe. Fully-developed flow conditions are reached shortly afterwards. Klein (1981) correlates data on developing flow in pipes from a range of sources, which include sharp-edged inlets. These data show that fully-developed flow is not established until greater than about 50 d downstream from the inlet. Other workers (Laws et al, 1979) have indicated that in the presence of significant obstructions at the inlet to a pipe, fully-developed flow is not established until significantly further downstream, possibly greater than 100 d. It should be noted that these studies are seeking flow profiles which are exactly the same with distance from the pipe inlet and so are rather rigorous in their requirements. As a practical rule of thumb Hinze (1975) recommends that turbulent fully developed flow can be assumed to occur in straight pipes after a minimum development length of 40 d. The turbulent developing flow in a straight square or rectangular duct behaves in a broadly similar manner, in that the distance to fully-developed flow is not short. Melling & Whitelaw (1976) present data which show that turbulent fully-developed flow in a square duct is reached at about 25 duct widths from the inlet. Other measurements, by Gessner et al (1977), show that fully-developed flow is not reached until much further downstream, at between 40 to 84 duct widths downstream from a rounded inlet. The development of the flow occurs in much the same manner as that for a circular pipe, with the gradual growth of boundary layers on the duct walls until they occupy the entire section of the duct. There is a further feature of flow in straight square or rectangular ducts which should be mentioned. This is the existence of a turbulence-driven secondary flow in the cross-stream direction. The secondary flow is of a very much smaller magnitude than the primary flow along the axis of the duct (Demuren & Rodi, 1984). Its effect is to slowly transport material from the centre of the duct towards the walls and back again. Overall this leads to an enhancement in turbulent mixing. This secondary flow presents a difficulty for CFD models. Simulations based on the commonlyused k-h turbulence model (Launder & Spalding, 1972) completely fail to capture this secondary flow (Speziale, 1996). More advanced turbulence models are available (Gatski & Speziale 1993, Gatski & Rumsey 2002) which do capture the secondary flow but they often tend to be more problematic to apply and require more computational resources. However, the secondary flow is not fully established immediately the flow enters a duct; in fact it is only completely established when fully developed conditions are met far downstream. In practice, offshore gas detectors for HVAC ducts are nearly always located close to the inlets. This is partly to minimise the volume of gas ingested following action to isolate an HVAC system upon detection of gas, and partly to provide access for maintenance and calibration. Therefore, although a standard k-h turbulence model cannot capture these secondary flows, the resulting error is unlikely to be significant in the few duct widths immediately downstream from an HVAC inlet, i.e. the region in which gas detectors are commonly found. The literature on developing and fully-developed flow in pipes and ducts gives a useful indication that equilibrium in the turbulent diffusion of momentum responsible for the creation of fully-developed flow profiles only occurs at a considerable distance downstream from the inlet to a straight pipe or duct. Of course, the offshore environment is somewhat different from the idealised experiments upon which this literature is based: offshore HVAC inlets do not have rounded edges; often they are sharp-edged and have obstructions such as grilles or weather louvres across their entire face. Figure 1 shows some examples of such features. 4

