Considering two-phase flow, homogenous flow assumes that the gas and liquid phases are flowing at the same velocity, while other models for two-phase flow, such as drift-flux assume a separated flow model with the phases allowed to flow at different velocities. Generally the vapor flow is faster in upward flow because of the density difference.

3.3.1 The homogenous equilibrium model

Homogenous Equilibrium Model (HEM) treats the two-phase flow as finely mixed and homogeneous in density and temperature with no difference in velocity between the gas and liquid phases.

A general expression for void fraction a, is given in (Feentra et al., 2000).

where pG and pL are the gas and liquid densities respectively and S is the velocity ratio of the gas and liquid phase (i. e. S = UG / UL ). The quality of the flow x is calculated from energy balance, which requires measurement of the mass flow rate, the temperature of the liquid entering the heater, the heater power, and the fluid temperature in the test section. The HEM void fraction aH is the simplest of the two-phase fluid modeling, whereby the gas and liquid phases are assumed to be well mixed and velocity ratio S in Equation 36 is assumed to be unity. The average two-phase fluid density p is determined by Equation 37.

P = apG + (1 — a)pL (37)

The HEM fluid density pH is determined using Equation 32 by substituting aH in place of a. The HEM pitch flow velocity VP is determined by

Vp = Gp / pH

Where Gp=Pitch mass flux

Radioactive waste should be disposed of in a way that guarantees its isolation from the biosphere. The release of potentially harmful substances — radionuclides — must be prevented or limited to levels that do not harm human health or the environment (IAEA, 1994). In this context, the issue of a proper siting process gains importance, especially in terms of selecting a site that has geological, hydrological, seismic, morphological and other characteristics that would not contribute to the release of radioactivity from a repository and subsequent exposure of the population. The site selection process is therefore a critical step in the overall site acquisition process. Countries are seeking their own ways on how to achieve these goals. As regard Slovenia it may be seen as a successful example concerning low and intermediate level waste (LILW) disposal. However, a strategy for the management of spent nuclear fuel and high level waste (HLW) is still under consideration.

The benefits of strategic environmental considerations in the process of siting a repository for LILW are clearly presented in Dermol & Kontic, 2011. The benefits have been explored by analyzing differences between the two site selection processes. One is a so-called official site selection process, which was implemented by the Agency for radwaste management (ARAO); the other is an optimization process suggested by experts working in the area of environmental impact assessment (EIA) and land use (spatial) planning. The criteria on which comparison of the results of the two site selection processes has been based are spatial organization, environmental impact, safety in terms of potential exposure of the population to radioactivity released from the repository, and feasibility of the repository from the technical, financial/economic and social points of view (the latter relates to consent by the local community for siting the repository). The site selection processes have been compared with the support of the multi-objective decision expert system named DEX — Decision EXpert (Bohanec & Rajkovic, 1999). The results of the comparison indicate that the sites selected by ARAO meet fewer suitability criteria than those identified by applying strategic environmental considerations in the framework of the optimization process. This result stands when taking into account spatial, environmental, safety and technical feasibility points of view. Acceptability of a site by a local community could not have been tested, since the formal site selection process had not yet been concluded at that time. Now the consent has been granted and ARAO is about to start construction of the repository in 2012. This approach to siting and comparison of the two site selection processes may serve as an example of transparent and inclusive — the local partnership has been established — way of dealing with radioactive waste disposal.

An analytical procedure for radiochemical characterization of radioactive waste material containing some of the radionuclides cited in Table 1 was developed. Radionuclides 242Pu, 238Pu, 239 + 240Pu, 241Am, 235U and 238U were determined by alpha spectrometry whilst 241Pu, 90Sr, 55Fe and 63Ni were determined by LSC and 59Ni by low energy gamma spectrometry. 242Pu, 238Pu, 243Am and 232U were used as tracers and Sr (2 mg/mL), Fe (3 mg/mL) and Ni (2 mg/mL) were used as carriers. In this work was developed a sensitive method for sequential analyses of the radionuclides in samples of radioactive waste. The samples analyzed were evaporator concentrate, resin and filter originated from Brazilian Nuclear Power Plants located at Angra dos Reis city (Reis et al, 2011).

