Monetary Policy Implications of Financial Frictions in the Czech Republic

As the global economy seems to be recovering from the 2009 financial crisis, we find it desirable to look back and analyze the Czech economy ex post. We work with a Swedish New Keynesian model of a small open economy which embeds financial frictions in light of the financial accelerator literature. Without explicitly modeling the banking sector, this model serves as a tool for understanding how a negative financial shock may spread to the real economy and how monetary policy may react. We use Bayesian techniques to estimate the model parameters to adjust the model structure closer to the evidence stemming from Czech data. Our attention focuses on a set of experiments in which we generate ex post forecasts of the economy prior to the 2009 crisis and illustrate that the monetary policy response to an upcoming crisis implied by the model with financial frictions is stronger on account of an increasing interest rate spread.Bayesian methods, financial frictions.

Macroparticle Charging in a Pulsed Vacuum Arc Thruster Discharge

Peer Reviewedhttp://deepblue.lib.umich.edu/bitstream/2027.42/77262/1/AIAA-2006-4499-411.pd

Microvacuum Arc Thruster Design for a Cubesat Class Satellite

This paper describes the University of Illinois 2-cube CubeSat (10 x 10 x 20 cm) designed for April 2003 launch. The Illinois Observing NanoSatellite (ION) includes a scientific mission to view the airglow layer of the atmosphere and a CMOS camera for space and Earth photography. ION will also be used as a test bed to demonstrate a number of technologies including an active 3-axis attitude control system, with a new propulsion system used for both attitude control as well as orbital maneuvers. The new vacuum arc thruster (VAT) propulsion system produces ion velocities of up to 30,000 m/s, driven mostly by local pressure gradients. A 12 V inductive energy storage circuit is used to provide the initial breakdown and to sustain the plasma. Four thruster heads can be controlled individually to produce arc pulses with adjustable pulse width and repetition rate. Size and mass have been driven by the CubeSat requirements and amount to 4 x 4 x 4 cm and 150 g, respectively. Thrust to power ratios are expected to be ≈10μN/W. The individual impulse will be close to 1μN-s/pulse. Challenges to the design and integration of the VAT into a CubeSat size satellite are presented. On board diagnostics and methods used to verify operation of the VAT are discussed