12 Figure 1: Examples of HVAC inlets Usually, dampers will be encountered a short distance inside an HVAC duct; to provide isolation from fire and gas in the event of an incident. The dampers will normally be open and parallel to the main flow direction, but they still present an obstruction to the flow, as does their supporting structure. Overall, the gross effect of obstructions such as grilles, louvres and dampers will be to generate turbulence which will lead to enhanced mixing. 2.2 DISPERSION IN PIPES AND DUCTS The above discussion has concentrated on the characteristics of turbulent flow in straight pipes and ducts. Our main interest is in the turbulent transfer of mass, or more specifically, in the turbulent diffusion of a contaminant such as natural gas at a relatively low concentration. It should be noted that at low concentrations in the flammable range and for the velocities typically encountered in offshore HVAC ducts, a natural gas mixture can be regarded as a passive contaminant, i.e. it is transported by, but does not directly influence the flow. This is because the Richardson number (Simpson, 1997) is likely to be at least an order of magnitude too low for any turbulence-modifying effects of a slightly buoyant gas to dominate over shearinduced turbulence. The turbulent diffusion of a contaminant which is essentially passive takes place in an analogous manner to that of the turbulent diffusion of momentum. Both are the result of the same turbulent flow-field. It can therefore be expected that the dispersion of a passive contaminant in a duct will follow roughly the same behaviour as that of the development of the velocity profile: if the development length is long then the distance to uniformly mixed conditions will also be long. A review of the literature was undertaken to obtain more information on the development of an initially non-uniform distribution of passive gas in a duct or pipe, in the presence or absence of obstructions. Initially this review focused on a search for papers in the field of gas dispersion, or detection, in ducts or HVAC systems. This produced little useful information on the distribution and mixing of gas in ducts. A search for papers in the field of sampling, rather than detection or dispersion, proved much more fruitful. Hence a number of papers were obtained on the sampling of exhaust duct stacks in the nuclear industry (Hampl et al 1986, Rodgers et al 1996, McFarland et al 1999a & 1999b, Anand et al 2003 and Seo et al 2006). These papers report research into the mixing of a passive tracer in circular, square and rectangular ducts in the presence or absence of bends and mixing elements. 5

13 In most cases the tracer was released from a single location on the axis of a duct. Multiple point concentration measurements were then made across the entire cross-section of the duct at a number of axial locations. This body of research was largely undertaken to support an improvement and updating of American standards on the sampling of gaseous radionuclide emissions from stacks. Sampling had historically been required at multiple points over the cross-section of a duct, with as many as 20 points being required for large ducts (Rodgers et al, 1996) to ensure that peaks in the concentration distribution were not missed. The sample plane was required to be no closer than eight duct diameters from the nearest upstream flow disturbance and two duct diameters from the nearest downstream disturbance. However, Hampl (1986) suggested that up to 50 duct diameters may be needed for near-uniform mixing of a passive tracer in a straight pipe, but less than two duct diameters for acceptable mixing downstream of two elbows. Proposals for a new approach were therefore made, based on a single point sample taken at a location where acceptable well-mixed conditions could be assured. To characterise the degree of mixing, a parameter known as the Coefficient of Variation (COV) was introduced. This is defined as: COV 1 N N ( C i C mean 1 i 1 C mean ) 2 where N is the number of samples at a particular downstream location, C i is the concentration of the ith sample and C mean is the mean concentration over all samples at that location, defined as: 1 N C C mean N i 1 i As an example, if the concentration distribution is such that across one half of a duct the concentration is a uniform 30% LEL whilst across the other half the concentration is a uniform 10% LEL, then the COV would be 50%. Appendix 1 provides examples of concentration distributions for a range of COV between 10% and 150%. The updated American standards (ANSI/HPS N ) allow for single point sampling of gaseous contaminants in a duct if the COV for both velocity and concentration of a tracer gas are less than 20% over the central two-thirds of a duct. We do not suggest or comment on the practicality or appropriateness of these criteria for offshore HVAC ducts. However, the notion of a COV is helpful in quantifying the uniformity of mixing in a duct: it is readily computed from CFD results. In addition, this body of research on stack sampling provides much useful data on how the COV is affected by a range of configurations; this can be used to inform the present research. Thus McFarland et al (1999a) showed that for a tracer released at a single point on the axis of a duct the COV can readily be reduced to ~5% following passage through a compact mixing chamber installed in the path of a duct. However, the pressure loss for this mixing chamber was not small (non-dimensional pressure coefficient of approximately 4.5, equivalent to a head loss of about 70 Pa for an average velocity of 5 m/s) and, more significantly, it would require reengineering of existing HVAC installations. Measurements of the COV in straight circular pipes by Anand et al (2003) show that the distance from a point release of a tracer to that at which the COV is less than 20% depends on the upstream turbulence intensity. For a low turbulence intensity of 1.5% the COV falls very slowly with distance and is still over 100% at 30 duct diameters downstream from the point of 6