The radiochemical procedure consists of three steps performed by anion-exchange chromatography, precipitation techniques and extraction chromatography, using TRU, Sr and Ni resins. In the first step, it was made the separation of 242Pu, 238Pu, 239 + 240Pu and 241Pu of the matrix by ion exchange chromatography using an anion exchange column (Dowex 1X8, Cl-form, 100-200 mesh, Sigma Chemical Co., USA). The separation is based on the formation of anionic complexes of Pu (IV) with NO3- or Cl — in concentrated HNO3 or HCl. In the second one, the effluent from the exchange column was used to separate Am and Sr by co-precipitation with oxalic acid of Fe, U and Ni that are retained in the filtrate. Americium and Sr isolation was done using commercially available resins, TRU resin and Sr Resin, respectively. These resins can be used for a number of analytical purposes, including the separation of actinides as a group from the matrix, separation of Sr from the matrix and sequential separation of individual actinides and Sr. In the third step Ni was separated by co-precipitation of Fe and U. And after that, Fe and U were separated by ion exchange chromatography using the anion exchange column (Dowex 1X8, Cl form, and 100-200 mesh) and Ni was isolated by Ni Resin extraction chromatography column from Eichrom Technologies, Inc. This work represents a fundamental step in establishing an analytical protocol for radioactive waste management system.

The safety planning for disposal of LLW and ILW radioactive waste takes account in special long half-life radionuclides. Both 59Ni and 63Ni are activation products of stable nickel, which was present as an impurity in fuel cladding materials or the uranium fuel of reactors (Kaye et al., 1994). 59Ni (half-life 7.6 x 104 years) is produced by neutron irradiation of 58Ni and decays by electron capture to stable 59Co with emission of 6.9 keV x-rays. 63Ni (half-life 100 years) emits only low-energy beta rays with a maximum energy of 67 keV, and is produced through neutron irradiation of 62Ni. Counting requirements dictated that prior the measurement these isotopes should be separated and purified with the purpose of removing the radiometric and chemical interferent elements so that they are essentially free of significant radioactive contamination.

Hou (Hou et al., 2005) proposed an analytical method for the determination of 63Ni and 55Fe in nuclear waste samples. Hydroxide precipitation was used to separate 63Ni and 55Fe from the interfering radionuclides as well as from each other. The separated 63Ni was further purified by extraction chromatography. According to him the recovery of Fe and Ni by hydroxide precipitation using NH4OH, was about 99, 9% and 21, 9%, respectively. Lee (Lee et al., 2007) proposed a sequential separation procedure developed for determination of 99Tc, 94Nb, 55Fe, 90Sr and 59/63Ni in various radioactive wastes. Ion exchange and extraction chromatography were adopted for the individual separation of the radionuclides. According to him Ni separation on the cation-exchange resin column was not selective enough therefore a further purification of Ni was performed by precipitation with dimetylglyoxime.

The aim of this work is the sequential analysis of nuclear waste containing several radionuclides (Pu, U, Am, Sr, Fe e Ni) where the last step consists in the separation of U, Fe and Ni. Thus we established the procedure for sequential separation of Pu, Am, Sr (Reis et al., 2011) in which we also included one step that is the hydroxide precipitation to separate U and Fe from Ni because Ni remains in solution in the co-precipitation of U and Fe.

2.3 Experimental

Pressure sharply decreases in DSs and RCS in the case of LOCA in Zone 4, but coolant circulation through the reactor core can be destroyed only in the latest stages of the accident, when pressure in RCS drops below 0.4 MPa. Thus, the characteristic fuel cladding and channel walls temperature peak within the first seconds of the accident (in case of LOCA in Zone 1 and Zone 2) is not met there. Guillotine breaks of one and two steamlines were considered for the large LOCA in RCS Zone 4: [3]

Many essential factors in selecting a measurement sensor for the RVI modularization were studied. Namely, the measurement environment, the measured object, size of the sensor, the weight of a sensor, measurement range, drive force of the sensor, accuracy and resolution were investigated

First, contact sensors and non-contact sensors were compared (Ko et al., 2009) (ABB-CE, 1995).