Performance and Heat Loss of a Coaxial Teflon Pulsed Plasma Thruster *

A coaxial Teflon pulsed plasma thruster, UIUC PPT-7, was tested and results were reported in a previous paper. More thrust data is taken while varying the stored energy for different geometries. A thermal model is used to determine energy lost as heat from the thruster to be 14% of the available energy. A similar thermal model is used to estimate the portion of heat lost due to conduction into the Teflon fuel. The plasma current is curve-fit to reveal a linearly increasing plasma resistance. The plasma temperature is estimated to vary from 9000 K to 21000 K. Introduction The pulsed plasma thruster is a robust, solid state device that has flown on a number of missions. In a previous paper 2 the performance of UIUC PPT-7, an electrothermal, coaxial pulsed plasma thruster was reported. The Two-Stream model was used to describe the thruster efficiency. Since then, more tests have been performed, resulting in a greater understanding of the performance versus energy of the thruster. Also, temperature data is examined to help discuss the heat loss of the thruster. Experiments Apparatus Experiments were performed to help determine the effects of geometry and energy on the performance of an electrothermal thruster. The thruster Tests with stored energy less than 50 J use 14 parallel mica capacitor sections, 3 with a total capacitance of 9.2 µF. In order to test at 70 J, 7 more sections are added, increasing the capacitance to 14.1 µF. The 50 J baseline current, shown in where ∫ = Ψ dt I 2 and E pl is the energy in the plasma. However, when this resistance is used to curve-fit to the experimental data, it becomes apparent that the resistance is not constant throughout the pulse. The plasma current is best fit with a linearly increasing resistance, as shown in A number of diagnostics were used during these tests. The UIUC compact thrust stand 4 was used in single pulse mode to measure the thrust. At least 10 thrust measurements were taken for each geometry tested. The current was measured using a Rogowski coil on the central electrode. A 1000:1 high voltage probe (Tektronix P6015) was used to measure the capacitor discharge voltage at the vacuum tank feed-through. Temperature measurements of the propellant and one capacitor pack were taken using type K thermocouples. The location of the thermocouples is shown on The thruster and capacitor surface temperature measurements were used to determine the heat loss from the thruster and the transfer loss. Starting from room temperature with a constant pulse rate, the capacitor temperature was observed to rise linearly at the outer surface. Assuming that the temperature is uniform throughout the capacitor the power lost to heating is calculated from the measured constant temperature rise rate and the thermal mass (mC p ), according to The thermal mass of the capacitors, listed in Because the heat from the Teflon cavity is flowing radially through the thruster body, the temperature is not uniform throughout. For this reason, the above equation cannot be used to determine the power lost as heat in the thruster. A heat flow model described later is employed to match the temperature rise rate at the outer surface of the fuel to the power lost to heating. Tests Each geometry tested is characterized by the cavity diameter near the rear electrode, the cavity length, and the cavity exit diameter. The baseline constant diameter geometry for these experiments is 14/35/14, denoting a 14 mm rear cavity diameter, 35 mm length, and a 14 mm front cavity diameter. From this baseline case, three other diameters and two other lengths were tested. Tests were also performed without a nozzle in order to determine its effect on performance. Two lengths of tapered cavities were also tested: 3.5/20/14 and 3.5/35/14 with 15° and 10° half angle respectively. These test parameters are illustrated in Test procedure Each test consisted of 1000 shots repetitive firing at 1 Hz to warm up the thruster and burn-in the fuel tube. After burn-in, 10 thrust measurements were taken in single pulse mode, followed by another 1000 shots to reduce error in the mass loss measurement. Each pulse produces a thrust stand position transducer (LVDT) output which is a slowly decaying sinusoid. 