Aerogenerador de eje vertical

Este proyecto nace de la necesidad de tener energía eléctrica en cada hogar, debido al aumento de nuevos aparatos eléctricos, del aumento del coste de la energía por parte de las compañías eléctricas y de la inminente desaparición de los materiales fósiles como el petróleo o el carbón para la generación de electricidad. Para ello se crea este proyecto, para que comunidades de vecinos o viviendas aisladas, tengan la posibilidad de autoabastecerse de energía eléctrica. A pesar de un primer desembolso de dinero para su implantación, tras su implantación se verá reducida la factura de la luz. Este proyecto se compone de dos grandes subgrupos, la parte mecánica y la parte eléctrica o electrónica. De estas dos, nos hemos centrado en la parte mecánica. Que se descompone en varios subconjuntos que son; la base del aerogenerador, la jaula completa y el posicionamiento o la parte superior del aerogenerador. Cada subconjunto se divide en mas subconjunto y finalmente en cada componente. Para ello se ha realizado un pequeño estudio aerodinámico de las zonas ideales de colocación del aerogenerador, altura mínima de colocación para una optima generación. Por otra parte, para la elección del numero de alabes del rotor se ha tomado en cuenta un estudio realizado en un túnel de viento realizado por Ben F. Blackwell, Robert E. Sheldahl y Louis V. Feliz. En la que se llega a la conclusión que mas alabes no aumenta la eficiencia del aerogenerador. Por lo que se optó por un aerogenerador de dos alabes. Puesto que la eficiencia era pequeña debido a que cuando el aire golpea en un rotor desnudo, disminuye la velocidad de giro de éste por que el aire golpea en sus partes cóncavas y convexas generando fuerzas en sentidos opuestos. Por lo que se desarrollo un estator para la canalización del flujo del aire a los alabes del rotor. Este estator es de aberturas regulables según el caudal de aire que se disponga, también funciona como mecanismo de seguridad en caso de velocidades muy grandes de viento, para evitar que el rotor se embale y genere daños dentro de este. Este mecanismo de posicionamiento de los alabes del estator se regulan mediante un PLC que tiene varios sensores por el aerogenerador para abrir o cerrar el estator cuando haga falta. Debido a que el estator es semiautomático, se han previsto una serie de medidas de prevención de riesgos para evitar daños físicos. También es necesario que se coloque una barandilla que limite el espacio del aerogenerador o por el contrario delimitar el acceso de las azoteas a personal autorizado. El posicionamiento de los alabes del estator se controlan desde la parte superior del aerogenerador, mediante un motor step, un reductor y un disco del cual salen vástagos con garfios en el extremo que se unen al alabe móvil. La fijación entre vástago y garfio se realiza mediante un pasador. El motor step es quien proporciona un torque pequeño que al pasar por el reductor aumenta hasta darnos el par necesario para mover el conjunto de los alabes del estator con rachas de viento hasta . El motor step va fijado mediante una brida metálica al soporte de reductor para evitar que se mueva. El reductor se fija a la pieza mediante la cual pivota el disco de posicionamiento. La pieza de pivote se le han realizado una serie de rebajes disminuir el peso, por lo que para su conformado se realizará mediante inyección de plástico al igual que el garfio y el disco de posicionamiento. El aerogenerador esta sujeto mediante seis pilares inferiores y un pilar central que se encarga de sustentar el rotor. Estos pilares reparten el peso del aerogenerador y a su vez sostienen la pletina exterior que esconde los elementos que hay debajo como; la multiplicadora, el alternador, el cardan y el PLC. La pletina tendrá una abertura por la que el operario tendrá acceso a sus partes. La pletina exterior estará formada por varias láminas de acero unidas por cordones de soldadura. La pletina estará sujeta mediante tornillería a los pilares. El montaje de los subconjuntos se realizarán en el sitio donde se vaya a colocar el aerogenerador a excepción del reductor que es posible su montaje en taller. Previamente se tendrán que colocar barras roscadas en el suelo de la azotea para la posterior colocación y amarre de los pilares. En ese instante se colocará la multiplicadora y el alternador. La jaula junto con los alabes se montará encima de los pilares y a su vez se colocará el rotor. Posteriormente se colocará la tapa y el mecanismo de posicionamiento de los alabes y la cúpula. Una vez fijado el rotor se colocará el cardan que unirá rotor y multiplicadora. Se colocará el acople entre alternador y la multiplicadora. Se finalizara con el cierre de la pletina. Se colocarán los aparatos electrónicos que harán que el aerogenerador se comporte como un aparato semiautomático. En un compartimento dentro del edificio se colocarán baterías que acumularán la energía generada. En este habitáculo se colocará un aparato donde se visualice la potencia que se esta generando así como la velocidad de rotación y la velocidad del viento. Junto a este aparato un pulsador de parada de emergencia. Alrededor del aerogenerador se colocarán señales que indiquen los peligros que se pueden dar así como, las precauciones a tener en cuenta. Las medidas vendrán escritas en un documento junto con los mantenimientos que se han de dar. En la puerta de acceso a la azotea y en la ventana de acceso a los interiores del aerogenerador habrá un resumen del documento anteriormente descrito.