Contact sensors are used to directly measure distance by moving the sensors toward the measured object. Non-contact sensors are used to indirectly measure comparatively very small distances by laser, high frequency and eddy current(Figliola & Beasley, 2000)( Beckwith et al., 1993).

Contact sensors are more suitable than non-contact sensors. Non-contact sensors are not ideal because their outside diameter is bigger than the hole diameter 8 [mm] of the CSB snubber lug, the measurement range is smaller than it is with contact sensors, and, most importantly, design changes to the CSB snubber lug are bigger in non-contact sensors than in contact sensors.

The most important principle can be directly applied without design changes to the existing reactor internals. The contact sensor of a remote distance measurement sensor must be inserted into the measurement hole of CSB snubber lug. Therefore, an essential condition of a sensor must be a probe type because a measurement sensor must pass through the measurement hole of the CSB snubber lug. The outside diameter must be 8 [mm] or less since the measured diameter of the hole in the CSB snubber lug is 8 [mm]. The length of the probe head must be 117.6 [mm] or more because the end point of the measurement hole in the CSB snubber lug must reach the RV after a zero point adjustment. Also, a sensor must span the distance from the CSB snubber lug to the RV when the RV and CSB are assembled, thus requiring the backward probe head to be 147.3 [mm] or less. Measurement range can be measured from 50 [mm].

The previous measurement was done by hand-measurement and the measured maximum value was 48.92 [mm] (Uljin #5 nuclear power plant in Korea). The resolution must be 0.0254 [mm] or below as the sensor has to measure until 0.0254 [mm] (1/1000") in the case of gap measurement between the CSB snubber lug and the shim on the RV core-stabilizing lug, after the shim was assembled onto the CSB snubber lug by cap screws.

A sensor was investigated that remote measurement was possible in at least 25 [m] or more. The following items were considered when selecting a sensor: material, space and outward shape of the reactor internals. Also, no additional devices should be installed around the RV core stabilizing lug and the CSB snubber lug in order to perform the remote measurement.

Table 1 shows the suitable specifications of a sensor that have been researched to measure gap. The shape of a sensor is probe type of a contact sensor and the measurement method is digital.

Finally, the SOLARTRON (UK) sensor (DT/20/P) was selected for the reduced-scale model system (Solartron-metrology, 2006).

The MASLWR (Modro et al., 2003; Reyes et al. 2007, Mascari et al., 2011a; Mascari et al., 2011e), figure 1, is a small modular integral PWR of 35 MWe developed by Idaho National Engineering and Environmental Laboratory, OSU and Nexant-Bechtel. During steady state condition the primary fluid, in single phase natural circulation, removes the core power and transfers it to the secondary fluid through helical coil SG. In transient condition the core decay heat is removed through a passive primary/containment coupling mitigation strategy based on natural circulation. The use of natural circulation reduces the number of active components simplifying the configuration of nuclear steam supply system. The reactor core and a helical coil SG are both located within the RPV. The integrated SG consists of banks of vertical helical tubes located in the upper region of the vessel outside of the HL chimney.

Its small size considered the prototypical MASLWR relatively portable and thus well suited for employment in smaller electricity grids but take into account its design simplicity, its simplified parallel construction, the consequent reduction of the capital costs, reduction of construction time, reduction of finance and operation cost, recognizes it to be able to reach larger electricity market in developing and developed regions (Modro et al., 2003; Reyes & Lorenzini, 2010; Mascari et al., 2011d).

As it is shown in figure 1, the primary coolant flows outside the SG tubes, and the Feed Water (FW) is fully vaporized resulting in superheated steam at exit of the SG. The safety systems are designed to operate passively. The RPV is surrounded by a cylindrical containment, partially filled with water. This containment provides pressure suppression and liquid makeup capabilities and is submerged in a pool of water that acts as the ultimate heat sink. The MASLWR steady-state operating conditions are reported in the table 1.

The RPV, figure 1, can be depressurized using the Automatic Depressurization System (ADS), consisting of six valves discharging into various locations within the containment. In particular two independent vent valves (high ADS valves), two independent depressurization valves (middle ADS valves) and two independent sump recirculation valves are considered in the MASLWR design.