14 release. Even with a high turbulence intensity of 10% - generated by passing the flow through an array of thick rods a COV of 20% was still not reached after 25 duct diameters downstream. McFarland et al (1999b) investigated the effect of bends and static mixing elements on the COV. They show that a single smooth 90 o bend in a circular duct still requires a distance of nine diameters downstream from the bend before the 20% COV criterion is met. The performance of the static mixing elements was very variable: all resulted in the 20% COV criterion being met within nine diameters downstream, but the most effective mixers were able to meet the criterion within three duct diameters of the mixing element. The most simple and effective mixing elements consisted of two large flow deflectors attached to opposite walls of a duct, leading to a slot-like opening in the centre of the duct. Two or more of these mixing elements were used in series. The common characteristic of the most effective mixing elements appears to be the introduction of cross-stream mixing from one side of a duct to another due to the introduction of large turbulent eddies (Seo et al, 2006). Mixing elements which only introduced flow swirl were less effective. One disadvantage of these simple deflector mixing elements is a relatively large non-dimensional pressure coefficient (5.0) for two elements in series. Seo et al (2006) examine the behaviour of the COV in square and rectangular ducts (aspect ratio of 3:1) with and without bends. They report that the COV is similar for circular and squaresection ducts at large distances downstream from the tracer release point, both with and without bends. This implies that the main findings of the previous work on circular-section ducts, discussed above, are likely to largely carry over to square-section ducts. However, a significant difference was found when the COV for the square and rectangular-section ducts were compared: typically the COV was much higher for the rectangular duct at any given distance downstream from the point of release, by about a factor of four. The COV for rectangular ducts with bends are also consistently higher than the same flow configuration in a square duct. Seo et al speculate that this is because turbulent eddies in a wide duct have less opportunity to effectively transfer mass and momentum from one side of a duct to another. The effect of grilles on the turbulence in a duct is relatively well understood. Laws & Livesey (1978) explain that the effect is either to suppress, or enhance turbulence, and this depends on the geometry of the grille. Thus a very fine grille, or mesh, will tend to suppress turbulence. Any turbulence which is introduced by the mesh decays quickly due to its small scale. A grid of relatively large diameter rods will enhance turbulence, although Laws & Livesey state that it is difficult to achieve a turbulence intensity of much higher than 10%. It is not clear whether grilles typically used to cover HVAC inlets will suppress or enhance turbulence. However, even if turbulence is significantly enhanced, Anand et al (2003) show that the COV will remain high for long distances downstream. For an inlet turbulence intensity of 10% created by an array of rods they found that the COV was still over 50% at 15 duct diameters downstream from the rod array. The reason for the relative ineffectiveness of such devices on mixing is that they introduce turbulence on too small a length-scale. Much of the above work is based on a flow which is well-controlled at the inlet to a duct, for example by use of a rounded entrance or other flow control devices. In the case of offshore HVAC installations this will not be the case: turbulence can be generated by large-scale flow separation at the sharp-edged entrance to a duct or is already present in the ambient flow outside of the duct. McFarland et al (1999b) note that the earlier work of Hampl et al (1986) was based on a sharp-edged inlet and in comparison to their later research with a well-controlled approach flow the COV were found to be reduced by between a factor of two to three: flow separation at the inlet enhances mixing inside a duct. Nevertheless, it should be recalled that Hampl et al still suggest that up to 50 duct diameters may be needed for near-uniform mixing of a passive tracer in a straight pipe, even with a sharp-edged inlet. It is difficult to know to what extent the effect 7