3 The initial (t = 0) slope is determined by curve-fitting an analytical damped sinusoid to the LVDT waveform, which gives the post-pulse thrust stand platform velocity u p . The impulse bit is u p multiplied by the platform mass, which is weighed for each test setup. Thermocouple data, including the capacitor and fuel temperature, were recorded during all repetitive pulsing phases of the test. The second set of data, measuring the variation of thrust versus energy at 8 mm diameter, were performed with a slightly different procedure. After the pumps reached operating pressure, the thruster was fired 1000 times at 50 J, 1 Hz to warm up the fuel and thruster. Then, approximately 10 single shot thrust measurements were taken at energies ranging from 50 J to 10 J. If the thruster fired repeatably at lower energies, thrust was also measured at energies below 10 J. Results The following plots represent the data taken. The specific mass loss (ML) vs energy is then curvefit using a cubic polynomial fit. The specific impulse vs energy curve fit is then calculated using the specific thrust and specific mass loss fits acording to the equation I sp = T sp /(g o ML). Specific thrust and specific mass loss vs diameter and length is also curve-fit using the above polynomial fit. The corresponding specific impulse curve fit is calculated as above. Some of the polynomial curve fits were determined without consideration of possible data outside the range of the data points. For this reason, extrapolation using these curve fits may be erronious. Analysis The thruster efficiency η t can be expressed as the product of component sub-efficiencies. 8 These efficiencies include the pulse energy transfer η tr , thruster heat loss η h , frozen flow η f , exhaust beam divergence η div , and exhaust velocity distribution efficiency η dist . The transfer efficiency takes into account the energy lost to the equivalent series resistance (ESR) of the capacitor. This energy is estimated from the temperature increase during a 50 W test, and the thermal mass of the capacitor. A small portion of the heating power is due to heat flow from the thruster, however that amount is insignificant compared to the thruster heating. For the baseline case, Q = 2.5 W This is 5.0% of the 50 W making the η tr = 0.95. Another way to estimate the transfer efficiency is from the relation capacitor heating/pulse = Ψ × ESR . A separately-measured capacitor ESR 3 of 2.8 mΩ for the baseline case gives 4.5 W, making η tr = 0.91. We adopt a value of η tr = 0.93. The energy in the Plasma E pl is therefore the stored capacitor energy reduced by capacitor and wall/electrode heat losses, or E pl = η tr η h E o . Mass loss. The pulse specific mass loss (µg/J) is relatively constant with energy for the baseline (14/35/14) geometry. For constant energy (50 J), specific mass loss is seen to increase for decreasing 4 diameters The specific thrust of the baseline geometry is relatively constant with energy at 35 -38 µN-s/J from 20 -70 J, and then decreases significantly (Eq. 3) at lower energies. For the 8 mm constant diameter cases, specific thrust increases slightly with energy from 20 -50 J. Below 20 J, the specific thrust decreases similarly to the baseline case. This behavior is consistent with a heat loss model of ∆Q loss = B + AE o . Assuming that the specific thrust scales as T sp = C E pl /E o , it can be shown that: which is mathematically of the same form as the curve fit of Eq. 3. By comparing Eq. 5 to the curve fit Eq. 3, values of the constant can be extracted. For the baseline geometry, the constant C is 46 µN-s/J, A = 0.06, and B = 2.8 J, so that T sp = 46(0.87 -2.8/E o ) and the heat loss is ∆Q loss = 2.8 + 0.06E o . At E o = 50 J, ∆Q loss = 5.8 J, approximately 50% from the constant term B and 50% from the term proportional to discharge energy. We have speculated as to the origin of the constant term of 2.8 J. One possibility is that the plasma pulse quickly raises the Teflon surface temperature to the vaporization point, where it is capped by the sublimation process, freezing the temperature profile and therefore the heat conducted into the solid. This picture suggests that shorter pulse lengths would reduce the constant term in the ∆Q loss equation. Applying this model to the specific thrust curve fit of Specific Impulse. Specific impulse generally increases with energy, from 330 s at 10 J to 490 s at 70 J (baseline geometry). With diameter, I sp reaches a maximum of 450 s at 14 mm (baseline). I sp increases to above 500 s for lengths < 25 mm, and some benefits (610 s) is shown for a short tapered cavity. Heat loss In this paper, we are mostly concerned with the heat loss efficiency, the other efficiencies are described elsewhere. where T is the temperature, r is the radius, r i is the inner radius, r o is the outer radius, and k 1 , k 2 , k 3 , kp 1 , kp 2 , kp 3 , are all constants. The model begins with temperature of 0° everywhere. Since all the calculations are based on temperature differences, starting from room temperature does not affect the outcome. The boundary conditions required involve a constant influx of energy at the inner radius, and zero efflux of energy at the outer radius. Due to a numerical problem, it is not possible to make both the inner and outer radius boundary conditions