The integral arrangement of the plant allows avoiding pressurized primary components outside the RPV eliminating the possibility of large break Loss of Coolant Accident (LOCA) and reducing the Small Break LOCA (SBLOCA) initiating event. Of particular interest is the SBLOCA mitigation strategy typical of the MASLWR design. Following, for example, an inadvertent opening of an ADS valve, a primary side blowdown into the pressure suppression containment takes place. The RPV blowdown causes a primary pressure decrease and a consequent containment pressure increase causing a safety injection signal. It automatically opens, figure 1, the high ADS valves, the middle ADS valves and the sump recirculation valves. As the primary and the containment pressures become equalized, the blowdown is terminated, and a natural circulation flow path is established. Infact, when the sump recirculation valves are opened the vapor produced in the core goes in RPV upper part and through the high ADS valve goes to the containment where it is condensed. At this point through the sump recirculation lines and the down comer the fluid goes to the core again. The pressure suppression containment is submerged in a pool that acts as the ultimate heat sink. This mechanism, based on natural circulation, permits the cooling of the core (Modro et al., 2003; Reyes & King, 2003; Reyes, 2005b; Reyes et al., 2007; Mascari et al., 2011).

The MASLWR concept design and its passive safety features were tested in a previous test campaign developed at the OSU-MASLWR experimental facility (Modro et al., 2003; Reyes & King, 2003; Reyes et al., 2007; Mascari et al., 2011a; Mascari et al., 2011e), figure 2. The planned work related to the OSU-MASLWR test facility will be not only to specifically investigate the MASLWR concept design further but also advance the broad understanding of integral natural circulation reactor plants and accompanying passive safety features as well. Furthermore an IAEA International Collaborative Standard Problem (ICSP) on the "Integral PWR Design Natural Circulation Flow Stability and Thermo-Hydraulic Coupling of Containment and Primary System During Accidents" is being hosted at OSU and the experimental data will be collected at the OSU-MASLWR facility. The purpose of this IAEA ICSP is to provide experimental data on single/two-phase flow instability phenomena under natural circulation conditions and coupled containment/reactor vessel behavior in integral-type reactors (Woods & Mascari, 2009; Woods et al.; 2011).

The acquisition of realistic operational load data in the power plant is one essential pillar of the AFC. Its function is to determine the realistic thermal loads. FAMOS was developed in the early eighties. At that time, German licensing authorities demanded for the realization of a comprehensive measurement program in one German NPP. This was in order to get detailed information on the real component loadings during plant operation. This proof should give the information that the real operating conditions are not different from the design data. At that occasion, the advantages of monitoring real operating loads and using the measured data as an input for fatigue analyses became obvious. Therefore, a sophisticated fatigue monitoring system was developed. As a consequence, many NPPs in Germany and abroad were equipped with FAMOS (see also [6]).

Depending on each power plant, a fatigue handbook is developed to identify the locations relevant to fatigue in the NPP. The instrumentation of these locations is specific for each plant and depends on system design and further requirements.

For the acquisition of load data FAMOS uses two different methods: the global fatigue monitoring and the local fatigue monitoring [7]. The global monitoring is made by existing operational measurement. The corresponding operational signals could be fluid pressure, fluid temperature, the position of valves etc. measured at different parts of the systems.

Local fatigue monitoring is located at fatigue relevant locations at the outer surface of pipes and is based on additional temperature measurement by means of thermocouples. The thermocouples are manufactured as measurement sections.

Figure 2 shows the typical locations of measurement sections in a pressurized water reactor (PWR). Indeed, FAMOS gathers measurement sections, which are mostly located on the:

• primary loops

• surge line

• spray lines

• volume control system

• feedwater system and further positions.

The different FAMOS measurement sections can be composed of seven or more thermocouples if some thermal events like stratification are suspected (horizontally installed pipes). However, in case of plug flow the application of only two thermocouples is sufficient (vertically installed pipes).

Thermocouple

Measuring tape

Protection shell

Fig. 3. FAMOS principle

Figure 3 shows how the application of thermocouples at the outer surface of a pipe is performed. More details on the technical bases of FAMOS can be found in [8].