15 of turbulence which is present in the ambient flow affects the mixing inside a duct. However, the gross effects of pre-existing turbulence in the outside ambient flow on mixing in a duct should be represented, albeit crudely, by the CFD modelling which follows in Sections 3, 4, and 5. It is also significant that both Anand et al (2003) and Seo et al (2006) note that the COV is littleaffected by the Reynolds number. Seo et al state that for a square duct there is only a small effect of the Reynolds number on COV over a range from 25,000 to 150,000. For a rectangular duct with a 3:1 aspect ratio there is a significant dependence on the COV for a Reynolds number below 50,000, but relatively little effect at higher Reynolds number. Both Anand et al (2003) and Seo et al (2006) conclude that, for fully turbulent flow, mixing is primarily dependent on geometry. 2.3 SUMMARY The main findings of relevance to the present project from the above literature review are as follows: x The Reynolds number for offshore HVAC ducts is sufficiently high that the flow will be fully turbulent. x A flow development region exists downstream from the inlet to a duct; in the absence of obstructions and for a smoothly rounded inlet the length of this region can be assumed to be a minimum of 40 duct diameters for a straight duct whereupon the flow is said to be fully-developed and does not change further downstream. x In the development region the flow may not be turbulent across the entire section of a duct: in this region it is more likely to be turbulent in the presence of a sharp-edged entrance to the duct and where the duct contains obstructions such as louvres. x A weak secondary flow exists in square and rectangular ducts which will enhance mixing. In CFD modelling this secondary flow is not captured by the most commonlyused k-h turbulence model. However, the secondary flow is not completely established until the flow is fully developed, so associated errors from use of the k-h model are unlikely to be significant close to an HVAC inlet. x A natural gas mixture which is ingested into offshore HVAC ducts at concentrations in the flammable range can generally be treated as a passive contaminant. x The turbulent diffusion of a passive contaminant in a duct will take place in a broadly analogous manner to that of the turbulent diffusion of momentum; long development lengths imply similarly long distances before well-mixed conditions are obtained. x The degree of mixing in a duct can be quantified by a Coefficient of Variation, COV. x Uniform mixing of a point release of a passive tracer in a duct typically requires 50 duct diameters downstream from the release for a straight circular duct. The presence of bends and mixing elements can significantly reduce this distance. x The turbulent mixing in a rectangular duct with a large aspect ratio (3:1 and above) is significantly less effective than that in a square or circular duct. x Effective mixing in a duct is driven by large-scale turbulent eddies which are able to transfer mass and momentum across the duct. x The most effective mixing elements are those which introduce large turbulent eddies. x Grilles at the entrance to HVAC ducts are likely to be ineffective in introducing largescale turbulent eddies and so will do little to enhance mixing. x For fully turbulent flow, mixing in a duct is primarily dependent on geometry. 8