contain only derivative terms with k 3 or kp 3 equal to zero, so a small temperature dependence is added to the efflux of energy at the outer radius. The empirically determined constants for the baseline case are listed in where k is the solid Teflon thermal conductivity, A is the surface area of the inside of the tube, and ∂T ∂r = k 3 . Surface Heat Conduction From the above model, it was determined that 6.4 Joules of energy is lost to thruster heating for every 50 J pulse. A portion of that heat is deposited into the electrode through sheath losses. The rest is deposited into the Teflon. As energy is deposited into the Teflon fuel, three processes are occurring. Heat is conducted into the fuel, heat is radiated out of the fuel, and heat is lost when fuel evaporates. This paper considers the convection of heat into the solid Teflon from the high temperature plasma. The timedependent temperature distribution inside the Teflon immediately after one pulse is obtained by using a model similar to the one used in the heat transfer analysis, with constant initial temperature throughout the Teflon except at the surface, according to the following equation. . the constants a and b are the inner and outer radius of the Teflon fuel, v 1 and v 2 are the surface temperatures at the inner and outer surface respectively. α n is the nth root of where J 0 and Y 0 are Bessel functions. U 0 is the relation and κ is the diffusivity which is K/ρCp or 7.8x10 -8 for Teflon. 6 The model requires knowing the surface temperature of the Teflon. It is assumed that the pressure in the cavity is the same as the vapor pressure of the Teflon. The vapor pressure of Teflon follows the equation: where p c = 1.84 x 10 15 and T c = 20800 K. Eq. 11 is a curve fit of data from reference 10, using the ClausiusClapeyron equation to fit the data. 1 The fit and data are shown in To estimate the length of time that the Teflon surface temperature remains at 675°C the sonic velocity of the plasma during the current pulse is determined by estimating plasma temperature. The plasma resistance, determined by subtracting the ESR from the circuit resistance used to curve-fit the current data lnΛ is determined according to: ln Λ = 23.4 -1.15 Log(n e ) + 3.45 Log(T ev ) The electron density n e (in cm -3 ) needed for lnΛ is determined using the Saha equation for a gas mixture, assuming only single ionization 13 : The partition function ratios are found from spectral data 14 and are 0.53 for carbon and 1.68 for fluorine. The following equations are then used to solve for the electron density n e p = (n e + n C + n C + + n F + n F + )kT (15) A rough estimate of conductive and radiative heating rate of the Teflon surface exposed to this temperature plasma at several atm predicts that the temperature rises to 675°C in a short time compared to the pulse length. Using the sonic velocity determined above, the length of time it would take an expansion wave to travel from the cavity entrance to the rear electrode is approximately 8 µs. For this model, since parts of the Teflon surface would be cooled before the expansion wave would reach the end of the cavity, it is assumed that the Teflon surface no longer absorbs heat from the plasma 4 µs after the current pulse. The measured heat loss of 6.4 J/pulse can be accounted for from two sources: a) heat stored in a thin surface layer of Teflon, and b) heat transferred to the electrodes by the voltage sheaths during the pulse. By using the thermal conduction model, a temperature distribution extending approximately 5 µm inside the Teflon is developed at the end of the pulse. From Ref 10 the specific heat of Teflon includes a phase change from crystalline to amorphous, corresponding to 59 J/g at 600 K. Using a constant C p of 1.4 J/kg-K the temperature distribution corresponds to 4.2 Joules of energy stored in the Teflon in the form of heat. This value is increased by 10% to account for the phase change to a final value of 4.6 Joules. This accounts for 72% of the energy lost due to thruster heating. The remaining heat (1.8 J) can be accounted for by the sheath loss at the electrodes, assuming ~20 V total drop. Conclusions The specific thrust vs energy for geometries with a smaller diameter cavity than the baseline show a specific Thrust variation of T sp =a-b'/E o , similar to the baseline case. A thermal model was used to show the 7 transfer efficiency η tr =0.93 and the heat loss efficiency η h = 0.86, corresponding to 6.4 J lost to the wall out of 50 J stored. The plasma current was curve-fit to reveal a linearly increasing plasma resistance from which the plasma temperature is estimated to vary from 9000 K to 21000 K. A thermal model is used to estimate that 72% of heat lost is by convection into the Teflon fuel and the remainder is due to the sheath drop