Each measurement section consists of the thermocouples installed on a thin metal tape and a robust protection shell to prevent the thermocouples from being damaged. Both, metal tape and protection shell are installed around the pipe under the piping insulation. The installation takes place at a certain distance to pipe welds. Thus, the dismounting of measurement sections during an ultrasonic testing of the weld is unnecessary. Furthermore, a distance to thick walled components (e. g. nozzles) is needed in order to minimize the thermal impact on the temperature measurement.

The measurement sections are designed for a fast mounting process. A very short installation time is absolutely necessary to ensure low dose rates for the mounting staff. Special manufacturing processes and thermal responses tests of the thermocouples guarantee realistic thermal load data. Actual measurement sections are characterized by a minimization of heat capacity effects and excellent thermal sensitivity.

All thermocouples are wired with extension lines to junction boxes where the cold junction compensation of the thermo voltage signal is performed. For the compensation a board with an isothermal terminal for up to 30 channels is used to connect the wires made of thermocouple material with a trunk cable made of copper. The temperature at the isothermal terminal is measured with a resistance thermometer. The voltage signals from different junction boxes are connected to the information and control system (I&C) of FAMOS. That system consists of two or more information modules with signal processing units and analog — to-digital converters. The information modules are connected by means of a data bus to a processing unit. The data of all thermocouples are analyzed, stored and transferred to a computer network in real time. To avoid a high data volume in the storage system a data reduction method is used for the samples of the signals. By application of the bandwidth method samples will only be stored if signals leave the bandwidth. Thus, it is ensured that all samples of fatigue relevant load data during transients are stored and only a few samples during steady state operating mode of the plant with no fatigue relevant transients.

The information modules and the processing unit are installed within the containment of the power plant. All requirements of the containment environment such as temperature, humidity, atmosphere and material restrictions have been considered during engineering and design of the I&C components. The system is equipped with redundant power supply and a diagnostic system to detect malfunctions in different modules. Thus, a long term monitoring and storage of load data in a containment environment is ensured.

Outside the containment a sever unit is installed in the safeguards building, main control side room or other computer rooms of the power plant. That server stores all load data in a database over years. In addition to the FAMOS data, load data of the global instrumentation are also stored in the sever unit. For the connection of the server unit with the processing unit inside the containment fibre optic cables or normal ethernet connections made of copper are used. These cables run in special penetrations through the pressure-sealed and gas tight containment wall.

By means of a network connection to the server unit all data of FAMOS and the global monitoring can be viewed in real time on any connected computer. The access to the database
features the view on the load data of several years of power plant operation. The real time data analyzing process done by the processing unit generates messages about high thermal loads or malfunctions automatically. Thresholds for temperature rate of change or thermal stratification can be set. Is the threshold reached during a transient the system generates a warning about high thermal loads. In combination with valve position signals from the global monitoring, the system detects insufficiently closed valves even if the valve position indicates it.

With the help of automated warnings, unfavourable thermal load events can be analyzed and operating modes optimized. For the purpose of preventive maintenance FAMOS offers all necessary information to prevent the plant’s components from avoidable load events.

The FAMOS I&C system operates in real time with signals of 120 or more than 300 thermocouple channels. The number of channels depends on the power plant layout, the number of primary loops and the monitored components. The engineering of FAMOS in a NPP starts with the generation of a FAMOS manual. This process contains a deep analysis to identify components relevant to fatigue in the primary, secondary, auxiliary and safeguards systems. To do so, design documents, operating experience and feedback from similar plants are considered. A measurement point plan is elaborated and all activities are coordinated with the plant operator and, if required, with independent experts.

Fig. 4. Development of fatigue usage factor considering local monitoring and improvements

If FAMOS is installed before commissioning of the plant the fatigue status can be calculated over the complete lifetime with realistic load data. Indeed, the commissioning phase is often characterized by the highest loads of the entire lifetime of the power plant. But an installation into a running plant is useful after several years of operation e. g. during lifetime extension.

Getting the real measurements at this stage implies a consequent reduction of the fatigue usage factor when the later detailed fatigue calculation is required as explained in Figure 4.

The objectives of FAMOS are summarized here below:

• to determine the fatigue status of the most highly stressed components

• to identify and optimize the operating modes which are unfavourable to fatigue e. g. valve leakages

• to improve the catalogue of transients used at the design phase

• to establish a basis for fatigue analysis based on realistic operating loads

• to use the results for lifetime management and lifetime extension.