Hamburg model basin : numerical and experimental investigations on the dependence of the fluctuations of thrust and torque on the blade area ratio of four-blade ship-propellers /

Contract no. N 62558-2223.

Composite load model decomposition : induction motor contribution

In this thesis, a new technique has been developed for determining the composition of a collection of loads including induction motors. The application would be to provide a representation of the dynamic electrical load of Brisbane so that the ability of the power system to survive a given fault can be predicted. Most of the work on load modelling to date has been on post disturbance analysis, not on continuous on-line models for loads. The post disturbance methods are unsuitable for load modelling where the aim is to determine the control action or a safety margin for a specific disturbance. This thesis is based on on-line load models. Dr. Tania Parveen considers 10 induction motors with different power ratings, inertia and torque damping constants to validate the approach, and their composite models are developed with different percentage contributions for each motor. This thesis also shows how measurements of a composite load respond to normal power system variations and this information can be used to continuously decompose the load continuously and to characterize regarding the load into different sizes and amounts of motor loads.

Automated symbolic and numeric procedure for manipulator rigid-body dynamic significance analysis and simplification

This paper describes an automated procedure for analysing the significance of each of the many terms in the equations of motion for a serial-link robot manipulator. Significance analysis provides insight into the rigid-body dynamic effects that are significant locally or globally in the manipulator's state space. Deleting those terms that do not contribute significantly to the total joint torque can greatly reduce the computational burden for online control, and a Monte-Carlo style simulation is used to investigate the errors thus introduced. The procedures described are a hybrid of symbolic and numeric techniques, and can be readily implemented using standard computer algebra packages.

Control issues for velocity sourced Series Elastic Actuators

The Series Elasic Actuator has been proposed as a method for providing safe force or torque based acutation for robots that interact with humans. In this paper we look at some outstanding issues in the implementation and control of Series Elastic Actuators. The study addresses issues in making the Series Elastic Actuator respond effectively in the presence of physical difficulties such as restriction, using a computation efficient controller. The improvement over previous implementations is achieved by treating the motor as a velocity source to the elastic element, rather than as a torque source.

Assessing the safety of a velocity sourced series elastic actuator using the head injury criterion

The Velocity Sourced Series Elastic Actuator has been proposed as a method for providing safe force or torque based actuation for robots without compromising the actuator performance. In this paper we assess the safety of Velocity Sourced Series Elastic Actuators by measuring the Head Injury Criterion scores for collisions with a model head. The study makes a comparative analysis against stiff, high impedance actuation using the same motor without the series elastic component, showing that the series elastic component brings about a massive reduction in the chance of head injury. The benefits of a collision detection and safe reaction system are shown to be limited to collisions at low speeds, providing greater interaction comfort but not necessarily contributing to safety from injury.