Quantitative detection of four pome fruit viruses in apple trees throughout the year

A one-step real-time RT-PCR assay (RT-qPCR) with melting curve analysis, using the green fluorescence dye SYBR Green I, was developed to detect and quantify RNA targets from Apple mosaic virus (ApMV), Apple stem grooving virus (ASGV), Apple stem pitting virus (ASPV) and Apple chlorotic leaf spot virus (ACLSV) in infected apple trees. Single PCR products of 87 bp (ApMV), 70 bp (ASGV), 104 bp (ASPV) and 148 bp (ACLSV) were obtained, and melting curve analyses revealed distinct melting temperature peaks for each virus. A dilution series using in vitro synthesized transcripts containing the target sequences as standards yielded a reproducible quantitative assay, with a wide dynamic range of detection and low coefficients of variance. The content of selected viruses in apple plant tissues was stable throughout the year, and their accumulation did not significantly change between different plant tissues. The only minor exceptions were for ApMV and ACLSV, in which noticeable differences in their concentrations in various biological material were observed within the year. This divergence did not influence their year-round detectability. This one-step RT-qPCR assay is a valuable tool for year-round diagnostics, and molecular studies of the biology of ApMV, ASGV, ASPV and ACLSV

Condensation-free radiant cooling using infrared-transparent enclosures of chilled panels

Radiant cooling power in the humid climates is inherently limited by condensation. This research investigates a type of radiant cooling methodology whereby the cold temperature source is convectively and conductively isolated from the environment with a membrane transparent to visible radiation to allow supply temperatures to be decreased for radiant cooling systems in humid climates. We conduct an FTIR analysis on three candidate membrane materials and fabricate a prototype experimental test panel that allows for thermal performance evaluation at different panel orientation and depths. Our study shows that for a 5 °C chilled panel temperature, the exterior membrane surface temperature reaches 26 °C in a 32 °C / 70% RH environment resulting in an effective panel temperature of 15.8 °C. Such a panel construction would avoid condensation in many humid environments and allow for radiant cooling without any latent load handling

Neutrophil apoptosis during experimentally induced Staphylococcus aureus mastitis

Abstract -The objective of this study was to determine whether neutrophil apoptosis and their consequent elimination by macrophages from the mammary gland is modulated by an infection caused by Staphylococcus aureus (S. aureus). The study was performed on twenty mammary glands of 5 virgin heifers. A buffered physiological solution (PBS) was administered as a means of control into the mammary glands of the heifers and after 168 h, the glands were inoculated with S. aureus. The samples of cell populations were obtained by lavages of the mammary glands in 4 intervals (24, 48, 72 and 168 h) after the experimental infection. Flow cytometry was used for determination of Annexin-V positivity and propidium iodide (PI) negativity of neutrophils. Light microscopy was used for determination of neutrophil karyopyknosis. Cytochemistry was used for the detection of myeloperoxidase-positive (MPO+) macrophages. Instillation of S. aureus resulted in an intramammary infection which persisted during the following experimental period. The total number of both Annexin-V-positive and PI negative neutrophils and karyopyknotic neutrophils peaked at 24 h after both of PBS and S. aureus administration. The highest percentages of Annexin-V-positive and PI negative neutrophils and karyopyknotic neutrophils were detected 48 and 168 h after PBS and S. aureus administration, respectively. The total number of MPO+ macrophages was the highest 24 h and 48 h after PBS and S. aureus administration, respectively; the percentage of MPO+ macrophages was the highest at 72 h in both cases. The dynamics of resolution of mastitis caused by S. aureus was very similar to the resolution of inflammatory response of the mammary gland after PBS administration. Mechanisms of cell pathogen elimination as well as inflammation resolution were very intensively involved; nevertheless, the mammary gland infection persisted. An early inclusion of the mechanisms of an acute inflammatory resolution thus paradoxically led to chronic infection

Thermal conditions in indoor environments: Exploring the reasoning behind standard-based recommendations

Professionals in the building design and operation fields typically look at standards and guidelines as a reliable source of information and guidance with regard to procedural, contractual, and legal scope and requirements that are relevant to accountability issues and compliance necessities. Specifically, indoor environmental quality (IEQ) standards support professionals to bring about comfortable thermal, air quality, acoustic, or visual conditions in buildings. In this context, it appears essential to regularly examine the IEQ standards’ applicability and scientific validity. The present contribution focuses on common thermal comfort standards in view of the reasoning and includes evidence behind their recommendations and requirements. Thereby, several international and national thermal comfort standards are examined via a structured matrix to assess basic parameters, design and performance variables targeted by the standards, suggested value ranges, and both general and specific evidence from the scientific literature. Finally, this paper discusses findings and points to the identified gaps in the chain of evidence from the results of scientific studies and the recommendations included in the thermal standards. As such, the present contribution has the potential to inform future developments regarding transparent and evidence-based thermal standards