An important connecting link between the recording of measurements and the stress and fatigue analysis is the load data evaluation (see Figure 1) and the specification of thermal loads in transient catalogues. Expert knowledge of the processes and the operation of the systems is essential. The first step to specify thermal loads is the identification of the operational processes leading to relevant transients of temperature and/or pressure. For these events an appropriate number of model transients is selected. Usually, it is necessary to split up these model transients into subclasses which are different e. g. in the temperature difference of the transient. As it is shown in Figure 1 the specified transients are input data for the detailed fatigue check.

Indeed, three graded methods were developed fulfilling the different requests in terms of fatigue. The choice of one method depends on the expected degree of fatigue relevance and the expected grade of details in fatigue calculation.

Step 1: Simplified fatigue estimation

Simple estimations of fatigue relevance of real loads for components are based on thermal mechanical considerations using the equation of completely restrained thermal expansion. A basic decision about fatigue relevance (yes/no) for the monitored position is made. In case of fatigue relevance a further evaluation is done according to step 2.

Step 2: Fast fatigue evaluation

A code-conforming usage factor U is calculated in a highly automated way based on the simplified elasto-plastic fatigue analysis. If U < Uadmissible the fatigue check will be success­fully finished. If U > Uadmissible further analyses will be based on step 3.

Step 3: Detailed fatigue calculation

Fatigue analysis is based on a detailed catalogue of transients. This catalogue of transients results from the evaluation of the real loads for the monitored component. The detailed fatigue check is based on finite element analyses mostly including elasto-plastic material behavior.

Both step 2 and step 3 allow for the consideration of EAF in the analysis process.

The developed approach for reduction of emergency planning zones (EPZ) was summarized in related IAEA document (IAEA-TECDOC-1652, 2010). It is applicable and recommended for all types of SMRs without on-site refuelling. The spatial extents of regulatory-mandated EPZ have historically been set according to conservative approaches for calculating bounding individual dose rates subsequent to a postulated accident sequence. The zones are not small — ranging up to 10 kilometers or even miles in radius. Moreover, regulations often require the reactor owner to provide for emplacement of infrastructure such as roads and bridges throughout the EPZ to facilitate public evacuation — as well as to periodic training and equipment supply to first responders. Current practice has been developed over many years specifically for the historical and current situation of large water-cooled reactor installations generating electricity for a regional grid.

Alternately, SMRs without on-site refuelling are being designed for local grids and some are even designed for cogeneration missions wherein the reactor must of necessity be placed very near the cogeneration application due to short heat transport distances. EPZ defined for large reactors on a one-size-fits-all basis can place a severe economic disadvantage on SMRs without on-site refuelling. For this reason, the IAEA CRP has conducted a review of the basis for the current regulations and has proposed a risk-informed methodology which could justify a reduced emergency planning zone extent on the basis of a smaller source term and a reduced probability of release for advanced SMRs, accounting for their passive safety and other risk reduction features. The methodology is not limited to small reactors without on-site refuelling, but is unique to many NPPs with innovative SMRs and larger reactors.

Within this methodology the information gathered from the PRA (both internal and external events) may be used to provide a basis for the redefinition of the EPZ defining criteria. The proposed approach consists of coupling the PRA results with deterministic dose evaluations associated to each relevant PRA sequence considered, and thus achieving a technically sound bases for the definition of a plant specific EPZ. In this approach the two basic components of risk (i. e. probability of occurrence and consequences of a given accident) are therefore explicitly combined. The EPZ radius then is defined as the distance from the plant such that the probability of exceeding the dose limit triggering the actuation of emergency procedure is equal to a specified threshold value. To identify this threshold value, detailed analysis of existing installations should be performed to infer the risk associated with the current EPZ definition.

The study conducted in the CRP included a sample application of the developed methodology for the IRIS-like SMR design under conditions of a particular site. This application indicated a potential for remarkable reduction of EPZ radius without increase in the public risk. However, to achieve this practically the proposed methodology first needs to be embraced by regulatory authorities. More details of the methodology and its trial application are provided in related IAEA document (IAEA-TECDOC-1652, 2010).