Modelling the foil rolling process

A review of the main rolling models is conducted to assess their suitability for modelling the foil rolling process. Two such models are Fleck and Johnson's Hertzian model and Fleck, Johnson, Mear and Zhang's Influence Function model. Both of these models are approximated through the use of perturbation methods. Decrease in the computation time resulted when compared with the numerical solution. The Hertzian model was approximated using the ratio of the yield stress of the strip to the plane-strain Young's Modulus of the rolls as the small perturbation parameter. The Influence Function model approximation takes advantage of the solution of the well-known Aerofoil Integral Equation to gain an insight into how the choice of interior boundary points affects the stability of numerical solution of the model's equations. These approximations require less computation than their full models and, in the case of the Hertzian approximation, only introduces a small error in the predictions of roll force roll torque. Hence the Hertzian approximate method is suitable for on-line control. The predictions from the Influence Function approximation underestimates the predictions from the numerical results. Better approximation of the pressure in the plastic reduction regions is the main source of this error.

Bone response to unloaded titanium implants in the fibula, iliac crest, and scapula: an animal study in the Yorkshire pig

The reconstruction of extended maxillary and mandibular defects with prefabricated free flaps is a two stage procedure, that allows immediate function with implant supported dentures. The appropriate delay between prefabrication and reconstruction depends on the interfacial strength of the bone–implant surface. The purpose of this animal study was to evaluate the removal torque of unloaded titanium implants in the fibula, the scapula and the iliac crest. Ninety implants with a sandblasted and acid-etched (SLA) surface were tested after healing periods of 3, 6, and 12 weeks, respectively. Removal torque values (RTV) were collected using a computerized counterclockwise torque driver. The bicortical anchored 8 mm implants in the fibula revealed values of 63.73 Ncm, 91.50 Ncm, and 101.83 Ncm at 3, 6, and 12 weeks, respectively. The monocortical anchorage in the iliac crest showed values of 71.40 Ncm, 63.14 Ncm, and 61.59 Ncm with 12 mm implants at the corresponding times. The monocortical anchorage in the scapula demonstrated mean RTV of 62.28 Ncm, 97.63 Ncm, and 99.7 Ncm with 12 mm implants at 3, 6, and 12 weeks, respectively. The study showed an increase of removal torque with increasing healing time. The interfacial strength for bicortical anchored 8 mm implants in the fibula was comparable to monocortical anchored 12 mm implants in the iliac crest and the scapula at the corresponding times. The resistance to shear seemed to be determined by the type of anchorage (monocortical vs. bicortical) and the length of the implant with greater amount of bone–implant interface.

Comparison between maximal lengthening and shortening contractions for biceps brachii muscle oxygenation and hemodynamics

Eccentric contractions (ECC) require lower systemic oxygen (O2) and induce greater symptoms of muscle damage than concentric contractions (CON); however, it is not known if local muscle oxygenation is lower in ECC than CON during and following exercise. This study compared between ECC and CON for changes in biceps brachii muscle oxygenation [tissue oxygenation index (TOI)] and hemodynamics [total hemoglobin volume (tHb) = oxygenated-Hb + deoxygenated-Hb], determined by near-infrared spectroscopy over 10 sets of 6 maximal contractions of the elbow flexors of 10 healthy subjects. This study also compared between ECC and CON for changes in TOI and tHb during a 10-s sustained and 30-repeated maximal isometric contraction (MVC) task measured immediately before and after and 1–3 days following exercise. The torque integral during ECC was greater (P < 0.05) than that during CON by ∼30%, and the decrease in TOI was smaller (P < 0.05) by ∼50% during ECC than CON. Increases in tHb during the relaxation phases were smaller (P < 0.05) by ∼100% for ECC than CON; however, the decreases in tHb during the contraction phases were not significantly different between sessions. These results suggest that ECC utilizes a lower muscle O2 relative to O2 supply compared with CON. Following exercise, greater (P < 0.05) decreases in MVC strength and increases in plasma creatine kinase activity and muscle soreness were evident 1–3 days after ECC than CON. Torque integral, TOI, and tHb during the sustained and repeated MVC tasks decreased (P < 0.01) only after ECC, suggesting that muscle O2 demand relative to O2 supply during the isometric tasks was decreased after ECC. This could mainly be due to a lower maximal muscle mass activated as a consequence of muscle damage; however, an increase in O2 supply due to microcirculation dysfunction and/or inflammatory vasodilatory responses after ECC is recognized.