16 3 INGESTION OF GAS INTO HVAC INLETS 3.1 INTRODUCTION The circumstances by which a non-uniform distribution of gas could be present immediately inside or outside of an HVAC inlet are of particular interest in the present project. If an HVAC inlet ingests a non-uniform distribution of gas then there is the possibility that this could be missed by the detection system. To understand how it might be possible that a non-uniform distribution of gas could be present at an HVAC inlet, it is useful to briefly examine the characteristics of a gas jet. For simplicity we examine a free round jet, although in practice a release is likely to be more complex. Nevertheless, this simple case does give a broad indication of the flow behaviour which might be expected of some more complex releases. Far downstream from a high pressure release from a round hole, in the region where the Mach number is low enough for the flow to be treated as incompressible (Ma < 0.3), the concentration on the centre-line of the resulting jet is inversely proportional to distance (Rodi, 1982): T m D T o z where T m is the concentration on the centre-line at distance z from a release of diameter D with an initial uniform concentration T R. Far downstream, the concentration varies continuously across the radius of a round jet and can be described by the following function (Rodi, 1982): T T e m r z where T is the concentration at radius r. An example is helpful to indicate the radial variation in concentration which is implied by the expression immediately above. Consider a release of pure methane such that the centre-line concentration of the resulting jet is equal to LEL 10 m downstream from an initial diameter D (to be defined). Concentrations of 100%, 50%, 20% and 10% LEL would be obtained at radii of 0 m, 1.05 m, 1.6 m and 1.9 m. In fact a jet of these dimensions could be expected from a high pressure release of methane at a stagnation pressure and temperature of 100 bar and 40 o C, from a hole of approximately 12 mm diameter, respectively. This is a large, but credible, release. Note that the hole diameter of 12 mm is not D in the expression above. D is the diameter of the jet when it has expanded to atmospheric pressure, 85 mm in this instance - considerably larger than the release hole. The expressions by which this fully expanded diameter has been calculated are given in Section 5. This example shows that for a large, but credible, high pressure jet release the radial concentration at 10 m downstream can vary from a maximum of 100% LEL at the jet centre-line to just 10% LEL at a radius of 1.9 m. This distance, over which concentration varies by a factor of ten, is broadly comparable to the dimensions of typical offshore HVAC inlets. If such a release were ingested into an HVAC inlet, then significant non-uniformity in gas concentration could be expected outside and inside of the HVAC duct. 9

17 In practice a release could be affected by many factors such as the wind flow in and around a module, the presence of obstructions, impingement on nearby objects, etc. The release could also be at lower pressure and more significantly influenced by buoyancy. The effects of these factors, and others, are difficult or impossible to predict from theory alone, but in principle can be captured by CFD modelling. Therefore two scenarios have been simulated using ANSYS CFX 10 to illustrate and examine circumstances in which a non-uniform distribution of gas and air may be ingested into HVAC inlets. The scenarios are idealised but still include many of the complicating factors mentioned above. Scenario 1 is a high pressure release from a riser on the underside of a platform. The release is directed horizontally underneath the platform towards an HVAC duct which faces vertically down and whose inlet protrudes from the underside of the platform. Scenario 2 is a low pressure release from within a module. The release is directed upwards and out of the module to pass up the downstream face of a platform. A horizontal HVAC duct protrudes from this downstream face of the platform and is intersected by the release. Full details of these idealised scenarios, including a description of the CFD modelling and results, are given below 3.2 SCENARIO Model geometry and computational domain The model geometry is shown in Figures 2 and 3. Although it is a very simplified representation of an offshore platform this geometry presents approximately the same obstruction to the approaching wind-field as a real platform, but without the inclusion of any platform-specific geometrical elements such as flare-stacks, helipad, etc. The modelled geometry consists of a cube of side 30 m whose base is located 25 m above sea level. The release is located 3 m below the platform, 6 m inboard from the platform edges, and is initially angled at 30 o across the underside of the platform towards its downstream edge. The release is also oriented at a very shallow angle upwards such that its axis as it exits from the underside of the platform is approximately 2 m below the base of the platform. The location and initial direction of the release are indicated in Figures 2 and 3. A single HVAC duct of external dimensions 3 m x 1.5 m and internal dimensions 2.84 m x 1.34 m is located 4 m from the downstream face of the platform and 12 m from the nearest edge of the platform. The size of the duct was guided by information supplied by Integrated Engineering Services Ltd (specialist designer and supplier of offshore HVAC systems). The axis of the duct is vertical, 20.9 m long and protrudes 2 m below the underside of the platform. The HVAC duct can be seen in Figures 2 and 3. Figure 2: Scenario 1, geometry (HVAC duct in blue, gas release in red) 10

18 a) Plan view b) Vertical elevation Figure 3: Scenario 1, dimensions (m) An extensive region of the atmosphere that surrounds the platform is included in the CFD model. This is to ensure that the simulated flow around the platform is not constrained by the presence of the nearby boundaries of the computational domain. In particular it allows for the presence of large-scale time-dependent flow features to develop naturally as a consequence of flow over the bluff body of the platform. The resulting computational domain is 120 m wide, 11