It must be noticed that the use of existing regulations and installations as the basis for this redefinition will not in any way impact the high degree of conservativism inherent in current regulations. Moreover, the remapping process makes this methodology partially independent from the uncertainties still affecting probabilistic techniques. Notwithstanding these considerations, it is still expected that applying this methodology to advanced plant designs with improved safety features will allow significant reductions in the emergency planning requirements, and specifically the size of the EPZ. In particular, in the case of IRIS it was expected that taking full credit of the Safety-by-Design™ approach of the IRIS reactor will allow a dramatic reduction in the EPZ requirement, while still maintaining a level of protection to the public fully consistent with existing regulations.

A summary of some recent efforts on the analysis of fluid-elastic instability in heat exchanger and steam generator tube bundles is given in Table 4.

1.2

Vorticity induced instability

Flow across a tube produces a series of vortices in the downstream wake formed as the flow separates alternatively from the opposite sides of the tube. This shedding of vortices produces alternating forces, which occur more frequently as the velocity of flow increases. For a single cylinder, frequency of vortex shedding fvs is given below by a dimensionless Strouhal number S.

_ SV vs _ D

where V is the flow velocity and D is the tube diameter. For a single cylinder, the vortex shedding Strouhal number is a constant with a value of about 0.2 (Chenoweth, 1993). Vortex shedding occurs for the range of Reynolds number 100 < Re <5×105 and > 2 x 106 whereas it dies out in-between. The gap is due to a shift of the flow separation point in vortices in the intermediate transcritical Reynolds number range. Vortex shedding can excite tube vibration when it matches with the natural frequency of the tubes. For tube banks with vortex shedding, Strouhal number is not constant, but varies with the arrangement and spacing of tubes, typical values for in-line and staggered tube bundle geometries are given in (Karaman, 1912, Lienhard, 1966). Strouhal numbers for in-line tube banks are given in Figure 6.

The vortex shedding frequency can become locked-in to the natural frequency of a vibrating tube even when flow velocity is increased (Blevins, 1977). Earlier on, the mechanism of vortex shedding has been investigated by a number of researchers. These include Sipvack (Sipvack, 1946) and, Thomas and Kraus (Thomas & Kraus, 1964) who investigated the vortex shedding of two cylinders arranged parallel and perpendicular to flow direction respectively. Grotz and Arnold (Groth & Arnold, 1956) measured for the first time systematically the vortex shedding frequencies in in-line tube bank for various tube spacing ratios.

The cause of vorticity excitation has been disputed in literature (Owen, 1965), but recent studies of (Weaver, 1993) and, (Oengoren & Ziada, 1993) have confirmed its cause of existence as periodic vortex formation. Vorticity shedding can cause tube resonance in liquid flow or acoustic resonance of the tube bundles or acoustic resonance of the tube bundles’ containers in gas flows (Oengoren & Ziada, 1995). (Chen, 1990), (Zaida & Oengoren, 1992) and (Weaver, 1993) have summarized the recent research efforts targeted at improvement in Strouhal number charts for vortex shedding and acoustic resonance for in­line tube bundles.

L

D

Fig. 6. Strouha! numbers for in-line tube banks (Karaman, 1912).

(Oengoren and Ziada, 1992) have investigated the coupling between the acoustic mode and vortex shedding, which may occur near the condition of frequency coincidence. They have investigated the system response both in the absence and in the presence of a splitter plate, installed at the mid-height of the bundle to double the acoustic resonance frequencies and therefore double the Reynolds number at which frequency coincidence occurs. They have also investigated the effect of row number on vortex shedding and have carried out flow visualization in Reynolds number range of < 355000. Figure 7 is a typical example of the mechanism of vortex shedding from the tubes of the first two rows displaying a time series of symmetric and anti-symmetric patterns (Oengoren & Ziada, 1993).

(Liang et al., 2009) has addressed numerically the effect of tube spacing on vortex shedding characteristics of laminar flow past an inline tube arrays. The study employs a six row in­line tube bank for eight pitch to diameter (^/^) ratios with Navier-Strokes continuity equation based unstructured code (validated for the case of flow past two tandem cylinders) (Axisa & Izquierdo, 1992) . A critical spacing range between 3.0 and 3.6 is identified at which mean drag as well as rms lift and drag coefficients for last three cylinders attain maximum values. Also at critical spacing, there is 180o phase difference in the shedding cycle between successive cylinders and the vortices travel a distance twice the tube spacing within one period of shedding.