Biceps brachii muscle oxygenation in electrical muscle stimulation

The purpose of this study was to compare between electrical muscle stimulation (EMS) and maximal voluntary (VOL) isometric contractions of the elbow flexors for changes in biceps brachii muscle oxygenation (tissue oxygenation index, TOI) and haemodynamics (total haemoglobin volume, tHb = oxygenated-Hb + deoxygenated-Hb) determined by near-infrared spectroscopy (NIRS). The biceps brachii muscle of 10 healthy men (23–39 years) was electrically stimulated at high frequency (75 Hz) via surface electrodes to evoke 50 intermittent (4-s contraction, 15-s relaxation) isometric contractions at maximum tolerated current level (EMS session). The contralateral arm performed 50 intermittent (4-s contraction, 15-s relaxation) maximal voluntary isometric contractions (VOL session) in a counterbalanced order separated by 2–3 weeks. Results indicated that although the torque produced during EMS was approximately 50% of VOL (P<0Æ05), there was no significant difference in the changes in TOI amplitude or TOI slope between EMS and VOL over the 50 contractions. However, the TOI amplitude divided by peak torque was approximately 50% lower for EMS than VOL (P<0Æ05), which indicates EMS was less efficient than VOL. This seems likely because of the difference in the muscles involved in the force production between conditions. Mean decrease in tHb amplitude during the contraction phases was significantly (P<0Æ05) greater for EMS than VOL from the 10th contraction onwards, suggesting that the muscle blood volume was lower in EMS than VOL. It is concluded that local oxygen demand of the biceps brachii sampled by NIRS is similar between VOL and EMS.

Flexural and torsional rigidity of colonoscope at room and body temperatures

A series of experiments have been conducted to determine the flexural, EI, and torsional, GJ, rigidity of an Olympus colonoscope CF‐140S and torsional rigidity of a Pentax colonoscope EC‐3870 and the dependency of these properties on temperature and on the presence of loops. Along the length of the colonoscope, the Olympus colonoscope flexural rigidity varied between 260 and 400 Ncm2 and torsional rigidity varied between 68 and 88 Ncm2/deg, with an average of 76 Ncm2/deg for tests involving 0.86 Nm of anti‐clockwise torque. Results show a significant decrease of 10% in torsional rigidity between clockwise and anti‐clockwise torque. For the Pentax colonoscope flexural rigidity was not tested; its torsional rigidity varied between 34 and 76 Ncm2/deg, with an average of 46 Ncm2/deg for tests involving 0.43 Nm of anti‐clockwise torque. An increase in temperature of the Olympus colonoscope from 24°C to 37°C reduces EI by an average of 17% and GJ by an average of 7%. A right‐handed loop caused a significant increase in flexural rigidity, but other looping configurations had no significant influence.

Dynamics of railway wagons subjected to braking torques on defective tracks

It is well known that track defects cause profound effects to the dynamics of railway wagons; normally such problems are examined for cases of wagons running at a constant speed. Brake/traction torques affect the speed profile due to the wheel–rail contact characteristics but most of the wagon–track interaction models do not explicitly consider them in simulation. The authors have recently published a model for the dynamics of wagons subject to braking traction torques on a perfect track by explicitly considering the pitch degree of freedom for wheelsets. The model is extended for cases of lateral and vertical track geometry defects and worn railhead and wheel profiles. This paper presents the results of the analyses carried out using the model extended to the dynamics of wagons containing less ideal wheel profiles running on tracks with geometry defects and worn rails.