19 115 m high, 240 m long and is illustrated in Figure 4. It includes the entire length of the HVAC duct. Note that the platform has been oriented at 20 o to the oncoming wind direction. Figure 4: Scenario 1, computational domain Boundary conditions The platform is located in the atmospheric boundary layer. Therefore at the upwind face of the computational domain a neutral stability atmospheric boundary layer profile was prescribed (Hargreaves & Wright, 2007). The wind speed was specified to be 1.5 m/s at 25 m above sea level. Although this is a low wind speed the results show that it is still sufficient to deflect the trajectory of the jet towards the HVAC duct. At the upper boundary of the computational domain a constant shear stress was applied by imposing a boundary velocity equal to that which would be expected at 115 m above sea level. At the downwind and two side boundaries in the atmosphere a condition was applied which allows for flow to leave or enter the domain. At sea level a rough wall boundary condition was applied. The sides, base and top of the platform were also specified as rough walls, with a surface roughness appropriate for corrugated metal (White, 1987). A mass flow rate of 19.2 kg/s was imposed at the upper end of the HVAC duct. This type of boundary condition allows the simulation to calculate an appropriate flow profile, rather than requiring a flow profile to be prescribed. The mass flow rate is equivalent to a uniform velocity of 4 m/s. Simulations have also been undertaken with a mass flow rate which is equivalent to a uniform velocity of 6 m/s and the results are qualitatively very similar. The release was assumed to be from a high pressure riser. Note that the riser is not explicitly included in the model since it will have negligible effect on the flow. In reality a gas release from a high pressure riser will initially be under-expanded and supersonic, i.e. it exhausts from the riser, accelerates to high Mach number then relaxes to atmospheric pressure by expansion through a complex shock structure. It is not practical or necessary to explicitly simulate this phase of a release. Instead, it is appropriate for such a release to be modelled from a location at 12

20 which expansion to atmospheric pressure has occurred and the flow is just sonic (Ivings et al, 2003). This has two advantages: it avoids the need to model the high Mach number region which can be problematic; and the dimension of the jet after expansion is much greater than that of the release hole and so is more readily resolved by the CFD mesh. In this scenario the release is modelled from the point at which expansion to atmospheric pressure has occurred. The release is sized and specified using the expressions in Section 3.1 for a free round jet as a rough guide. In this case the release is fixed at approximately 23 m from the HVAC duct. A peak concentration of 50% LEL was decided upon as a target concentration at this location since this is comparable to current action levels upon detection of gas at 50% to 60% LEL (Walsh et al, 2005). Using the expressions in Section 3.1 implies that after expansion to atmospheric pressure a jet which meets these target conditions should be specified as being 0.1 m in diameter. Note that for a high pressure release this actually corresponds to a much smaller hole size. For example, if the riser pressure and temperature were 100 bar and 40 o C, then the hole size would be approximately 14 mm diameter for a release of 100% methane. Section 5 provides expressions by which this effective hole size can be calculated. The release was therefore modelled as a sonic source of 100% methane over a nominal diameter of 0.1 m. The temperature of gas in the riser was assumed to be 40 o C which gives a sonic velocity of 428 m/s. The release rate is approximately 2.5 kg/s Flow physics The flow over a bluff body such as an offshore platform is known to be inherently transient. Large vortices are typically shed from a bluff body to form an oscillating wake. In fact initial attempts to compute this scenario as steady-state failed to converge. Therefore, all the simulations undertaken and presented in this report are time-dependent. The flow around the platform and in the HVAC duct will be turbulent. The commonly-used k-h turbulence model (Launder & Spalding, 1972) has been used throughout this work. Wall functions were used at all solid boundaries, appropriately modified to account for surface roughness. The physical properties for methane were taken from Santon et al (2002), including its LEL of 4.4% by volume Mesh An unstructured mesh was used to resolve the geometry. The mesh was refined in regions where high gradients of velocity and concentration were expected; across the methane jet, the HVAC inlet and duct. A total of 22 mesh elements were used to resolve the release source. On the walls of the platform four layers of prismatic mesh elements were used to aid resolution of the nearwall flow and help ensure that the first mesh node correctly falls within the log-law region for a turbulent boundary layer. A mesh with a total of 317,000 nodes (control volumes) was generated, constructed from 1.7 million mesh elements. The mesh is illustrated in Figures 5 to 7, overleaf. 13