(Williamson & Govardhan, 2008) have reviewed and summarized the fundamental results and discoveries related to vortex induced vibrations with particular emphasis to vortex dynamics and energy transfer which give rise to the mode of vibrations. The importance of mass and damping and the concept of "critical mass", "effective elasticity" and the relationship between force and vorticity. With reference to critical mass, it is concluded that

as the structural mass decreases, so the regime of velocity (non-dimensional) over which there is large amplitude of vibrations increases. The synchronizing regime become infinitely wide not simply when mass become zero but when a mass falls below special critical value when the numerical value depends upon the vibrating body shape.

(Williamson & Govardhan, 2000) present a large data set for the low branch frequency flower plotted versus m* (mass ratio) yielding a good collapse of data on to single curve base equation 7.

(7)

This equation provides a practical and simple means to calculate the frequency attained by vortex induced vibrations. The critical mass ratio is given by

m*crit = 0.54 ±0.002 (8)

Below which the lower branch of response can never be attained. With respect to combine mass-damping parameter’s capability to reasonably collapse peak amplitude data in Griffins plot, a number of parameters like stability parameter, Scrutom number and combined response parameter termed as Skop-Griffins parameter given by (SG):

Where S stands for single vortices and Sc is the Scruton number.

(Hamakawa & Matsue, 2008) focused on relation between vortex shedding and acoustic resonance in a model (boiler plant) for tube banks to clarify the interactive characteristics of vortex shedding and acoustic resonance. Periodic velocity fluctuation due to vortex shedding was noticed inside the tube banks at the Reynolds number (1100-10000) without acoustic resonance and natural vortex shedding frequency of low gap velocities. Kumar et al., 2008 in their review stated that controlling or suppressing vortex induced vibrations is of importance in practical applications where active or passive control could be applied.

(Paidoussis, 2006) specially addressed real life experiences in vortex induced vibrations and concludes with this mechanism in addition of other already clarified mechanisms of flow induced vibrations. Vortex induced vibrations of ICI nozzles and guide tubes in PWR for ICI thimble guiding into the core of the reactor to monitor reactivity may witness breakage of ICI nozzles resulting in strange noises experience in the reactor. Analysis of shedding frequencies confirmed the vortex induced vibrations to be the culprit partially due to the large values of varying lift coefficients and partially due to lock-in.

(Hamakawa & Fukano, 2006) also focused vortex shedding in relation with the acoustic resonance in staggered tube banks and observe three Strouhal number (0.29, 0.22 and 0.19). In cases with no resonance inside tube banks, the last rows of tube banks and in both regimes respectively. The vortices of 0.29 and 0.22 components alternatively irregularly originated.

(Pettigrew & Taylor, 2003) discussed and overviewed procedures and recommended design guidelines for periodic wave shedding in addition to other flow induced vibration considerations for shell and tube heat exchangers. It concludes that the fluctuating forces due to periodic wave shedding depends on the number of considerations like geometric configuration of tube bundles, its location, Reynolds number, turbulence, density of fluid and pitch to diameter ratio.

1. The channel box of the internal CSB and the air hose, compressor, electric power cord, power supply and signal cables of the external CSB were connected. Electric power was then supplied.

2. Using a software program running on a remote measurement computer, the length was measured a total of five times. The pressure for the measurement was adjusted to 0.8 ~ 1 bar; this was set to have a zero length to ensure that the data were entered correctly. The average value of the measured lengths was used as the data. Once the measurement was completed, the data were stored and the measured length was recorded.

3. After the gap measurements were completed, all electric power was turned off. After the RV and the CSB assembly were detached, the air hose, electric power cord and signal cable were respectively separated from the compressor, electric power, USB hub and RS-485 converter. Once separated, the air hose, electric power cord and signal cable were temporarily fixed in the CSB assembly to ensure that they would not interfere with the disassembly of the CSB assembly.