21 Figure 5: Scenario 1, mesh - vertical elevation coincident with the release Figure 6: Scenario 1, mesh plan view coincident with the release 14

22 Figure 7: Scenario 1, mesh detail near the HVAC inlet Solution numerics and convergence All simulations in this work have been undertaken using high-order-accuracy spatial and temporal discretisation schemes in an effort to reduce numerical errors. Time-dependent simulations were undertaken with a short time-step of 1 s to aid convergence and also to help resolve the main time-dependent flow features. A non-dimensional frequency known as the Strouhal number, S, can be identified for the time-dependent flow over a sharpedged building: S fw /U where f is frequency, W is width in the cross-wind direction and U is mean velocity. For a cube of side 30 m in a 1.5 m/s flow the Strouhal number has a value of approximately 0.1 (Schetz and Fuhs, 1996), giving a frequency of Hz, equivalent to a time-period of 200 s. Hence, a time-step of 1 s is short in comparison to this long time-period. The transient simulation was run until the flow appeared periodic. A total of 1600 time-steps were computed. The residuals for all transport equations decreased by between at least two and usually three orders of magnitude, to very small rms values, and the global imbalances of mass, momentum, energy and gas mass fraction all reduced to very much less than 0.05% of reference values. This indicates that the solution is well-converged. 15

23 3.2.6 Results The flow-field around the platform is shown in Figure 8 for two time-steps separated by 60 s. Although the gross flow is similar consisting of massive flow separation in the wake of the platform the detail of the flow in the wake region is dissimilar. a) time t b) time t + 60 s Figure 8: Scenario 1, velocity field in plan view at mid-height of the platform 16

24 The flow-field in a vertical elevation is shown in Figure 9. Again, the separated region in the wake of the platform can be seen. Figure 9: Scenario 1, velocity field shown on a vertical plane The time-dependent wake flow has a minor effect on the trajectory of the gas jet, but sufficient to affect the concentration distribution at the HVAC inlet and inside the duct. An iso-surface of gas concentration at 50% of LEL is shown at two times in Figures 10 and 11. These times represent the two extreme states in the quasi-periodic time-dependent flow. a) time t b) time t + 60s Figure 10: Scenario 1, iso-surface at 50% LEL, plan view 17

25 a) time t b) time t + 60s Figure 11: Scenario 1, iso-surface at 50% LEL, iso-metric view The very small difference in the trajectory of the jet at these two time-steps can be seen to result in a more significant influence on the flow within the HVAC duct. This is especially evident in Figure 11, in which a more prominent tongue of gas can be seen in the duct at t + 60 s. The full extent of the gas jet is shown in Figure 12, below. Gas concentrations of 10% LEL and above are evident a significant way downstream from the platform. Figure 12: Scenario 1, gas concentration in % LEL, plan view 0.5 m below the HVAC inlet The distribution of gas at 0.5 m below the HVAC inlet, 1 m and 3 m inside the HVAC duct, is shown in Figure 13 at time t and t + 60 s. There is a large gradient in concentration across the duct in each case, with concentration differing by a factor of approximately five or six across the breadth of the duct. However, there is only a small difference in the concentration distribution between planes located immediately outside and inside of the duct. Although gas concentration is somewhat higher in the HVAC duct at time t + 60 s, the overall flow behaviour 18